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Cosorption processes of triethylene glycol in a packed-bed liquid desiccant dehumidifier.


The environmental awareness that occurred in the 1980s has led to the public demand for healthy outdoor and indoor environments. Today, the common approach adopted for cleaning contaminated indoor air is diluting indoor air with clean outdoor air in order to achieve acceptable indoor air quality. Air cleaning may be an attractive alternative when the outdoor air quality is poor or when there is a desire to reduce energy costs associated with high outdoor air exchange rates. Accordingly, substantial research efforts has been carried out to identify effective methods to clean indoor air while minimizing energy consumption.

Although water vapor is not considered to be an air contaminant, it has a significant impact on both comfort and health of human occupants. Unlike other contaminants in the air, water vapor cannot be completely removed from the airstream. Extremely low relative humidity can lead to eye irritation, mucous dryness, and other health problems (Arundel et al. 1992). On the other hand, high indoor relative humidity can lead to condensation on cold faces, such as ducts, dryers, windows and various building materials, and can also promote the growth of various micro-organisms. Therefore, the control of the indoor relative humidity is an important design parameter for HVAC industry practitioners and researchers.

In conventional vapor-compression systems, the process of dehumidification and humidity control is brought about by cooling the air below its dew point and reheating the air to the desired relative humidity. This method of humidity control is extremely energy intensive. In contrast, systems employing desiccants can help control the humidity by removing water from the air through a sorption process.

Desiccants are broadly classified as solid and liquid. Solid desiccants include molecular sieves and silica gel, whereas liquid desiccants include both inorganic salt such as lithium bromide and an organic solution like triethylene glycol (TEG). Generally, systems that employ solid desiccants can achieve a higher degree of dehumidification than systems that use liquid desiccants.

Although liquid desiccant systems have lower drying capacities than solid desiccant systems, systems that use liquid desiccants do have some advantages. These include the possibility of cooling the air by manipulating the air and liquid flow rates, the lower regeneration temperatures required by liquid desiccants, and the fact that liquid systems generally do not require complex dehumidifier geometries and thus can be easily applied to various commercial and industrial applications.

In liquid desiccant systems, air and desiccant can be brought into contact in several absorber configurations. This can be achieved by either (a) bubbling the air through the liquid, as in a tray column; (b) spraying the liquid in a fine dispersion in an upward airstream, as in a spray tower; (c) spraying the liquid over a bank of cooling tubes past which air is blown; or (d) passing the air and liquid streams through a packed bed. In this study, we focus on a packed bed, since it usually offers lower pressure drop but has the greatest interfacial area between the air and the liquid streams. This leads to high interchange rates of heat and mass between the liquid desiccant and the airstream.

Apart from studies on the heat and mass transfer characteristics of desiccants, desiccant potential for removing indoor air contaminants has also begun to be recognized. Hines and Ghosh (1993) revealed the capabilities of solid desiccants such as silica gel and a molecular sieve for removing contaminants from the air. Activated carbon fiber has been shown to have a strong capability of removing volatile organic compounds (VOCs) (Das et al. 2004). Hydrpohobic zeolites, in membrane form, have also been found have an ability to selectively remove one or more organic pollutants from humid airstreams (Chitawar and Greene 1997; Aguado et al. 2004). Consequently, solid absorbents have been employed in many industrial applications in the US and Asia for independently controlling humidity and VOCs.

On the other hand, Moschandreas and Relwani (1990) demonstrated that a liquid desiccant-based, gas-fired dehumidification system using lithium chloride solution (LiCl) had the capacity to remove indoor pollutants. However, no systematic study has been conducted to assess the full adsorption potential of LiCl solution. It was not until the work done by Hines et al. (1992), a comprehensive study conducted to determine the removal capabilities of both organic and inorganic liquid absorbents, that organic compounds such as TEG solution were found to have a stronger capability to remove the organic contaminants from the airstreams than inorganic salts such as lithium chloride. The researchers attributed this phenomenon to the fact that the removal capacity of the LiCl solutions depend to a great extent on the solubility of the particular pollutant in the water portion of the solution while the pollutant removal capacities of the TEG solution depends on the solubility of the pollutants in the absorbent portion of the solution.

Although research on the absorption properties of TEG for water and various VOCs have been ongoing for some time (Peng and Howell 1981; Ng et al. 1983; Grasso et al. 1994), there is no published experimental or numerical data on the simultaneous absorption of water vapor and air contaminants by TEG. Chau and Worek (2007) developed a numerical model to simulate the cosorptive characteristics of TEG solutions for both water vapor and VOC air contaminants under various operating conditions through packed-bed absorbers. Toluene, which has widely been recommended as a reasonable surrogate and representative for indoor total VOCs in desiccant adsorption studies (Liu 1990, 1993), has been selected as another component in the air-stream. In this paper, the influence of various operating conditions on the cosorptive performance of packed dehumidifiers is examined. Also, the optimal operating conditions in which TEG solutions can absorb the maximum amount of toluene in the presence of water vapor are presented.


In order to model the cosorption processes that occur in the dehumidifier, the dehumidifier is divided into a sequence of stages. Figure 1 shows

a schematic representation of a packed dehumidifier with a typical stage enlarged. The following assumptions were made when deriving the model:


1. no accumulation of mass and energy fluxes at the interface of each stage;

2. plug flow of each phase (i.e., absence of radial gradients of velocity, temperature, and composition for the bulk flow);

3. the pressure drop across the packed column is small compared to the absolute pressure, which is assumed to be close to atmospheric pressure;

4. uniform and constant concentrations at the inlet streams;

5. ideal gas behavior for the vapor phase;

6. constant specific heats for both the liquid and vapor phases;

7. equal heat and mass transfer areas;

8. finite and constant mass and heat transfer coefficients;

9. negligible axial conduction or diffusion of thermal energy; and

10. negligible droplet carry-over effect despite this may commonly exist in packed columns but can be mitigated by incorporating a droplet filter or demister in the system.

Conservation of Mass

Applying the conservation law of mass for pollutant i in the vapor phase at stage j, we have the following:

[Vapor flow from stage j + 1 = Vapor flow from stage j + Molar mass flux from vapor to liquid phase]

[v.sub.ij] - [v.sub.i,j+1] + [N.sub.ij.sup.V][a.sub.j] = 0 (1)

For the liquid phase we have the following:

[Liquid flow from stage j = Liquid flow from stage j - 1 + Molar mass flux from vapor to liquid phase]

[l.sub.ij] - [l.sub.i,j-1] + [N.sub.ij.sup.L][a.sub.j] = 0 (2)

Since we are concerned with only toluene vapor and water vapor transport in the inert air and TEG solution, we have two equations for vapor mass conservation and two equations for liquid mass conservation. With assumption 1, [N.sub.ij.sup.V]given in Equation 1 must be equal to that given in Equation 2.

Conservation of Energy

Similarly, applying the conservation of energy for the vapor phase, we have the following:

[Vapor enthalpy for stage j + Heat input + Energy transfer from vapor to liquid phase = Vapor enthalpy for stage j + 1]

[V.sub.j][H.sub.j.sup.V] - [V.sub.j + 1] [H.sub.j+1.sup.V] + [Q.sub.j.sup.V] + [e.sub.j.sup.V][a.sub.j] = 0, (3)

For the liquid phase we have the following:

[Liquid enthalpy for stage j + Heat input + Liquid enthalpy for stage j + 1 = Energy transfer from vapor to liquid phase]

[L.sub.j][H.sub.j.sup.L] - [L.sub.j - 1] [H.sub.j-1.sup.L] + [Q.sub.j.sup.L] + [e.sub.j.sup.L][a.sub.j] = 0, (3)

Mass and Energy Transfer Rates

The vapor mass transfer rate equation for the multi-component mass transfer process can be described by the following:

[Overall mass transfer = Mass transfer due to diffusion + Mass transfer due to convection]

([N.sub.V]) = [[k.sup.V]]a([y.sub.i.sup.V] - [y.sub.i.sup.I]) + [N.sub.t.sup.V] ([y.sub.V]) (5)

On the other hand, the liquid mass transfer rate equation can be represented by

([N.sub.L]) = [[k.sup.L]]a([y.sub.i.sup.I]-[y.sub.i.sup.L]) + [N.sub.t.sup.T] ([y.sup.L]). (6)

Once again, as we are concerned only with toluene and water vapor, this gives two vapor mass transfer rate equations and two liquid mass transfer rate equations.

Similarly, the energy flux rate equation, which is also made up of a diffusive and a convective component, can be represented by the following:

[Overall energy transfer = Energy transfer due to conduction + Energy transfer due to convection]

[e.sub.j.sup.V] = [q.sub.j.sup.V] + [3.summation over (i = 1)][[~.H].sub.ij.sup.V][N.sub.ij] (7)

[e.sub.j.sup.L] = [q.sub.j.sup.L] + [3.summation over (i = 1)][[~.H].sub.ij.sup.L][N.sub.ij] (8)

Using assumption 1, that gives [e.sub.j.sup.V] = [e.sub.j.sup.L], Equations 7 and 8 can be reduced to

[q.sub.j.sup.L] + [3.summation over (i = 1)][[~.H].sub.ij.sup.L][N.sub.ij] = [q.sub.j.sup.V] + [3.summation over (i = 1)][[~.H].sub.ij.sup.V][N.sub.ij]. (9)

Interface Equilibrium Relations

The interface equilibrium equations, which relate the mole fractions of vapor to liquid phases, for both water vapor and toluene vapor in a TEG solution, are given by

[K.sub.ij][x.sub.ij.sup.I] - [y.sub.ij.sup.I] = 0 for i = 1, 2 and j = stage number. (10)

In Equations 1 through 6 and Equation 9, there are fifteen total unknowns with thirteen equations. In order to solve the equations, we can make use of the molar ratio definition that the summation of the molar ratios of all the species in liquid or vapor phases is equal to unity. We have the following summation relationships in the interface for vapor and liquid phases, respectively:

[3.summation over (i = 1)][y.sub.ij.sup.I] - 1 = 0 (11)

[3.summation over (i = 1)][x.sub.ij.sup.I] - 1 = 0 (12)

By using Equations 11 and 12, the number of unknowns is reduced to thirteen for each stage j(i.e., [v.sub.ij], [l.sub.ij], [y.sub.ij.sup.I], [x.sub.ij.sup.I]), [N.sup.ij], [T.sub.j.sup.V], [T.sub.j.sup.L], [T.sub.j.sup.I]

Mass Transfer Rate Equation

The matrix of multi-component diffusive mass transfer coefficients, [[k.sup.V]], in Equation 5 can be calculated from

[k.sup.V]a = [[B.sup.V].sup.-1]a[[PHI].sup.V][{exp[[[PHI].sup.V]] - [I]}.sup.-1], (13)

and the elements of the matrices [B] and [[PHI]] are defined with the elements


and [M.sub.ik] = [-m.sub.i]([1/[k.sub.ik.sup.V]a] - [1/[k.sub.i2.sup.V]a]) i [not equal to] k = 1, 2, (15)

where the [k.sub.ij.sup.V] are the mass transfer coefficients of the binary pairs i-k in vapor phase and are most accurately estimated by the Onda et al. correlation (Chau and Worek 2007). The other property correlations used for estimating the binary mass transfer coefficients, [k.sub.ij.sup.V], are given in Table 1.


Heat Transfer Rate Equations

The conduction heat fluxes of the vapor and liquid phases are related to the heat transfer coefficients by the following two equations:

[q.sub.j.sup.V] = [h.sub.j.sup.V]([T.sub.j.sup.V] - [T.sub.j.sup.I]) (16)

[q.sub.j.sup.L] = [h.sub.j.sup.L]([T.sub.j.sup.L] - [T.sub.j.sup.I]) (17)

Liquid and Vapor Enthalpies

The total vapor enthalpy for stage j, [H.sub.j.sup.V], in Equation 4 can be determined from the following equation:


Assuming all the species in the vapor phase behave as an ideal gas, the molar heat enthalpy of the individual components is equal to


However, the latent heat of vaporization term for air at each stage j, [MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII], has not been included in Equation 19 since air exists in vapor form at normal room temperatures. Although toluene exists in liquid form at room temperature, we are concerned with toluene existing as a contaminant vapor in an airstream in this study. Thus, the [MATHEMATICAL EXPRESSION NOT REPRODUCIBLE IN ASCII] term for toluene should also not be included when we construct Equation 19 for toluene.

On the other hand, the total liquid enthalpy can be determined by


where [[~.H].sub.i,j.sup.L] is the molar heat enthalpy of component i at stage j and is equal to

[[~.H].sub.i,j.sup.L] = [C.sub.P,ij.sup.L]([T.sub.j.sup.L] - [T.sub.j.sup.o]). (21)

The integral heat of solution in Equation 21, [DELTA][H.sub.s], can be determined by the following curve-fitted equations using the data given by Konnecke et al. (1958) and Peng and Howell (1981):



[c.sub.1] = 2.0757612 - [49.40696/T], [c.sub.2] = [(-5.2417769 x [10.sup.-5] + 7.2199564 x [10.sup.-6]).sup.-1], [c.sub.3] = [(5.2137659 x [10.sup.-5] - 7.2086831 x [10.sup.-6]).sup.-1], [c.sub.4] = -2.32200443 + [79.815955/T], [c.sub.5] = 1.3500517 - [73.93278/T]

The partial molar enthalpies of water and TEG in Equation 21 can be evaluated by the integral heat of solution data in the water-TEG system through the use of the following set of equations:




Equilibrium Relationships

The equilibrium constant [K.sub.i,j] in Equation 10, which relates the molar ratio of vapor to liquid, can be expressed as

[k.sub.i,j] = [[[gamma].sub.i][P.sub.i.sup.s]/P], (26)

where P and [P.sup.s] are the total system pressure and the saturated water or toluene vapor pressure, respectively. On the other hand, the activity coefficients of water vapor and toluene vapor in TEG solution can be determined from the reported semi-empirical relationships by Parish et al. (1990) and Gupta et al. (1989), respectively, which are listed in Table 1.


In this study, we divided the entire packed dehumidifier into ten stages and the resulting 130 equations were solved simultaneously by the Newton Method. The equations were grouped together by stages beginning at the top stage and going sequentially stage by stage down the column as proposed by Napthali and Sandholm (1971). When grouping in this way, the partial derivatives for a particular stage of the Jacobian matrix depend only on the variables for its own stage and the two adjacent stages. The resulting Jacobian has a stage-wise sparse block tri-diagonal structure that can easily be solved using the Thomas Algorithm subroutines for general matrix forms as written by Hindmarsh (1977).

Experimental data relating to the binary mixture are needed for validating the numerical model. However, no cosorption data is available for the TEG solution. In order to validate the model, a two-phase validation approach was adopted. Initially, single-component water vapor experimental data presented by Chung et al. (1993) and Chung (1994) on a packed bed of TEG solution absorbing water vapor were compared to the results predicted by numerical model. The predicted results agreed within 5% with the experimental results.

In the second validation phase, an experimental setup was constructed to examine the water and toluene cosorption characteristics of TEG solutions and to validate the numerical data derived by the developed numerical model. The absorption study was carried out in a packed dehumidifier that contained 6.0 x 6.0 x 1.0 mm glass raschig rings, which were dumped randomly in the dehumidifier to a height of 17.8 cm. The middle section of the dehumidifier, containing the packing, had a glass perforated plate that not only supports the packing but also acts as an air distributor. The top cap had two 6.35 cm (1/4 in.) tubes for the air to exit and to introduce the liquid. The air was introduced in the solution through a 6.35 cm diameter tube connected at the bottom side. The specification of the packed dehumidifier and the relevant operating parameters are summarized in Table 2.
Table 2. Configuration of Packed Dehumidifier Used in Experiments

Descriptions                          Range of Values

Diameter of column                    5.08 cm (2 in.)

Type of packing       6 x 6 x 1.0 mm glass raschig rings with specific
                          surface area of 885 [m.sup.2]/[m.sup.3]

Packing height                               cm

Concentration of TEG                    95% and 100%

Inlet air and liquid               20[degrees]C-25[degrees]C
and temperatures

Inlet airflow rate                    2300-2800 mL/min

Inlet liquid flow                        256 mL/min

Inlet air relative                        30%-70%

Toluene                                 1000-2000 ppm

A schematic diagram of this experimental setup is shown in Figure 2; more details are given in Chau and Worek (2007). Again, the numerical results agreed very well with the experimental results.



The performance of the packed dehumidifier was evaluated in terms of the moisture and toluene removal rate and the dehumidifier efficiency. The moisture and toluene removal rate for a packed dehumidifier were calculated by Equation 27:

[m.sub.i] = [m.sub.a]([y.sub.i,inlet] - [y.sub.i,outlet]) (27)

The dehumidifier efficiency is defined as the ratio of the actual change in moisture content or toluene concentration of the air leaving the dehumidifier to the maximum possible change in moisture content or toluene concentration under a given set of operating conditions. The dehumidifier effectiveness can be expressed as

[[epsilon].sub.i] = ([[y.sub.i,in] - [y.sub.i,out]]/[[y.sub.i,in] - [y.sub.i,eqm]]), (28)

where [y.sub.i,egm] is the humidity ratio or toluene concentration of the air, which is at equilibrium with the TEG solution at the inlet concentration and temperature.

The removal rate and the dehumidifier efficiency are influenced by climatic and operational variables such as the temperature and relative humidity of the ambient air, the inlet temperature and the concentration of the TEG solution, as well as airflow and solution flow rates. In this research, the impacts of these operational variables on the removal rate and removal efficiency of moisture and toluene vapor for the dehumidifier were investigated.

In order to study their impacts, a packed dehumidifier was assumed to be operated in a countercurrent mode with 90%, 95%, and 100% TEG liquid desiccant flowing from one side and air with various relative humidities flowing from the other side. Sixteen millimeter diameter polypropylene rings were used as packing for the packed dehumidifier. All the other configurations, which are given in Table 3, were based on those employed in the experiments conducted by Chung et al. (1993) and Chung (1994).
Table 3. Configuration of the Packed Dehumidifier used in the Numerical

Item                              Descriptions

Diameter of column                  0.1524 m

Type of packing       1.588 cm (5/8 in.) dia. polypropylene
                         flexi rings with 342 [m.sup.2]/

Packing height                        42 cm

Concentration of TEG            90%, 95%, and 100%

Inlet air and liquid     25[degrees]C, 33[degrees]C, and
and temperatures                  40[degrees]C

Inlet airflow rate                  18.9 L/s

Inlet liquid flow              0.0947 to 0.177 L/s

Inlet air humidity    0.006-0.020 kg [H.sub.2]O/kg dry air

Toluene                              100 ppm

A series of computer runs were performed to determine the optimum dehumidifier height. The results show that increasing the packing height generally improves the water and toluene removal rates, as this provides more areas for heat and mass transfer. However, as shown in Figures 3a and 3b, an optimal height was found to exist such that further increase of the packing height would not lead to a significant increase in the removal rates or removal efficiencies for moisture and toluene vapor.

It can be observed that the optimal height for the removal of water vapor was different from that of toluene vapor. The moisture removal rate increased until it reached an optimal height of 0.6 m, and further increase of the packing height did not lead to a significant increase in the moisture removal rate. If the packing height was increased beyond 0.6 m, more toluene vapor would be removed per unit time by the TEG solution. The increase in toluene removal rate was observed with the increase in the packing height until the packing height reached approximately 0.9 m, where the TEG solution was saturated with both water and toluene vapor. Increasing the packing height beyond 0.9 m did not lead to any observable reduction in moisture removal rate, but it still resulted in an observable reduction in toluene removal rate. On the whole, less than 1% reduction in the moisture and toluene removal efficiencies was observed. However, these optimum values for packing height observed from Figures 3a and 3b are valid only for the liquid-to-gas ratio of 6.25. In reality, the optimum packing height normally comes from a trade-off between removal efficiency and air-side pressure drop.

The effect of relative flow rates of liquid desiccant to air (i.e., liquid-to-gas flow ratios) has always been identified to exert great influence on the dehumidification performance of dehumidifiers (Peng and Howell 1981; Perry and Green 1982; Lowenstein and Gabruk 1992). However, the effect of liquid-to-gas flow ratios on the removal rates and efficiencies of binary air components (i.e., water and toluene vapors) was not examined in these papers.

The profiles of the molar fluxes of water and toluene vapors at various air and liquid inlet temperatures are shown in Figures 4 through 6. These results would give some insight into the phenomenon occurring at different liquid-to-gas flow ratios. The molar fluxes of water and toluene vapors, which are denoted by in Equation 1, give an indication of the mass transfer driving potential, and the total area under the curve gives the total number of moles of gaseous components being removed by the TEG solution.

At low liquid-to-gas flow ratios, the molar flux of water vapor is greater at the top of the dehumidifier (i.e., at a height around 0.4 m). The effect was more prominent for an inlet air temperature of 33[degrees]C than at 25[degrees]C. Hence, the amount of water vapor being removed by the TEG solution and the degree of dehumidification is greater than at the top of the dehumidifier. The heat being released during the dehumidification process causes the air temperature profile to reach a maximum at the top of the dehumidifier, which causes an increase in the vapor pressure of the liquid stream. Therefore, the mass transfer driving force between the liquid and the air-stream and the molar flux of both components decrease from the top to the bottom of the dehumidifier.

As the liquid-to-gas flow ratio increases, the maximum air temperature decreases and the location of the maximum air temperature moves to the lower part of the dehumidifier. At a liquid-to-gas flow ratio of 6.25, the molar flux decreases rapidly from 1.9 x [10.sup.-5] to 1 x [10.sup.-6] kmol/[m.sup.2]*s. This suggests that greater amounts of water vapor and toluene are being removed by TEG and more dehumidification occurs in the lower portion of the dehumidifier. In the region above it, the air loses its heat to the cooler liquid, which will immediately exit the dehumidifier.

Figure 7a shows that both the moisture removal rate and the moisture removal efficiency increased with the liquid-to-gas flow ratio. Figure 7b shows a similar trend for both the toluene removal rate and the toluene removal efficiency. However, the removal efficiency of toluene vapor experienced a greater increase than that of water vapor. The moisture removal efficiency increased from 35% to 75% while the toluene vapor removal efficiency increased from 35% to 92% when the packed dehumidifier was running at a liquid-to-gas ratio of 6.25 instead of 0.625. This can be attributed to the fact that a higher liquid-to-gas flow ratio results in more liquid being available to remove both the water vapor and the toluene vapor from the airstream.

From the perspective of indoor air quality, it would be advantageous to operate the packed dehumidifier with the highest possible liquid-to-gas flow ratio, as this leads to a greater toluene removal rate. However, doing so is subject to the requirements of greater pumping power and higher operating costs.

Figure 8a shows the moisture removal rate and moisture removal efficiency of the TEG solution versus the change in the air inlet temperature when the packed dehumidifier was running at a liquid-to-gas flow ratio of 6.25. Figure 8b shows the toluene removal rate and toluene removal efficiency of the TEG solution versus the change in the air inlet temperature when the packed dehumidifier was running at a liquid-to-gas flow ratio of 6.25. The removal rates and the removal efficiencies for both water vapor and toluene vapor did not change as the inlet air temperature increased.

The influence of the liquid inlet temperature on the removal rate and efficiency of both water vapor and toluene were also investigated. Figure 9a shows that an increase in the liquid inlet temperature generally leads to a lower water vapor removal rate and efficiency, and the effect is prominent with the increase in the concentration of TEG solution. This is similar to the observation by Gandhidasan et al. (1987) in their studies focusing on the water vapor absorption characteristics of another liquid desiccant, lithium bromide. This phenomenon is due to the fact that fewer TEG molecules per unit volume are available for attracting both the water and the toluene molecules from the airstream.

In particular, results show that, for 90% TEG solution, the water vapor removal efficiency decreased from 50% to -380% as the liquid inlet temperature increased from 25[degrees]C to 40[degrees]C. The negative water removal efficiency indicates that the packed dehumidifier was actually releasing water vapor into the air instead of removing water vapor from the air.

In contrast, no release of water vapor was observed for the 95% TEG solution. This is due to the increase in the concentration of TEG in the TEG solution, which increases the vapor pressure between the airstream and the liquid solution. However, Figure 9a does show that the water removal efficiency decreased from 78% to 3% for the 95% TEG solution as the liquid inlet temperature increased from 25[degrees]C to 40[degrees]C. The water vapor removal rates for 100% pure TEG liquid remained constant at 98% for the range of liquid inlet temperatures being investigated.

As the liquid inlet temperature increases, the vapor pressure of the liquid increases correspondingly. The higher the vapor pressure of the liquid, the smaller the vapor pressure differences between the liquid and the airstream and thus the lower the mass transfer driving force from the airstream to the liquid. As a result, less water vapor is being removed, and in some instances water vapor is being released at a particular operating condition.

Similarly, Figure 9b shows the effect of liquid inlet temperature on the toluene removal rate and toluene removal efficiency. Generally, the liquid inlet temperature had weak sensitivity in the toluene removal rate and toluene removal efficiency for 90%, 95%, and 100% TEG solution when the dehumidifier was running at liquid-to-gas ratio of 6.25. In particular, the toluene removal efficiency for 90% TEG solution remained constant at about 89%, for 95% TEG solution at about 91%, and for 100% TEG solution at about 93% for the range of liquid inlet temperatures being investigated.


In this study, the effect of different design parameters on the performance of a packed-bed dehumidifier using TEG solution was investigated. The effects of liquid-to-air ratios, TEG solution concentrations, TEG inlet temperatures, and air inlet temperatures were reported on the moisture and toluene removal rates as well as the moisture and toluene removal efficiencies of the packed dehumidifier. Running the packed dehumidifier at a higher liquid-to-gas flow ratio leads to higher removal rates and efficiencies for both toluene vapor and water vapor. The inlet temperature of the TEG solution was found to exert great influences on the removal rates and removal efficiencies for moisture but not on those for toluene vapor. The high moisture and toluene vapor removal rates and efficiencies under certain particular operating conditions made TEG suitable to be employed in the packed dehumidifiers for better indoor air quality control.

Nevertheless, the characteristics of the TEG solution also present some shortcomings that make it currently unfavorable for indoor air quality control. For example, TEG has a small vapor pressure that may cause some of the glycol to evaporate into the airstream (Grossman and Johannsen 1981; Oberg and Goswami 1998), which may eventually be carried over into the conditioned space and may condense on walls windows and equipment (Abdul-Wahab et al. 2004). High viscosities of the TEG solution also create a higher pumping cost (Chung and Luo 1999). More research should be directed at resolving these technical issues before TEG can be fully used for indoor air quality control.


a = interfacial area on a stage, [m.sup.2]

[a.sub.d] = area of packings, [m.sup.2]

[a.sub.w] = wetted areas of packings, [m.sup.2]

[B] = matrix for Maxwell-Stefan equation

[C.sub.p] = molar heat capacity, kJ/[kmol.sup.-1]*[K.sup.-1]

d = diameter of packing, m

D = diffusion coefficients or mass transfer coefficients, [m.sup.2]*[s.sup.-1]

[e.sub.j] = interphase energy flux of stage j, W*[m.sup.-2]

g =gravity, [m.sup.2]*[s.sup.-1]

G = gas flow rate, [kmol*[s.sup.-1]

h = heat transfer coefficients, W*[m.sup.-2]*[K.sup.-1]

[h.sub.fg] = latent heat of vaporization, kJ*[kmol.sup.-1]

[H.sub.e] = Henry's law constant, Pa, Pa*[m.sup.3]*[kmol.sup.-1]

[[~.H].sub.i,j] = partial molar enthalpy of species i at stage j, kJ*[kmol.sup.-1]

[DELTA][H.sub.s] = integral heat of solution, kJ*[kmol.sup.-1]

[J] = Jacobain matrix

[k.sub.ik] = binary mass transfer coefficients of species i-k, m*[s.sup.-1]

[k] = multi-component mass transfer coefficient matrix, m*[s.sup.-1]

K = equilibrium ratio or partition coefficient

l = liquid component flow rate, kmol*[s.sup.-1]

L = liquid flow rate, kmol*[s.sup.-1]

[m.sub.i]= mass flow air of component i, kg/s

N = mass transfer rate, kmol*[m.sup.-2]*[s.sup.-1]

P = pressure, atm

T =temperature, K

v = vapor component flow rate, kmol*[s.sup.-1]

V = vapor flow rate, kmol*[s.sup.-1]

[w.sub.l] = mass fraction of liquid, dimensionless

q = heat transfer flux, kW*[m.sup.-2]

x = liquid phase composition

y = vapor phase composition

Greek Symbols

[gamma] = activity coefficient, dimensionless

[rho] = density, kg*[m.sup.-3]

[lambda] = thermal conductivity, W*[m.sup.-2]*[K.sup.-1]

[mu] =viscosity, [m.sup.2]/s

[alpha] = heat transfer coefficient, W*[m.sup.-2]*[K.sup.-1]

[epsilon] = interface energy flux, W*[m.sup.-2]

[sigma] = surface tension, N*[m.sup.-1]

[[sigma].sub.c] = critical surface tension, N*[m.sup.-1]

[phi] = association factor, dimensionless

[LAMBDA] = constants used in Wilson equation

[GAMMA] = thermodynamic factors, dimensionless

[THETA] = matrices of correction factors that account for the influence of mass transfer on the mass transfer coefficient

[[PHI].sub.H] = Ackermann correction for mass transfer, dimensionless

[XI] = heat transfer correction factor, dimensionless


i,k = component number

j =stage number

g = vapor phase

l = liquid phase

mix =mixture

[H.sub.2]O = water

TEG = triethylene glycol

air =air

t =total

eff = effective

eqm = equilibrium


I =interphase

L = liquid phase

V = vapor phase

B = bulk phase

sat = saturated pressure

o = standard reference state


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C.K. Chau, PhD

Associate Member ASHRAE

W.M. Worek, PhD


Received October 4, 2006; accepted September 7, 2008

C.K. Chau is an associate professor in the Department of Building Services Engineering, The Hong Kong Polytechnic University, Hung Hom, Hong Kong. W.M. Worek is professor and head of the Department of Mechanical and Industrial Engineering, The University of Illinois at Chicago, Chicago, IL.
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Author:Chau, C.K.; Worek, W.M.
Publication:HVAC & R Research
Article Type:Report
Geographic Code:1USA
Date:Mar 1, 2009
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