A Dissertation Entitled Membrane Drying of Ionic Liquid by Xi Du Submitted to the Graduate Faculty As Partial Fulfillment Of

A Dissertation

entitled

Membrane Drying of Ionic Liquid

Xi Du

Submitted to the Graduate Faculty as partial fulfillment of the requirements for the Doctor of Philosophy Degree in Engineering

1 Dr. G. Glenn Lipscomb, Committee Chair 1 Dr. Sasidhar Varanasi, Committee Member 1 Dr. Maria R. Coleman, Committee Member 1 Dr. Sridhar Viamajala, Committee Member 1 Dr. Yong Wah Kim, Committee Member 1 Dr. Patricia R. Komuniecki, Dean College of Graduate Studies

The University of Toledo December 2012

This document is copyrighted material. Under copyright law, no parts of this document may be reproduced without the expressed permission of the author. An Abstract of

Membrane Drying of Ionic Liquid

Xi Du

Submitted to the Graduate Faculty as partial fulfillment of the requirements for the Doctor of Philosophy Degree in Engineering

The University of Toledo

December 2012

Room temperature ionic liquids (RTILs or, simply, ILs) are liquid salts at temperatures near or slightly above room temperature. ILs consist entirely of bulky, asymmetric organic cations and a variety of anions and, as a consequence, ILs have non-measureable vapor pressure. Due to their unique physical properties, RTILs can be used and recycled as environmentally benign solvents without loss due to evaporation. Envisioned applications range from carbon dioxide capture to battery electrolyte replacement to cellulose dissolution and processing.

Nanostructural organization of aqueous ILs and its change with water concentration are of interest in many recent studies. Water molecules can form hydrogen bond with anions in ILs. At low water concentrations water molecules form complexes mostly with anions rather than with other water molecules and begin to form clusters at high water concentrations.

Diffusion is the macroscopic result of random thermal motion on a microscopic scale which is useful to understand the movement of a single molecule while viscosity is a collective transport property for the entire system. Viscosity can increase dramatically as

iii water concentration decreases and is accompanied by large decreases in water diffusivity and equilibrium water vapor pressure. Additionally, solvation properties can change dramatically and necessitate careful control of water concentration. Experimental measurements of [C2mim][OAc] cation and anion self-diffusivity, water self-diffusivity, and water vapor pressure are investigated. The role of temperature and water concentration on the physical-chemical characteristics is studied.

Pervaporation is a membrane separation process which has been developed for the recovery of dilute solutes from aqueous or organic bulk solvents. The removal of water from aqueous RTIL solutions using pervaporation is reported here. The driving force for separation is the difference in partial pressure of the components on the two sides of the membrane. Pervaporation may prove more versatile than solvent extraction techniques.

Additionally, membrane processes may have dramatically lower energy requirements than distillation if one can identify an appropriate membrane and process for the separation problem.

A number of membranes of varying composition and molecular weight cut-off are evaluated for pervaporation with a dry gas sweep. The dependence of overall water transport rates on temperature, liquid flow and gas flow is evaluated. The mass transfer resistance of each flow channel and the membrane are determined from the results. The effects of membrane properties and the temperature dependence of water vapor pressure on performance are reported.

Dedicated to my family

Acknowledgements

I would like to thank my advisor, Dr. Glenn Lipscomb, for giving me this great opportunity to work with him and carry through this research. His encouragement, patience, and guidance make this research and the completion of this dissertation possible.

I would also like to appreciate my dissertation committee: Dr. Sasidhar Varanasi, Dr.

Maria R. Coleman, Dr. Sridhar Viamajala and Dr. Yong Wah Kim for their advice and contribution to this research.

I want to specially thank Rob Dunmyer and Tom Jacob, who helped me a lot to set up the experiment system and fix problem of the instruments. I got many suggestions and learnt a lot from them.

I gratefully acknowledge SuGanit Systems Inc. and the University of Toledo for providing financial support to this research.

I would also like to thank Pingjiao Hao, Rahul Patil, Sricharan Nanduri, Ashkan

Iranshahi, and Yuecun Lou for helping me through my work. Finally, I thank all friends in the University of Toledo for giving me good memories during the years I am here.

vii

Table of Contents

Abstract iii

Acknowledgements vii

Contents viii

List of Tables xii

List of Figures xiii

List of Symbols xviii viii

1 Introduction ...... 1

1.1 Research Objective ...... 1

1.2 Literature Review ...... 2

1.2.1 Room Temperature Ionic Liquids ...... 3

1.2.2 Challenges of ILs ...... 7

1.2.3 Methods for ILs Recovery ...... 8

1.2.4 Membrane Separation ...... 9

viii

1.2.5 Characteristics of Reverse Osmosis and Pervaporation ...... 11

1.2.6 Mass Transport in Membranes ...... 14

1.2.7 Mass Transport in Pervaporation...... 16

1.2.8 Concentration Polarization ...... 20

1.3 Structure of Thesis ...... 24

2 Vapor Pressure, Conductivity, Viscosity, and Diffusivity of Aqueous 1-ethyl-3- methylimidazoulium Acetate ([C2mim][OAc]) ...... 26

2.1 Introduction ...... 26

2.2 Vapor Pressure of Aqueous [C2mim][OAc] ...... 27

2.2.1 Experiment ...... 27

2.2.2. Results ...... 32

2.3 Conductivity of Aqueous [C2mim][OAc] ...... 33

2.3.1 Experiment ...... 33

2.3.2. Results ...... 33

2.3 Viscosity of Aqueous [C2mim][OAc] ...... 34

2.3.1 Experiment ...... 34

2.3.2. Results ...... 35

2.4 Diffusivity of Aqueous [C2mim][OAc] ...... 36

2.4.1 Experiment ...... 36

2.4.3 Results ...... 39

2.5 Conclusion ...... 48

3 Membrane Drying of Aqueous 1-ethyl-3-methylimidazoulium Acetate ([C2mim][OAc])

...... 50

3.1 Instruction ...... 50

3.2 Straight RO Membrane Separation ...... 51

3.2.1 Experiment ...... 51

3.2.2 Results and Discussion ...... 53

3.3 Pervaporation separation ...... 55

3.3.1 Introduction and Theory ...... 55

3.3.2 Experimental Setup ...... 55

3.3.3 Results ...... 57

3.3.3.1 Membrane Types ...... 57

3.3.3.2 Performance of Aqueous IL Pervaporation with RO AK Membrane...... 60

3.4 Conclusion ...... 62

4 Mass Transfer in Pervaporation for Aqueous Ionic Liquid ...... 63

4.1 Mass Transfer Resistance for Pervaporation ...... 63

4.2 Mass Transfer Coefficient in Gas Phase (Pervaporation of Pure Water) ...... 64

4.2.1 Theory...... 64

4.2.2 Results ...... 65

4.2.2.1 Mass Transfer Coefficient in Gas Phase ...... 65

4.2.2.2 Effect of Temperature on Mass Transfer Coefficient in Gas Phase ...... 73

4.3 Calculation for Mass Transfer Resistance in Liquid Phase and Membrane

(Pervaporation of Aqueous IL) ...... 76

4.3.1 Theory...... 76

4.3.2 Results ...... 77

4.4 Mass Transfer Coefficient in Membrane...... 88

4.4.1 Introduction ...... 88

4.4.2 Results ...... 90

4.4 Conclusion ...... 93

5 Energy Consumption of Vacuum Evaporation vs. Pervaporation for Aqueous IL

Dehydration...... 94

5.1 Introduction ...... 94

5.2 Vacuum Evaporation for Aqueous IL ...... 94

5.2.1 Introduction ...... 94

5.2.2 Theory...... 96

5.2.2.1 Single-effect Evaporator ...... 96

5.2.2.4 Enthalpy for Aqueous 1-ethyl-3-methylimidazolium Acetate ...... 98

5.2.2.5 Energy Consuming from Compressor and Vacuum ...... 98

5.2.3 Results ...... 99

5.2.3.1 Enthalpy of Aqueous IL ...... 99

5.2.3.2 Energy Cost Calculation for Evaporation ...... 100

5.2.3.2.1 Single-effect Evaporator ...... 100

5.2.3.2.2 Double-effect Evaporator ...... 102

5.3 Pervaporation ...... 105

5.3.1 Introduction ...... 105

5.3.2 Membrane Area and Energy Cost Calculation ...... 105

5.3.2.1 Theory ...... 105

5.3.2.2 Results ...... 106

5.4 Conclusion ...... 111

6 Conclusions and Recommendation for Future Work...... 112

6.1 Conclusions ...... 112

6.2 Future Work ...... 113

Reference ...... 115

xii

List of Tables

3.1 IL concentration results of RO separation at 350 psi ...... 54

3.2 IL concentration results of RO separation at 400 psi ...... 54

3.3 IL concentration results of RO separation at 450 psi ...... 54

3.4 Water flux at different IL concentrations for ultralfiltration membrane with MWCO of 1,000 ...... 58

3.5 Water flux at different IL concentrations ...... 59

4.1 Calculation of Rov at different Qgas with inlet DP = -10ºC at 24ºC ...... 68

4.2 Calculation of Rov at different Qgas with inlet DP = -15ºC at 24ºC ...... 69

4.3 Calculation of Rov at different Qgas with inlet DP = -20ºC at 24ºC ...... 70

4.4 Calculation of Rov at different Qgas with inlet DP = -60ºC at 24ºC ...... 71

4.5 Parameters of mass transfer coefficient in gas phase with different inlet dew points at 24ºC...... 73

4.6 Calculation of Rov at different Qgas with inlet dew point of -60ºC at 40ºC ...... 75

4.7 Parameters of mass transfer coefficient in gas phase with nitrogen sweep at 40ºC . 76

4.8 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 60 ml/min ...... 79

4.9 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 90 ml/min ...... 80

4.10 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for

14.0 mole% of IL with QIL = 120 ml/min ...... 81

4.11 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for

14.0 mole% of IL with QIL = 150 ml/min ...... 82

4.12 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for

14.0 mole% of IL with QIL = 180 ml/min ...... 83

4.13 Parameters for mass transfer resistance in membrane and liquid side at 24ºC ...... 86

5.1 Energy cost for vacuum of single-effect evaporator ...... 102

5.2 Energy cost for vacuum of effect II evaporator ...... 104

5.3 Calculation process of integration in the first concentration interval ...... 109

vii

List of Figures

1-1 Schematic diagram of Reverse Osmosis ...... 11

1-2 Schematic diagram of the pervaporation process. (a) Vacuum pervaporation, (b)

Purge gas pervaporation ...... 12

1-3 Schematic representation of the membrane transport mechanism. (a) pore-flow model, (b) solution-diffusion model ...... 15

1-4 Chemical potential, pressure, and activity profiles through a pervaporation membrane ...... 17

1-5 Concentration gradients adjacent to the membrane ...... 21

2-1 Figure of experimental apparatus used to measure the vapor pressure of aqueous IL

(IL molar concentration < 20%) ...... 28

2-2 Water vapor pressure test for IL-water system (IL molar concentration > 20%)..... 30

2-3 Vapor pressure of IL-water at different concentrations and temperatures ...... 32

2-4 Conductivity of aqueous IL at 40ºC ...... 33

2-5 Viscosity test of aqueous [C2mim][OAc] at different temperatures...... 35

2-6 Full range of 1H NMR spectrum of [C2mim][OAc] containing water ...... 41

2-7 2D DOSY Plots of aqueous [C2mim][OAc] with IL mole concentration = 18.09% at

20ºC ...... 42

viii

2-8 Diffusivity vs. IL concentrations for cation, anion and water in IL-water system at

40ºC obtained using DgcstelSL_CC NMR ...... 43

+ - 2-9 Temperature vs. diffusivity for both [C2mim] and [OAc] in pure 1-ethyl-3- methylimidazolium acetate obtained using DgcstelSL_CC NMR ...... 45

+ - 2-10 Temperature vs. diffusivity for [C2mim] , [OAc] and water in aqueous 1-ethyl-3- methylimidazolium acetate (IL mole% = 65.98) obtained using DgcstelSL_CC NMR .. 46

+ - 2-11 Temperature vs. diffusivity for [C2mim] , [OAc] and water in aqueous 1-ethyl-3- methylimidazolium acetate (IL mole% = 6.79) obtained using DgcstelSL_CC NMR .... 47

3-1 Schematic of the reverse osmosis process ...... 51

3-2 Water flux measurement from straight RO membrane ...... 52

3-3 Pervaporation for aqueous IL ...... 56

3-4 Schematic diagram of IL leaking using polysulfone ultra-filtration membrane ...... 58

3-5 Schematic diagram of boundary layer increment due to IL penetration ...... 59

3-6 Water flux vs. IL concentration at different flow rate in feed and permeate channel at room temperature ...... 60

3-7 Water flux vs. IL concentration at different operation temperatures ...... 61

4-1 Water flux vs. gas flow rates with different inlet dew points ...... 66

4-2 Water flux vs. gas flow rates with N2 sweep at 40ºC ...... 74

4-3 Jwater vs. QIL with different nitrogen flow rates (DP=-60ºC) at room temperature for

14.0 mole% of IL...... 78

4-4 Henry’s law constant for aqueous IL vs. IL mole concentrations at 24°C and 40°C 85

4-5 Mass transfer resistance vs. IL mole concentration at 24ºC with QN2 = 5 L/min, and

QIL = 180 ml/min ...... 87

4-7 Schematic view of the RO AK membrane ...... 89

4-8 Mass transfer coefficient in membrane at 24ºC ...... 91

4-9 Mass transfer coefficient in membrane at 40ºC ...... 92

5-1 Material and enthalpy balances in evaporator ...... 96

5-3 Enthalpy of aqueous IL at different temperatures ...... 100

5-2 Pervaporation of aqueous IL...... 105

5-3 Regression of water partial pressure vs. IL weight fraction at 40°C ...... 107

5-4 Regression of overall mass transfer coefficient vs. IL weight fraction at 40°C ..... 108

Chapter 1

Introduction

1.1 Research Objective

The objective of the proposed project is to investigate and optimize processes for concentrating aqueous IL (ionic liquid), especially membrane processes. Optimization of membrane processes includes selection of membranes, flow rates on liquid and gas sides, and operating temperatures. The permeability, transport properties of the membranes and chemical-physical properties of aqueous IL are critical to process viability. Specific tasks to be completed in this work are:

(i) Measure the water vapor pressure of IL-water at a series of temperatures and

concentrations.

(ii) Analyze the self-diffusivity of IL-water system using NMR technique and

investigate the nanostructure of aqueous IL.

(iii) Measure the viscosity and conductivity of IL-water system.

(iv) Compare the performance of two different membrane processes: reverse

osmosis and pervaporation.

(v) Determine the water flux at different operation conditions and estimate the

mass transfer resistance of each channel as well as the membrane in

pervaporation system.

(vi) Design a membrane separation and a vacuum evaporation process for

dehydration of aqueous IL. Compare the energy cost of pervaporation with

that of vacuum evaporation.

This project investigates a novel process for IL recovery. The physical and chemical properties of 1-Ethyl-3-methylimidizolium acetate are investigated for future application.

The transport properties of this ionic liquid through membranes are evaluated. The economics of pervaporation is compared with vapor-liquid equilibrium process.

1.2 Literature Review

RTILs or ILs are molten salts at temperatures around room temperature which are composed of an organic cation and an inorganic anion. They have negligible vapor pressure and represent environmentally safe and non-corrosive media [ 1 , 2 , 3 ].

Many unique properties of ILs have raised increasing interest in their applications in organic synthesis, catalytic reactions, separation processes, and extraction [4, 5]. ILs are usually hygroscopic and various properties of ILs such as solubility, polarity, viscosity, vapor pressure and diffusivity highly depend on water concentration, hence the water content in ILs needs to be carefully controlled. As ILs are still quite expensive media, their recovery is essential for their use to be economically viable [6].

Traditional recovery methods for aqueous ILs are distillation [7] and extraction [8].

However, distillation is energy consuming and might cause thermal degradation of ILs [9] while extraction needs to add auxiliary substance which introduces further separation.

During the last two decades, membrane processes have been of great importance for the purification and recycling of a large variety of different compounds [10, 11]. Reverse osmosis (RO) is normally used for concentrating aqueous solutions containing a low molecular weight solute and pervaporation (PV) can be used for dehydration of solvents and other volatile organics [12].

This work seeks to investigate the physical-chemical properties of 1-ethyl-3-methyl imidazolium acetate ionic liquid + water system including viscosity, vapor pressure, and diffusivity. The nanostructure of IL and water is helpful to understand the macro behavior of the system. The separation performances of reverse osmosis and pervaporation are evaluated. The mass transfer resistances of each channel and the membrane are studied for pervaporation. Additionally, the effect of temperature on the mass transfer coefficients is also studied. Finally, the energy cost of pervaporation is calculated and compared with traditional thermal based evaporation process.

1.2.1 Room Temperature Ionic Liquids

Recently, researchers have started to explore a particular class of liquids called room temperature ionic liquids or simply ionic liquids to investigate their unique properties for different applications. An ionic liquid generally consists of a large nitrogen-containing organic cation and a smaller inorganic or organic anion. Its asymmetric structure reduces the lattice energy of the crystalline structure and results in melting points as low as -96ºC.

ILs consist entirely of ions and therefore are considered an environmentally benign alternative to traditional solvents since they do not have a measurable volatility. ILs may be utilized as extractive media in liquid/liquid extraction processes. They are good solvents for a wide range of inorganic, organic, and polymeric materials. Due to their unique structures, ILs have also been widely exploited in applications in the fields of battery electrolytes [13], biocatalysis [14] and carbon dioxide sequestration [15].

High viscosity is inherent to many ionic liquids. Strong intermolecular forces (charge- charge interaction) between ions cause high viscosity. It is known the viscosity is highly variable among ILs [16]. IL viscosity depends on the cation and anion combination.

Although typical organic solvents may exhibit viscosities ranging from 0.2 to 10 cP, viscosities of ILs can have variations spanning several orders of magnitude.

ILs usually are hygroscopic which can absorb a significant amount of water [17] and sometimes show an ability to absorb water from the surrounding gas phase [18]. The presence of water in the ILs can dramatically affect their physical properties [19].

Viscosity can increase steeply as water concentration decreases and is accompanied by large decreases in water diffusivity and equilibrium water vapor pressure. Additionally, solvation properties can change dramatically and necessitate careful control of water concentration. When researchers controlled the water activity in an ionic liquid, the driest conditions gave the highest enzyme activity [20, 21]. An addition of water also can change chemical reactions of solutes in ILs. In the example of palladium catalyzed toluene oxygenation in 1-n-butyl-3-methylimidazolium tetrafluoroborate ([bmim][BF4]), water addition can lead to the main product from benzaldehyde to benzoic acid [22].

Moreover, like conventional liquids, the viscosity of ILs is a strong function of

4 temperature. Other physical and chemical properties of ILs also can change dramatically with system temperature.

Another important factor responsible for the increasing popularity of ILs lies in their extremely low vapor pressure [23]. Negligible vapor pressure eases ILs processing and purification, facilitates their use in the reaction and extraction procedures, and has led to the recognition as environmentally friendly solvents. The vapor pressure and its dependence on solute concentration, temperature and the enthalpy for vaporization are among those fundamental properties needed to be investigated to contribute to the development of reasonable molecular models.

The study of diffusion in ILs receives increasing interest as slow diffusion restricts their applicability in practice. A thorough understanding of diffusion in ILs is therefore highly desirable. Self-diffusion of ILs-containing systems was extensively studied in the past.

This quantity can be measured using pulsed-field gradient nuclear magnetic resonance

(PFG-NMR) spectroscopy.

Nanostructural organization of aqueous ILs and its changes with water concentrations has been investigated in many recent researches [ 24 ]. Particularly, studies from IR spectroscopy and density functional calculations revealed water molecules can form hydrogen bond with anions in imidazolium-based ILs [25]. These studies showed that at low water concentrations water exists in a free (not self-associated) state and water molecules prefer to form complexes mostly with anions rather than with other water molecules. Those results are in a qualitative agreement with the study of recent molecular dynamics (MD) simulations, which show the tendency that water preferentially solvates

5 the anions and isolates from each other at small water concentrations [26]. On the contrary, water molecules tend to form clusters and combine with other water molecules at high water molar concentrations exceeding 75% [ 27 ]. The physical-chemical properties of aqueous ILs are vital parts of understanding how ILs participate as components in a mixed solvent system.

In recent years, ILs based on the 1-alkyl-3-methylimidazolium cation [Cnmim] have received high attention. It was found that the [Cnmim] cations possess an inherent amphiphilicity. Therefore, it can be anticipated that those ILs will exhibit the interfacial and aggregation behavior analogous to that displayed by conventional cationic surfactants.

It may be expected that the amphiphilicity of the [Cnmim]-based ILs signifies that interfacial phenomena will play a major role in the physical and chemical properties, the synthesis and purification, and the formation of dispersed or phase-separated systems in

IL-containing system. Some of these points have guidance on the chemical reactivity of the IL-containing systems and have also generated interest in the theoretical aspects of the solvation of ILs [28].

Interest in ILs with acetate anions stems from the previous studies of the dissolution of cellulose using 1-alkyl-3-methylimidazolium chloride [29] and the subsequent explosion of the ILs pretreatment in biomass processing to generate energy products [30]. The alternative ILs with anions of dimethylphosphate and acetate could also dissolve polysaccharides and provide comparable solubilizing characteristics [ 31 ]. 1-ethyl-3- methylimidazolium acetate ([C2mim][OAc]) is perhaps the “best” cellulosic solvent [32] and also exhibits enzyme and environmentally-friendly characteristic [33] and significant solubility of CO2 [34].

However, imidazolium cation based ILs are found to thermally degrade, yielding volatile degradation products at certain temperatures. Significant evolution of degradation by- products can take place at temperatures much lower than previously predicted [35]. For example, [C2mim]Cl thermally decomposes when heated to 190ºC in vacuum and the volatile decomposition products can be condensed to generate a new ternary ionic liquid system. Consequently, the operation temperature for ILs needs to be carefully controlled.

1.2.2 Challenges of ILs

The unique properties of ILs and the ability to design their properties by appropriate choice of anion, cation and substituents create many more processing options for applications. However, high cost, lack of physical and chemical property and toxicity data restrict the advantageous use of ILs as process chemicals.

The essential challenges for using ILs in the chemical industries must be also addressed as well as their advantages. The major challenge is cost. The price of a kilogram of ILs is about 30,000 times greater than a common organic solvent such as acetone. Though proper choice for composition of the cation and anion and the scale of production can reduce the price of the ILs [36, 37], the cost is still discouraging. Researches need to find alternative ways to recycle ILs as many processes for cleaning up ILs involve washing with water or VOCs which creates another waste stream.

Another challenge for the application of ILs is the incomplete physico-chemical data. At the present most available data are focused on bulk physical properties such as melting point, viscosity, and density. Relatively little is studies about the microscopic physical properties. These properties are helpful to investigate the influence of ILs on chemical

7 reaction rate and new ILs design with precisely tailored properties as well as their specific applications.

1.2.3 Methods for ILs Recovery

RTILs are more expensive than traditional solvents. Reuse and recovery of RTILs is essential to satisfy both economic and environmental requirements for sustainable processes but largely remains an unresolved issue.

The impact of structural characteristics of different ILs on their solubilities in water is related to further applications such as selecting a proper treatment process, and understanding their environmental fate. The loss of ILs into the aqueous phase may be an important factor in estimating the costs of recycling and water treatment.

Freire et al. demonstrated that the mutual solubilities between imidazolium-based ILs and water are primarily defined by the anion hydrophobicity followed by the cation alkyl side chain length [38]. The miscible ionic liquid is highly polar and therefore soluble in water.

The recovery of IL wastes mainly focuses on wastewater streams due to the physico- chemical properties and applications of ILs such as biomass pretreatment. The amount of

ILs in wastewater streams can be as large as the concentration equivalent to their solubilities. Concentrations of hydrophobic ILs in wastewater can be relatively high even though poor solubilities.

Since ionic liquids have very low melting points, they exist in a liquid phase at room temperature or lower temperature, making it difficult to purify by recrystallization.

Distillation and extraction are the primary techniques used for IL purification.

Distillation is energy-intensive as energy is required to evaporate the solvent(s) and to heat the bulk ionic liquid. Consequently, distillation may be expensive especially at low

IL concentrations. Moreover, the severe drop in water vapor pressure for IL-water systems at high IL concentrations can cause significant problems in vapor-liquid equilibrium separation process and impose limitations on the purity that can be achieved.

Some ILs show evidence of decomposition to volatile products in distillation, especially for ILs with halides, sulphates or carboxylates anions. The principal mechanism of decomposition is believed to be dealkylation or transalkylation of the cation [39, 40].

Extraction processes offer potential to purify ILs but process efficacy is limited by the recovery of ILs and the cross contamination of both phases. Supercritical carbon dioxide extraction is environmental clean, but it is technically demanding and involves operation at elevated pressure which incurs additional processing costs. As the extraction fluid directly contacts the ionic liquid, it may be difficult to guarantee that there is no loss of ionic liquid into the supercritical fluid, which represents an important limitation.

New separation methods that are less energy consuming, providing high recovery and operating under mild conditions are demanded.

1.2.4 Membrane Separation

A membrane acts as a thin barrier allowing selective passage of different species through it. This difference in selectivity is utilized for separations from biotechnology to water treatment. These membranes commonly are made with polymeric materials in flat sheet, spiral-wound, tubular and hollow fiber form. Present day industrial membrane processes involves microfiltration, ultrafiltration, nanofiltration and reverse osmosis. Other

9 commercial membrane processes are electrodialysis, membrane electrolysis, diffusion dialysis, pervaporation, vapor permeation and gas separation.

Reverse osmosis (RO) is used for aqueous solutions containing a low molecular weight solute. The process involves the application of pressure to the liquid feed mixture forcing smaller components of the solution to pass and leaving large molecules as a retentate.

Reverse osmosis is most commonly used in drinking water purification from seawater.

ILs can be purified by RO technique. However, the huge energy costs from the need to operate at high pressure must be considered.

Pervaporation (PV) is a membrane process in which the feed side is a liquid while the permeate side is a vapor. Phase change is obtained by lowering the partial pressure of the permeate side either by gas-sweeping or with a vacuum pump. The separation principle of pervaporation is based on the preferential partitioning of a solute from a liquid feed phase in the membrane through which it diffuses according to its chemical potential gradient. The driving force for species to transport is the difference in partial pressure between both sides of the membrane. Pervaporation is governed by solute-polymer interactions instead of vapor-liquid equilibrium as for the case of distillation.

Pervaporation is a promising alternative to distillation and extraction for separation of low molecular weight molecules. It becomes an emerging membrane technology for liquid separation due to the advantages of high selectivity, energy-efficiency and eco- friendliness for molecular-level liquid separation. Pervaporation offers the potential to reduce energy consumption and operate at milder conditions to avoid degradation of temperature sensitive compounds. Pervaporation also may prove more versatile than

10 solvent extraction if an appropriate membrane can be identified for the separation problem. The separation potential of pervaporation may be much higher than that of reverse osmosis, though both processes are based on the same transport mechanism.

However, the rates of transport in both processes are approximately equal.

In general, the use of membrane technology for IL recovery is regarded as advantageous in comparison to distillation and extraction. In membrane separations, processes operate under mild conditions and no auxiliary substance is required for phase addition.

1.2.5 Characteristics of Reverse Osmosis and Pervaporation

Most reverse osmosis and pervaporation membranes use dense membrane to perform the separation and thus, they are promising method for liquid separations, especially for small molecular weight components removal from a liquid mixture.

In reverse osmosis process, the pressure exerted is greater than the osmotic pressure of the feed, causing solvent to flow through the membrane. Figure 1-1 provides a schematic representation of the process.

Figure 1-1 Schematic diagram of Reverse Osmosis

Reverse Osmosis process is a modern technology to purify water for a wide range of applications. The driving force of the reverse osmosis process is applied pressure in the feed side.

Pervaporation is a relatively new membrane process in which the feed side is a liquid mixture while the permeate side is a vapor due to application of a vacuum or use of a purge gas (Figure 1-2). The permeate vapor can be condensed and collected or released as desired.

Figure 1-2 Schematic diagram of the pervaporation process. (a) Vacuum pervaporation,

(b) Purge gas pervaporation

A fundamental understanding of pervaporation is still emerging. Pervaporation is commonly described as a three-step process: solution-diffusion-evaporation. Separation takes place because of selective solution and diffusion. The separation principle is based on the physical-chemical interactions of the membrane material and the permeating molecules, not the relative volatility as in distillation. Therefore, pervaporation is a beneficial complement to energy-intensive distillation in azeotropic and close-boiling mixtures separation.

Pervaporation involves a phase change of permeating species from liquid to vapor state and vaporization energy is supplied during the process. The heat of vaporization required for permeation can be supplied either in the feed liquid or directly heating the membrane.

Hence, pervaporation is not an isothermal process, even if the temperature of the incoming feed is carefully controlled. Vaporization of the permeate at the downstream membrane surface causes local cooling to induce a temperature gradient across the membrane. As a consequence, the temperature of the flowing retentate also decreases as it proceeds along the module.

Pervaporation is commonly operated at low feed pressures and ambient temperatures.

Moreover, no additional chemicals are needed for separation. Therefore, pervaporation can be applied in separation of labile components.

Transport is driven by osmotic pressure for reverse osmosis while separation of pervaporation is achieved by lowering the chemical potential of the permeate stream on the downstream side. Separation potential of pervaporation is higher compared with reverse osmosis separation. For example, both the permeation flux and separation factor

13 in pervaporation are higher than reverse osmosis to concentrate ethanol in aqueous solution from 85% to more than 99 wt % [41].

1.2.6 Mass Transport in Membranes

A proper understanding of the mechanism of permeation through a membrane may provide direct information on research and development of an appropriate membrane.

There are two models [42] to describe mass transport in membranes as in Figure 1-3: (a) the pore flow model and (ii) the solution-diffusion model.

(a)

(b) Figure 1-3 Schematic representation of the membrane transport mechanism. (a) pore- flow model, (b) solution-diffusion model

In pore-flow model, permeants are transported by pressure-driven convective flow through pores. Separation occurs as the larger components are excluded from some of the pores in the membrane while other components transport through. In solution-diffusion model, permeants dissolve in the membrane and then diffuse through it down a concentration gradient. Separation occurs due to the differences in the solubilities of the species in the membrane and the differences in diffusion rates through the membrane.

The solution-diffusion model is more widely accepted to explain transport in reverse osmosis, gas permeation, and pervaporation. The overall driving force producing movement of a component through a membrane is the chemical potential gradient which is related to pressure, temperature, concentration, and electrical potential. The flux, Ji

(g/cm2·s), of i component, is described by the equation

d J  L i (1) i i dx

Where dµi/dx is the chemical potential gradient of component i across the membrane and

Li is a proportional coefficient related to chemical potential gradient and flux. Chemical potential gradient can be expressed in terms of differences in concentration, pressure, temperature, and electrical potential depending on different systems. For example, if driving force is generated by concentration and pressure gradient, the chemical potential can be written as

di  RTd ln( i xi )  vi dp (2) where xi is the mole fraction of component i, γi is the activity coefficient, p is the pressure, and vi is the molar volume of component i.

For incompressible phases as a liquid or a solid membrane, volume does not change with pressure. Integrate Equation (2) with respect to concentration and pressure as

   0  RT ln( x )  v ( p  p ) i i i i i i sat (3)

0 where µi is the chemical potential of pure i at the saturation vapor pressure, pisat.

For compressible phases as gases, the molar volume changes with pressure. Assume ideal gas to integrate Equation (2) gives

p    0  RT ln( x )  RT ln i i i i p i sat (4)

1.2.7 Mass Transport in Pervaporation

Based on solution-diffusion model, the penetrant species is assumed to selectively absorb at the upstream face of the membrane and then diffuse through the membrane, followed

16 by desorption at the downstream side. The bulk fluids on either side of the membrane are assumed to be in equilibrium with the membrane material at the interface which means the gradient of chemical potential in the membrane is continuous. The gradients of chemical potential, pressure, and activity across the membrane are illustrated in Figure 1-

4 [42]:

Figure 1-4 Chemical potential, pressure, and activity profiles through a pervaporation membrane

At the feed-membrane interface, the chemical potential of a solute in the feed liquid is equal to the chemical potential of the solute in the membrane. According to Equation (3), we can derive:

0 L 0 i  RT ln( i o xi o )  vi ( po  pi sat )  i  RT ln( i o(m) xi o(m) )  vi ( po  pi sat ) (5)

The term o represents the positions of the feed interface of the membrane, and the super script L is used here and later to represent liquid phase activity coefficient, sorption

L coefficient and permeability coefficient. Thus the term γio represents the activity

17 coefficient of component i in the liquid in contact with the membrane at the feed interface.

The subscript m represents the membrane phase. Thus, γio(m) is the activity coefficient of component i in the membrane at the feed interface. Equation (5) leads to

L ln( i o xi o )  ln( i o(m) xi o(m) ) (6) and thus

 i o(m) xi o  L xi o(m) (7)  i o

3 The more practical term concentration ci (g/cm ) is defined as

ci  mi xi (8)

3 where mi is the molecular weight of i (g/mol) and ρ is the molar density (mol/cm ),

Equation (7) can be written as

L  i o  m L ci o(m)  ci o  Ki ci o (9)  i o(m) 0

L where Ki is the liquid phase sorption coefficient.

At the permeate-membrane interface, the pressure drops from p0 in the membrane to pl in the permeate vapor. The chemical potential relationship between gas and membrane at the interface is

0 G pl 0 i  RT ln( i l xil )  RT ln( )  i  RT ln( i l(m) xil(m) )  vi ( po  pi sat ) (10) pisat

Rearranging Equation (10) gives

G  i l pl  vi ( po  pi sat ) xil(m)    xi l exp{ } (11)  i l(m) pisat RT

Because the term –vi(po–pl)/RT is small, the exponential term in Equation (11) is close to one, and Equation (11) can be written as

G  i l pl xil(m)    xi l (12)  i l(m) pisat

The product xil pl can be replaced by the partial pressure pil, thus

G  i l pi l xil(m)   (13)  i l(m) pisat substituting concentration for mole fraction from Equation (8),

G  il pil G cil(m)  mi m   Ki pil (14)  il(m) pisat

G where Ki is the gas phase sorption coefficient.

According to Fick’s law,

dc J  D i (15) i i dx

Integrating over the thickness of the membrane, then,

D (c  c ) J  i i0(m) il(m) (16) i l

Substitute Equation (9) and (14) in to Equation (16) as:

D (K Lc  K G p ) J  i i i0 i il (17) i l

The interconversion of liquid and gas phase sorption coefficient can be derived by assuming a hypothetical vapor in equilibrium with a liquid solution. The relationship can be written

0 L L 0 G G p i  RT ln( i xi )  vi ( p  pisat )  i  RT ln( i xi )  RT ln( ) (18) pisat

Follow the same steps from Equation (10) to (14), Equation (18) becomes

G G L  i pi Ki ci  mi  L  L  pi (19)  i pisat Ki

Substitution of Equation (19) into Equation (17) yields

D K G ( p  p ) PG J  i i i0 il  i ( p  p ) (20) i l l i0 il

Based on Henry’s law

L ci  H i  pi (21) then the Henry’s law constant Hi in Equation (19) can be written as

L L Ki  i pisat Hi  G  G (22) Ki mi  i

Equation (20) can also be written as

PG P L J  i (c H  p )  i (c  p / H ) (23) i l i0 i il l i0 il i

Equation (20) expresses the driving force in pervaporation in terms of the vapor pressure

G difference. Pi is the permeability coefficient of component i.

1.2.8 Concentration Polarization

During separation by pervaporation, concentration gradients form in the fluids on both sides of the membrane because the feed mixture components permeate at different rates.

The phenomenon is called concentration polarization. Concentration polarization [43] affects membrane separation processes and reduces the effective permeance and selectivity. There is an accumulation of the less permeable species and a decrease in

20 concentration of the more permeable components in the boundary layer adjacent to the membrane. Figure 1-5 illustrates the boundary layer of the more permeable component that forms at the membrane surface. As ionic liquid cannot evaporate and permeate through the membrane, water is the more permeable component in pervaporation.

1 2 Feed Membrane Permeate

c1 m cb1f

c1 f

c2 m

p c2 cb2p

Figure 1-5 Concentration gradients adjacent to the membrane

In Figure 1-5, 1 and 2 represent the interfaces of membrane with the feed liquid and

f p permeate gas, respectively. cb1 and cb2 are the concentration of water in the bulk feed

f p and permeate, respectively. c1 and c2 are concentration of water in contact with the

m m membrane in the feed and permeate, respectively. c1 and c2 are concentration of water in the membrane at the interface in contact with the feed and permeate, respectively.

f m Based on Equation (9), c1 is equilibrium with c1 , and the relationship can be written as:

m r f c1  H c1 (24) where Hr is the partition coefficient, the ratio at equilibrium of concentration of water in the membrane on surface 1 to the concentration in liquid phase at the feed interface.

p m Similarly, c2 is equilibrium with c2 as:

m p p c2  H c2 (25) where Hp is the Henry’s law constant, the ratio at equilibrium of concentration of water in the membrane on surface 2 to the concentration in gas phase at the permeate interface.

The flux of water across the feed side boundary layer J can be written as

f f f f J  k fB (cb1  c1 )  k fB  H'(pb1  p1 ) (26) where kfB is the mass transfer coefficient in the feed side, and H' is the Henry’s law constant which is the ratio at equilibrium of concentration of water to the partial pressure of water in the bulk. The flux across the boundary layer of the permeate side can be written as

p p p p k J  k (c p  c p )  k ( 2  b2 )  pB ( p p  p p ) (27) pB 2 b2 pB RT RT RT 2 b2

p p where kpB is the mass transfer coefficient in the permeate side, and p2 and pb2 are the

p p water partial pressures in equilibrium with c2 and cb2 respectively. The flux across the membrane can be written as

D D J  k ( p m  p m )  (H'' p f  H'' p p )   H''( p f  p p ) (28) m 1 2 l 1 2 l 1 2 where km is the mass transfer coefficient of membrane which equals the ratio of membrane diffusivity D to membrane thickness l. H'' is the partition coefficient, the ratio

22 of partial pressure of water in the membrane to the partial pressure of water at the interface outside the membrane.

f f Combine Equations (26), (27) and (28), and cancel p1 and p2 on the right sides of the equations:

1 1 1 J f p J (   )   pb1  pb2  p (29) k fBH ' H ''(D / l) k pB / RT K

The “resistances-in-series” model has been used to describe the effect of concentration polarization. As shown in Figure 1-5, we can distinguish three different resistances related to the transport:

 the boundary layer in the feed side adjacent to the membrane;

 the membrane;

 the boundary layer in the permeate side adjacent to the membrane.

Therefore, the overall resistance Rov consists of the resistance Rm in the membrane together with the resistances RfB and RpB in the boundary layers on the feed and permeate sides, respectively:

1 1 1 Rov  R fB  Rm  R pB    (30) k fB H ' H ''(D / l) k pB / RT

The boundary layer mass transfer coefficients are known from experiment which are determined by the fluid flow velocity, the relationship can be given as:

n kB  k0Q (31)

23 where Q is the fluid flow rate through the membrane module, k0 and n are experimentally dependent on operation conditions such as temperature, fluid compositions and so on.

The resistance of the boundary layer can be expressed by

n RB  k0Q (32)

1.3 Structure of Thesis

Chapter 1 (this chapter) provides an introduction to the characteristics of ILs, the traditional recovery method for aqueous ILs (distillation and extraction), membrane separation process, and the theory of membrane separation especially reverse osmosis and pervaporation. The “resistances in series” model used to evaluate the mass transfer coefficients is reviewed as well.

Chapter 2 investigates the physical-chemical properties of aqueous ionic liquid. The viscosity, vapor pressure, conductivity and diffusivity of the aqueous ionic at different temperatures within the full range of ILs concentrations are tested. The relationship of viscosity and diffusivity can be quantifies using Stokes-Einstein relation. Those parameters of physico-chemical properties can help understanding of the nanostructure organization of aqueous ILs system. Moreover, the vapor pressure data for aqueous IL could be used to analyze the mass transport mechanism for membrane separation of the system.

Chapter 3 claims two types of membrane techniques for aqueous ILs purification: reverse osmosis and pervaporation. The first part of this chapter reports the performance of reverse osmosis at different operation pressures. The second part reports the performance

24 of pervaporation at different flow rates on both channels at varied temperatures. The processes of reverse osmosis and pervaporation are compared.

Chapter 4 investigates the mass transport for pervaporation dehydration of aqueous ILs.

The mass transfer resistance in the liquid channel, the gas channel and the membrane are calculated individually. The impact of IL concentrations, liquid and gas flow rates and temperature on water flux and mass transfer resistance is also analyzed. The characteristic of the RO AK membrane is investigated to understand the transport properties of the membrane.

Chapter 5 establishes two systems for 1-ethyl-3-methulimidazolium acetate recovery from water: vacuum evaporation and membrane pervaporation. The processing costs of both systems are calculated and compared.

Chapter 6 is the conclusions and recommendations for future work.

Chapter 2

Vapor Pressure, Conductivity, Viscosity, and Diffusivity of Aqueous 1- ethyl-3-methylimidazoulium Acetate ([C2mim][OAc])

2.1 Introduction

Ionic liquids are attracting great attentions as a potential green alternative to volatile molecular organic solvents to be applied in different areas including catalytic and organic reactions and electrochemical and separation processes [44]. Over 200 types of ionic liquids are known but for most of them physico-chemical data are incomplete or lacking.

The number of potential ionic liquids is huge (evaluated as > 1014), however, generally only a few imidazolium-based salts are used in synthesis. The physico-chemical properties of the most used ionic liquids, which are relevant to synthesis and operation, are critical to investigate. The presence of water and organic solvents in IL not only reduce the purity of IL but also affect the physico-chemical properties of IL and further applications [45, 46]. For instance, the presence of water could reduce the activity of catalysts for biomass pretreatment [47].

To facilitate the selection of an optimum IL for a particular application, it is useful to consider their transport and thermodynamic properties. This chapter will discuss the physic-chemical properties of vapor pressure, conductivity, viscosity and diffusivity for aqueous 1-ethyl-3-methylimidazoulium acetate ([C2mim][OAc]). The properties of the

26 system can supply information to understand the microstructures and interactions of ILs at molecular level, as pure compounds or in the presence of dissolved species. The chemical constitution of ILs determines the nature of intramolecular and intermolecular interactions which results in the impact on the macroscopically observable properties such as thermodynamic and transport properties [48].

2.2 Vapor Pressure of Aqueous [C2mim][OAc]

2.2.1 Experiment

The ionic liquid in our research is 1-ethyl-3-methylimidazolium acetate ([C2mim][OAc]) which is completely miscible with water over the entire composition range at the temperatures of study (293.25K-333.25K).

The driving force for pervaporation of IL-water mixtures is dependent on the water vapor pressure difference between the bulk of the liquid side and the permeate side. The extremely low water vapor pressure of IL-water systems, especially at high IL concentrations, makes it difficult to separate water from ILs. Therefore, the water vapor pressure of IL-water mixtures is a critical parameter and must be characterized to design a separation process.

For low IL concentrations (IL molar concentration < 20%), the water vapor pressure of the mixture at varied concentrations and temperatures was measured by Vernier. When a liquid was added to the flask as in Figure 2-1, it would evaporate into the air in the flask until vapor-liquid equilibrium reached between the rate of evaporation and the rate of condensation. At the point of equilibrium, the vapor pressure of the liquid is equal to the partial pressure of the vapor in the flask. Pressure and temperature data were collected

27 using a Gas Pressure Sensor and a Temperature Probe. The flask was placed in a water bath of constant temperature. The effect of temperature on vapor pressure was determined by changing the bath temperature. A diagraph of the experimental apparatus is provided as follows [49]:

Figure 2-1 Figure of experimental apparatus used to measure the vapor pressure of aqueous IL (IL molar concentration < 20%)

The experimental procedure consists of the following steps:

1. Set the bath temperature to be 60ºC, connect the temperature probe to the

computer.

2. Purge the flask with nitrogen gas for at least one hour and then connect the flask

with the Gas Pressure Sensor.

3. Install the flask into the water bath and connect the Gas Pressure Sensor to the

computer.

4. Open the “Logger Pro” software and collect data. Wait until the temperature and

vapor pressure become stable.

5. Introduce the aqueous IL into the flask by pushing in the plunger of the syringe.

Return the plunger of the syringe back to the initial mark, then close the 2-way

valve.

6. Monitor and collect pressure and temperature data.

Repeat all above steps to test another data at a different concentration or temperature.

For higher IL concentrations (IL molar concentration > 20%), high resolution vapor pressure sensor is needed to measure extremely low vapor pressure. A Setra CDG vacuum pressure gauge (model #: 769) which has a range of 10 Torr was used for vacuum measurements. An RO membrane cell was modified to measure the water vapor pressure of IL-water mixtures at different concentrations and temperatures. Polyamide

RO AK membrane from GE labstore was used to test the vapor pressure. The experiment set up is shown in Figure 2-2:

Figure 2-2 Water vapor pressure test for IL-water system (IL molar concentration >

20%)

As IL does not permeate through the RO membrane, the water vapor pressure of the permeate side reveals the partial pressure of water in IL-water system. The experimental procedure consists of the following steps:

1. Pump the feed side of the membrane cell with an IL-water mixture.

2. Turn on vacuum pump of the permeate side until the pressure goes to minimum

and becomes constant.

3. Close valve of the permeate side and turn off vacuum pump.

4. Wait until the pressure transducer displays a constant value.

5. Adjust the oven temperature to test another water vapor pressure at different

temperatures.

Repeat all the steps to test any IL-water concentration at different temperatures.

The vapor pressure of aqueous IL with low water content was also measured and verified with moisture sensor (Dewpoint transmitter) with model of DMT 242 from Vaisala. The aqueous IL is maintained at certain temperature and the moisture sensor is directly contacting the vapor phase above the aqueous IL in a sealed tube. As only water evaporate, the water content tested by the moisture sensor stands for the vapor pressure of the system.

Goff-Gratch water dew point correlations [50] as Equation (36), (37), (38) are used to convert dew points to water vapor concentrations.

log p  7.90298(a 1)  5.02808log (a) 1.3816107 (1011.344(11/ a) 1) 10 v _ water 10 (36) 3 3.49149(a1)  8.132810 (10 1)  log10 1013.246

a  373.16/(273.15 Dp) (37)

pv _ water  xwater ptotal (38) pv_water is the water vapor pressure, ptotal the total pressure of gas mixture, Dp the dew point, and xwater the mole fraction of water vapor in the gas mixture. In Goff-Gratch equation, the pressure unit is millibar and dew point unit is ºC.

2.2.2. Results

The vapor pressure of the IL-water system at 24ºC, 40ºC and 60ºC were measured. Data are shown in Figure 2-3.

Figure 2-3 Vapor pressure of IL-water at different concentrations and temperatures

It can be seen that departure from ideality is significant for all the IL concentration range.

The vapor pressure can be extremely low when IL molar concentration exceeds 40%. The figure indicates the water vapor pressure decreases dramatically as IL concentration increases. Temperature dependence of vapor pressure shows a great change in magnitude.

Increasing the temperature increases the vapor pressure at all concentrations.

2.3 Conductivity of Aqueous [C2mim][OAc]

2.3.1 Experiment

The conductivity measurement apparatus was model 19101-00 supplied by Cole Parmer.

The aqueous IL was kept in isothermal box to maintain the temperature.

2.3.2. Results

The conductivity of the system versus the weight percent of IL at 40ºC is shown in

Figure 2-4.

Figure 2-4 Conductivity of aqueous IL at 40ºC

It can be seen from Figure 2-4 that the measured electrical conductivity presents a maximum at IL weight percent of 35% at 40ºC. The curves presented in Figure 2-4 are similar to those published previously on aqueous solutions of the precursor salt including

EMIM-Br and EPYR-Br, where also a peak in the conductivity appeared [51]. A peak in the conductivity appears for the majority of concentrated electrolyte solutions. A satisfactory quantitative theory for explanation of the peak apparition and its position for aqueous ILs do not exist and the electrolyte solution transport theory is not clear. Some papers explain the relationship of conductivity vs. concentration qualitatively based in the presence of a glassy transition in the aqueous solution at concentration around the peak

[52]. Other researchers affirm that two different mechanisms are present in the electrical conduction. One of them is the number of ions present to transport charge which causes σ to increase with the IL concentration. The other one is related with the mobility of the ions in the solution which will decrease if the number of ions increases, and as a consequence, σ will decrease with IL concentration. At the peak the addition of both effects is optimal making conduction most effective [53].

2.3 Viscosity of Aqueous [C2mim][OAc]

2.3.1 Experiment

Viscosity is a key physical property of ILs. When used as a solvent, a low viscosity is generally desired to minimize pumping costs and increase mass transfer rates. Ionic liquids are much more viscous than typical organic solvents. As an example, BMIM·BF4 has a viscosity of 19.6 cP at 25ºC [54] versus 0.5 cP at 25ºC for methanol or 0.6 cP at

25ºC for toluene. High viscosity is probably inherent to ionic liquids due to strong

34 charge-charge interactions which lead to molecular clustering and interaction of larger transient aggregates during flow.

The viscosity of IL-water mixtures was measured with a Brookfield Viscometer, model

LVDV-III Ultra.

2.3.2. Results

Data obtained for temperatures of 25ºC, 30ºC, 35ºC, 40ºC, 45ºC and 50ºC over the full range of IL concentrations is presented in Figure 2-5.

Figure 2-5 Viscosity test of aqueous [C2mim][OAc] at different temperatures

From the figure we can see that viscosity of pure IL at room temperature is almost 130 times more viscous than water. High viscosity of the system arises from strong

+ - association of [C2mim] and [OAc] . IL viscosity is highly dependent on water concentration, especially at low concentration. The water addition in [C2mim][OAc] decrease the viscosity due to the ion pair dissociation by water which causes decrease of

[C2mim][OAc] cohesion. Viscosity at high IL concentration increases dramatically with decreasing temperature.

The viscosity is related to diffusion coefficient based on Stokes-Einstein equation:

kBT Dij  (39) 6j ri

Where Dij is the diffusion coefficient of solute i in pure solvent j, kB is Boltzmann constant, T is the temperature, ηj is the solvent viscosity, and ri is the radius of the solute molecule.

Based on Stokes-Einstein equation, diffusivity is inversely proportional to viscosity. High viscosity in aqueous ILs will lead to low diffusivity of the system. In the next section, the diffusivity of aqueous [C2mim][OAc] will be discussed.

2.4 Diffusivity of Aqueous [C2mim][OAc]

2.4.1 Experiment

Diffusivity or diffusion coefficient is a measure of the rate of molecular transport of one substance relative to another. This fundamental property is therefore involved in every transport equation either individual or collective. The rate of transport is characterized by

Fick's law:

dc J  D i (40) i ij dx

36 where Dij is the diffusion coefficient of solute i in pure solvent j, ci is concentration of i component, x is the direction of mass transport.

Self-diffusion coefficients provide uniquely detailed and easily interpreted information on molecular organization and phase structure [55]. Moreover, self-diffusion data can reveal structural change in the environment. For example, the experimental self-diffusion coefficient is directly related to molecular displacement in the laboratory coordinates for colloidal or macromolecular systems in solution which provides information about restricted molecular motion.

Among all the self-diffusion coefficient measurement experimental techniques, NMR pulsed field gradients (PFG) is probably one of the most convenient and accurate [56].

The technique has become a powerful tool to monitor molecular mobility in a variety of materials.

In the case of self-diffusivity measurement in the PFG experiment, the echo attenuation can be expressed by:

 A2τ  2 2 2  δ  Ln   γ Dδ G  Δ-  (41)  A0 2τ  3  where:

τ – Time interval for two pulses,

A – Echo attenuation with field gradient,

A0 – Echo attenuation without field gradient,

γ – Gyromagnetic ratio of the perturbed nuclei,

D – Self-diffusivity,

δ – Length of time the field gradient pulse is applied,

G – Pulse field gradient,

Δ – Time between start of first field gradient pulse and the start of the second field gradient pulse.

The experiments were typically performed with Δ fixed and δ regularly incremented while keeping G constant. If we plot δ against the echo attenuation A, then from the slope the effective diffusion coefficient can be extracted.

In the experiment, about 0.5 ml of a proper NMR solvent (Dimethyl Sulfoxide-D6) was added into a 5mm NMR tube and the test sample was injected into a narrow co-axial

NMR tube (4mm/2mm) which was then centered into the 5mm standard NMR tube afterwards to make the sample and NMR solvent vertical height to be approximately

50mm.

1H Pulsed Field Gradient (PFG) NMR diffusion studies was carried out using a Varian

NMR spectrometer operating at a proton resonance frequency of 600 MHz. Magnetic field gradients were generated using a PENTA probe.

DOSY (Diffusion Ordered Spectroscopy) experiment can be used to separate the NMR signals of mixture components based on different diffusion coefficients.

By applying two gradient pulses between 90 and 180 degree pulse (spin echo) and measuring the intensity of the echo, we can get DOSY spectrum in which x-axis is the chemical shifts for different protons and y-axis is the diffusion coefficients corresponding to the specific proton. If the sample is a mixture, the DOSY spectrum will show separated lines in coordinate with different species with their chemical shift on the x-axis.

2.4.3 Results

In this research, we used the pulse sequence DgcstelSL_CC (DOSY Gradient

Compensated Stimulated Echo with Spin Lock and Convection Compensation), which is an enhancement of the classical PGSE (Pulsed Gradient Spin-Echo) pulse sequence.

The parameter DAC_to_G must be calibrated first. This parameter is a conversion for the units of the gradient amplifier (DAC units) to gradient in gauss/cm. This was done with a sample of known diffusion coefficient. In our research, the diffusivity of HDO in D2O

(the residual solvent peak in D2O) at 25ºC was acquired. The temperature was regulated for at least 10 minutes to equilibrate, a diffusion experiment was run on the prepared sample. After Fourier transformation and baseline correction, the expected value of the diffusivity value (for HDO at 25ºC, the diffusivity is 19.02×10-10 m2/s) was input into the software to finish the calibration.

+ - After calibration, the diffusivity of [C2mim] , [OAc] and water in the system was measured and 2D-Dosy spectra of IL-water in deuterated DMSO were acquired.

Diffusion measurements of ions and water were performed using the standard PFG NMR stimulated echo pulse sequence in a broad range of temperatures between 297.15K and

323.15K for full IL concentration range.

Diffusion data were obtained from dependencies of the intensity of the PFG NMR signal

(A) on the amplitude of the magnetic field gradients (G). The signal intensity was determined by integrating the area under selected line(s) of the frequency-domain NMR spectra recorded by the PFG NMR stimulated echo pulse sequence. Different lines in such spectra can correspond to different species. Hence, diffusion data for a chosen type

39 of species in a sample can be obtained by selecting an appropriate line in the spectrum for data processing. For the NMR lines exhibiting no significant overlap with the lines of other types of ions or molecules in a sample the diffusivity (D) was determined from the measured attenuation of the PFG NMR signal (Ψ≡A(2τ)/A0(2τ)) corresponding to equation (35).

The signal attenuation was measured under the conditions when only the value of G is varied and all other parameters in the PFG MNR stimulated echo sequence remained constant. The effective diffusion time was changed by changing the value of Δ.

An example of 1H NMR spectra of 1-ethyl-3-methylimidazolium acetate ([C2mim][OAc]) containing water with IL mole concentration of 18.09% is shown in Figure 2-6. The area under each peak in the 1H NMR spectrum is proportional to the number of protons causing that peak. The concentration of IL was calculated based on the integrated area of the peaks for water and IL. The diffusivity spectra were recorded using the free induction decay NMR sequence. Due to high performance characteristics of our PFG NMR

+ spectrometer, the proton for position 5 in [C2mim] at 0.9ppm and anion proton lines at

1.3 ppm respectively can be completely resolved as shown in Figure 2-6. Water has a peak at the chemical shift of 4.8 ppm. The recorded PFG NMR attenuation curve for the cation and anion lines were monoexponential in agreement with Equation (35).

6 2

3 1

5 4 Water in aqueous ILs

6 Water in Anion DMSO

5 1 2 3 4

DMSO

Figure 2-6 Full range of 1H NMR spectrum of [C2mim][OAc] containing water

Figure 2-7 shows the 2-D DOSY NMR spectrum test of diffusivities for aqueous IL (IL mole%=18.09) at 20ºC.

Figure 2-7 2D DOSY Plots of aqueous [C2mim][OAc] with IL mole concentration =

18.09% at 20ºC

In the diffusivity spectrum, the x axis is the chemical shift for different protons in the

[C2mim][OAc] + water system and y axis is the corresponding diffusivity data for different components which has the unit of 10-10 m2/s. For aqueous IL with mole concentration of 18.09% at 20ºC, the cation and anion have the same diffusivity of 0.79

×10-10 m2/s and water has higher diffusivity of 1.84×10-10 m2/s.

Figure 2-8 gives the collected results of diffusivity vs IL concentrations for cation, anion and water in aqueous IL at 40ºC.

Figure 2-8 Diffusivity vs. IL concentrations for cation, anion and water in IL-water system at 40ºC obtained using DgcstelSL_CC NMR

From Figure 2-8 we can tell that diffusion coefficients for all three components decrease with IL concentration. Based on Stokes-Einstein equation (33), diffusivity is inversely proportional to viscosity. Viscosity of the system is increasing with IL concentration which results in decrease of diffusivities. As the water amount increases, the polar network of cation and anion is continuously broken up by the intruding water.

Correspondingly, the ions diffusivities increase continuously with water content, though they do so nonmonotonically with a sluggish increase for less than 40% (mole concentration) water content.

Diffusion coefficient of water molecule is about two times larger than the ions. Stoke’s law states that for a simple liquid the diffusion coefficient of a species is also inversely proportional to its effective hydrodynamic radius. Smaller molecule size of water will diffuse faster. However, the cation and anion have close self-diffusion coefficients despite their very different size and shape. This phenomenon arises from the fact that

+ - [C2mim] and [OAc] are strongly associated and therefore partially move together.

An addition of water reduces significantly the magnitude of the anomalous difference between the diffusivities of the cations and anions. This can be attributed to a partial screening of the electrostatic interaction between the cations and anions by water molecules. As a result of such screening the domain structure is expected to become more fluid and the cooperative character of the ions diffusion less pronounced. Addition of water also raises the ion pair dissociation rate and as a consequence to cause a larger increase of Danion compared with Dcation due to their different sizes.

+ - Figure 2-9 Temperature vs. diffusivity for both [C2mim] and [OAc] in pure 1-ethyl-3- methylimidazolium acetate obtained using DgcstelSL_CC NMR

Figure 2-9 shows the diffusion coefficients of the ions of pure 1-ethyl-3- methylimidazolium acetate as determined from DgcstelSL_CC NMR spectroscopy as a function of temperature. A cation diffusion coefficient of 1.6×10-11 m2/s and anion diffusion coefficient of 1.3×10-11 m2/s at 297.15 K have been obtained. These results are comparable to previous experimental measurements by Bowron et al for the same system

[57] which produce a cations diffusion coefficient of 1.5×10-11 m2/s and anion diffusion coefficient of 1.3×10-11 m2/s at 297.15 K.

Smaller ions will have a larger diffusivity for non-interacting species based on Stoke’s law. However, for interacting species particularly ionic system, the effective hydrodynamic radii of the ions, ri, have the potential to be increased due to association between cations and anions. Previous studies have found that cations still diffuse faster than anions, even when their corresponding effective hydrodynamics radii are calculated to be larger than those of the counterion [58]. The same observation has been made in this investigation as shown in Figure 2-9. This phenomenon implies that the anion diffuses as part of larger ion aggregates instead of single ions.

+ - Figure 2-10 Temperature vs. diffusivity for [C2mim] , [OAc] and water in aqueous 1- ethyl-3-methylimidazolium acetate (IL mole% = 65.98) obtained using DgcstelSL_CC

NMR

+ - Figure 2-11 Temperature vs. diffusivity for [C2mim] , [OAc] and water in aqueous 1- ethyl-3-methylimidazolium acetate (IL mole% = 6.79) obtained using DgcstelSL_CC

NMR

+ - Figure 2-10 and Figure 2-11 show the diffusion coefficients of [C2mim] , [OAc] and water of aqueous IL at IL mole concentration of 65.98% and 6.79% respectively as determined from NMR spectroscopy as a function of temperature. It is observed that at high IL concentration the larger cation diffuses faster than anion while at low IL concentration the smaller anion diffuses faster.

Rollet et al [59] hypothesis that water and ions do not occupy the same domains and water is not homogenously mixed with IL but forms small aggregates whose size and connections increase with water content. At low water content, water molecules are

generally hydrogen bonded to anions so as to form pore walls. While at high water content, the water would be released and bind together. Voth et al [60] state that a local water “pool” exists in several ILs in numerical simulation studies. Those researches are in good agreement with our results which imply that for low water concentrations, water molecule could break up the cation-anion polar network. The anion-water H bond interactions cause the water molecule to be attracted to the anion. Hence, anion diffuses slower because of the increase of its hydrodynamic radius. As the water content increases, more and more water molecules gather around anions so that the water aggregation reaches a limit at which the hydrogen bonding with the anions becomes saturated. After water concentration increases to a certain amount, the active water-water interactions dominate, and more water can be released. At high water concentration, water can combine with water to form clusters and also intrude into the cation-anion network which induces the anion diffusivity increment.

The predominate interaction of anion and water at low water content makes dehydration of aqueous IL at high concentrations difficult. The next chapter will discuss membrane methods to dry IL.

2.5 Conclusion

The physicp-chemical properties of ILs including vapor pressure, conductivity, viscosity and diffusivity are highly dependent on solution temperature and water concentrations.

Water molecules preferentially interact with anions through hydrogen bonding at low water concentrations. The hydrogen bonding of water and anion results in smaller diffusivity of anion compared with larger cation and extremely low vapor pressure for

aqueous IL with small water content. As a consequence, the removal of water with low water concentration against the hydrogen bond of water-anion becomes more difficult.

The vapor pressure of aqueous IL can be directly used to calculate the mass transfer resistance for membrane separation system. Self-diffusivity of aqueous IL can help to estimate the boundary layer thickness if mass transfer in liquid dominates the water transport.

Chapter 3

Membrane Drying of Aqueous 1-ethyl-3-methylimidazoulium Acetate ([C2mim][OAc])

3.1 Instruction

The traditional method to remove water or organic solvents is to heat the mixture under vacuum. Particularly, the reduction of water commonly needs long period of time under low pressure at high temperatures. This method needs long time operation to remove water at low ILs concentrations and high requirement for the vacuum especially for high

ILs concentrations due to the extremely low vapor pressure. The long-term thermal stability of ILs has been investigated and it is found that degradation occurs when the temperature is held at an elevated level for a longer time (about 10 hours at T > 150°C)

[61]. After the drying procedure, the ILs which are referred to as “dried” ILs may still contain water [62]. Both hydrophilic and hydrophobic ILs can absorb water from the atmosphere. This method could be difficult to decrease the water content to a low level.

The energy cost of heat and vacuum for evaporation needs to be concerned.

Membrane separation offers an alternative method to remove water from aqueous ILs.

The potential advantages of membrane operation have been successfully demonstrated at the lab scale and pilot scale, including improved product quality, lower energy consumption, and easy scale-up. This chapter will discuss two different membrane

operations for drying aqueous ILs: straight RO separation and pervaporation. The performance of both will be illustrated and compared.

3.2 Straight RO Membrane Separation

3.2.1 Experiment

In reverse osmosis, the feed stream containing a solute is pressurized on one side of a semipermeable membrane. Pressure greater than the osmotic pressure of the feed is exerted to force solvent through the membrane. Figure 3-1 provides a schematic representation of the process.

Figure 3-1 Schematic of the reverse osmosis process

Reverse osmosis is used widely to concentrate aqueous solutions containing a low molecular weight solute. Applying pressure to the feed solution forces smaller components of the feed through the membrane while larger components are rejected. The

most common use for reverse osmosis is in the desalination of water. Such a process can also be used to remove water from an IL-water mixture.

Figure 3-2 illustrates the experimental setup used to evaluate the potential of reverse osmosis to dehydrate IL-water mixtures.

Figure 3-2 Water flux measurement from straight RO membrane

The membrane cell system used in this research is GE SEPA CF II system. The membrane cell accommodate any 7.5 in × 5.5 in (19 cm × 14 cm) flat sheet membrane for a full 22 in2 (140 cm2) of effective membrane area. The heights of the feed and permeate

slot in the Sepa CF are 0.172 cm and 0.035 cm respectively. The membrane thickness is

0.016 cm. The diamond nonwoven spacer in the feed channel has thickness of 0.864 mm and the one in the permeate channel has thickness of 0.2 mm.

The feed spacer was placed into the slot of the bottom cell body half below the membrane and the permeate spacer with 0.2 mm thickness was placed on the top of the membrane which fits over the guideposts. The cell body was inserted into the cell holder. Hydraulic pressure was applied to the top of the holder. The same membrane system was also used for pervaporation procedure of aqueous IL.

As shown in Figure 3-2, the IL-water mixture was pumped into feed side of the membrane cell. The pressure on the feed side was controlled by adjusting the valve at the retentate outlet. The retentate and permeate products were recycled to the feed reservoir to allow continuous operation at constant feed flow rate and pressure. HPLC and NMR were used to measure the concentration of IL in the products streams.

3.2.2 Results and Discussion

Reverse osmosis performance was evaluated at room temperature for an IL feed concentration of 5.0%, feed flow rate of 10 ml/min and membrane (polyamide RO) area of 0.014 m2. Tables 3.1, 3.2 and 3.3 report the concentration and flow rate of the retentate and permeate streams produced at feed pressures ranging from 350 to 450 psi. Such a pressure range is similar to that encountered in reverse osmosis processes used for seawater desalination.

Table 3.1 IL concentration results of RO separation at 350 psi

Concentrated outlet Permeate outlet IL weight percent IL flow rate IL weight percent IL flow rate (%) (ml/min) (%) (ml/min) 5.44 8.5 0.30 1.5

Table 3.2 IL concentration results of RO separation at 400 psi

Concentrated outlet Permeate outlet IL weight percent IL flow rate IL weight percent IL flow rate (%) (ml/min) (%) (ml/min) 5.57 8.4 0.33 1.6

Table 3.3 IL concentration results of RO separation at 450 psi

Concentrated outlet Permeate outlet IL weight percent IL flow rate IL weight percent IL flow rate (%) (ml/min) (%) (ml/min) 6.47 8.2 0.33 1.9

From the experimental data, we can see that IL purification using RO will require high feed pressures to overcome the osmotic pressure of the IL-water mixture even at low IL concentrations. Separation of higher concentration of IL will need even higher pressure.

Additionally, the concentration change from feed inlet to retentate outlet is small and the permeate possesses an undesirably high concentration of IL. Therefore, RO does not appear to offer promise for IL recovery, due to energy requirements and IL losses, and pervaporation was considered as an alternative.

3.3 Pervaporation separation

3.3.1 Introduction and Theory

Pervaporation is an emerging membrane technique designed to separate a liquid mixture by partly vaporizing it through a nonporous permselective membrane. The feed is maintained to contact one side of the membrane and a fraction of it passes through the membrane to the opposite side leaving in the vapor state, which is kept under vacuum or is purged with a stream of inert carrier gas. In this work, we used nitrogen or dry air as sweep in the permeate side.

3.3.2 Experimental Setup

The experiment set up for pervaporation of aqueous IL is illustrated as Figure 3-3:

Figure 3-3 Pervaporation for aqueous IL

In the experiment, the compressed air was introduced to an oil removal filter to remove oil particulate that is typical of many compressed air systems. The cleaned air went into a hollow fiber dehydration module to decrease the water concentration in the air. The water concentration in the sweep of the pervaporation system was adjusted by a valve after the dehydration module. The dried air swept the permeate side of the membrane to take water vapor out of the system. The aqueous IL was pumped into the feed side of the membrane, exhibiting counter-current flow with the air sweep. The membrane cell was held in a GE

Sepa CF II system which was wrapped with a heating tape and the whole system was maintained in an isothermal box. The mass of aqueous IL was measured with a weight balance. The water flux is calculated from the weight loss during a period of time. The concentration of aqueous IL was measured with 1H NMR. By changing the heating temperature of thermal tape, water flux at different temperatures was measured.

3.3.3 Results

3.3.3.1 Membrane Types

A number of membranes of varying composition and molecular weight cut-off were evaluated for pervaporation. The first membranes tested were polysulfone ultra-filtration membranes with MW cut-offs of 60,000 and 30,000. Both membranes leaked IL from the feed to the permeate. This was due to the pore size being sufficiently large that capillary forces were not able to hold the liquid within the membrane pores as illustrated in Figure

3-4.

Figure 3-4 Schematic diagram of IL leaking using polysulfone ultra-filtration membrane

Subsequently, a composite polyamide ultrafiltration membrane with MW cut-off of 1,000 was tested. Experiments were performed at temperature of 40ºC, feed flow rate of 30 ml/min, and a nitrogen sweep at 1.5 L/min. Table 4 reports the measured water flux at different IL concentrations:

Table 3.4 Water flux at different IL concentrations for ultralfiltration membrane with

MWCO of 1,000

IL wt% water flux (g/hr/m2)

75.0 14.1

77.5 12.5

81.3 10.0

81.9 non-measurable

The data indicate that the water flux decreases as the IL concentration increases.

Moreover, the rate of water transport is not measurable at a concentration of 81.9%. This phenomenon appears to arise from the formation of a concentration boundary layer

within the liquid contacting the membrane and penetrating the membrane pores as illustrated in Figure 3-5.

Figure 3-5 Schematic diagram of boundary layer increment due to IL penetration

Experiment with same operation conditions was conducted using a membrane with a polyamide RO membrane with a lower MW cut-off of 200. Table 3-5 shows water flux results at different IL concentrations:

Table 3.5 Water flux at different IL concentrations

IL wt% water flux (g/hr/m2)

80.6 15.1

85.6 9.2

89.4 4.0

92.7 1.2

94.5 0.3

97.5 0.1

The water flux is significantly higher for this membrane and the concentration at which the water flux approaches to zero also is significantly higher, > 97%. This suggests the boundary layer thickness is much smaller than for the composite polyamide ultrafiltration membrane.

3.3.3.2 Performance of Aqueous IL Pervaporation with RO AK Membrane

As discussed above, RO AK membrane has the best performance for pervaporation. We measured the water flux with different IL-water concentrations at different flow rates on both liquid and gas sides. Figure 3-6 shows the water flux results for experiments conducted at room temperature using a nitrogen sweep.

Figure 3-6 Water flux vs. IL concentration at different flow rate in feed and permeate channel at room temperature

Figure 3-6 indicates that the water flux increases dramatically if both gas and liquid flows are doubled. The origin of this increase can be rationalized by using the mass transfer

theory in pervaporation separation. The contribution of the mass transfer resistance in gas phase, liquid phase and membrane needs to be analyzed individually to understand the transport mechanism of pervaporation dehydration of aqueous IL.

The physical and chemical properties of aqueous IL are highly dependent on temperature.

The driving force of pervaporation is the water partial pressure difference in the bulk feed and bulk permeate. Water partial pressure of aqueous IL is highly sensitive to temperature based on the data in the previous chapter. Temperature increment should have significant improvement for the water flux of aqueous IL pervaporation dehydration.

Figure 3-7 compares the water flux vs. IL concentration at room temperature with 40ºC.

Figure 3-7 Water flux vs. IL concentration at different operation temperatures

Comparing water flux through the membrane at different temperature, it is very clear that temperature plays an important role on water transport rate. This phenomenon is due to the water vapor pressure increment with temperature. In next chapter, the mass transfer resistances in the liquid side, gas side as well as the membrane will be analyzed separately dependent on different temperatures.

3.4 Conclusion

Both straight RO and pervaporation can be used to remove water from aqueous IL. It needs high energy cost even for low IL concentration for straight RO process. Moreover, the concentration rate is low and IL can be lost into the permeate side. RO AK membrane has the best performance for pervaporation of aqueous IL in this research. Pervaporation is more promising compared with straight RO process as the IL can be concentrated to more than 95 wt% to satisfy the requirement for the biomass pretreatment. The water flux is increasing with temperature.

Chapter 4

Mass Transfer in Pervaporation for Aqueous Ionic Liquid

4.1 Mass Transfer Resistance for Pervaporation

Mass transport by pervaporation across the membrane can be divided into three successive steps [63]:

1. Upstream partitioning of the feed components between the flowing liquid mixture

and the swollen upstream surface layer of the membrane.

2. Diffusion of the components in the surface layer through the membrane.

3. Desorption of these components at the downstream surface of the membrane.

The overall mass transfer coefficient can be determined from the bulk feed concentration and the steady state permeate flux of investigated component. As discussed in Equation

(30) in Chapter 1, the reciprocal value of the overall mass transfer coefficient is equal to the sum of the boundary layer resistance (gas phase and liquid phase) and the membrane resistance. The boundary layer resistance for a given solution is mainly determined by the flow conditions and module design. With increasing flow velocity (increasing Reynolds number) the boundary layer resistance decreases and the membrane resistance becomes more dominant. The membrane resistance is directly proportional to the effective membrane thickness and inversely proportional to the component permeability.

The overall mass transfer coefficient is an important parameter to evaluate the mass transport efficiency. Based on equation (29) and the water vapor pressure test data of water-IL system as a function of temperature, we can determine the overall mass transfer resistance. To separate the contributions of the liquid side, gas side, and membrane mass transfer resistance, the water flux through the membrane with pure water as feed as well as aqueous IL needs to be measured first.

4.2 Mass Transfer Coefficient in Gas Phase (Pervaporation of Pure Water)

4.2.1 Theory

Based on equation (29), the overall mass transfer coefficient is the sum of resistance in the liquid phase, gas phase and membrane. The liquid phase mass transfer resistance can be neglected if using pure water as feed.

When dry gas flows continuously along the membrane surface, the water vapor concentration difference between the feed and permeate side varies throughout the length of the membrane. To account for this condition, the corrected mean water vapor pressure difference is the logarithmic-mean value:

in out out in ( p f  p p )  ( p f  p p ) p  plogmean  in out (42) p f  p p ln( out in ) p f  p p where Δplog mean is the log mean water vapor pressure difference between the feed and

in out in permeate, pf and pf are the inlet and outlet water vapor pressure of the feed while pp

out and pp are the inlet and outlet water vapor pressure of the permeate. The partial pressure of water vapor of inlet could be adjusted through changing the sweep dew points.

With pure water, the Δplog mean can be expressed as:

in out p p  p p plogmean  in out (43) p f  p p ln( out in ) p f  p p

in out As feed is pure water, pf and pf in Equation (43) is equal to saturated vapor pressure of

in water. pp can be calculated based on measured dew point using the Goff-Gratch water dew point correlations discussed in Chapter 1 to calculate the vapor pressure in the gas

out sweep. pp is calculated based on the water flux through the membrane and the initial inlet gas condition.

Also, the liquid side mass transfer resistance is zero. Therefore, the overall resistance given by Equations (30) and (31)

l RT Rov  Rm  RpB   n (44) DH '' QG

The parameters of β and n can be determined experimentally by varying the gas flow rates and regression of Rov vs. QG. The mass transfer coefficients of membrane and gas side can be analyzed based on the calculated parameters. Liquid side mass transfer coefficient can be evaluated by using IL-water mixtures as feed with varied liquid side flow rates.

4.2.2 Results

4.2.2.1 Mass Transfer Coefficient in Gas Phase

Mass transfer coefficients in the gas phase are evaluated first. When the feed is filled with pure water, the water flux through pervaporation membrane can be expressed as follows based on equation (43) and (44):

in out l RT p p  p p J (  m )  p  in out (45) DH '' QG p f  p p ln( out in ) p f  p p The water concentration of inlet gas can was calculated from the measured dew point using the Goff-Gratch correlation. The dew point of the gas was measured using a

Vaisala DMT242 Dew point Transmitter. Experiments were conducted using a pure water feed and a sweep with variable flow rate and inlet dew point (i.e., water concentration), operating at room temperature. The measured water fluxes are presented in Figure 4-1.

Figure 4-1 Water flux vs. gas flow rates with different inlet dew points

Mass transfer resistances calculated using Equation (45) for a dew point of -10ºC, -15 ºC ,

-20 ºC and -60 ºC are reported in Table 4.1, 4.2, 4.3 and 4.4.

Table 4.1 Calculation of Rov at different Qgas with inlet DP = -10ºC at 24ºC

in in in in out out Qgas pp cp np J Qwater Xwater pp Δp Rov

mol/s Pa mol/m3 mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

1.33E-03 286.04 1.16E-03 2.27E-06 2.25E-03 1.89E-03 2.31% 2340 1364 6.06E+05

2.69E-03 286.04 1.16E-03 4.57E-06 2.87E-03 3.85E-03 1.47% 1494 1973 6.87E+05

4.55E-03 286.04 1.16E-03 7.73E-06 4.58E-03 5.00E-03 1.39% 1411 2025 4.42E+05

8.68E-03 286.04 1.16E-03 1.48E-05 5.95E-03 7.09E-03 0.95% 964 2283 3.84E+05

68 1.30E-02 286.04 1.16E-03 2.21E-05 8.45E-03 5.65E-03 0.90% 914 2311 2.74E+05

1.04E-02 286.04 1.16E-03 1.76E-05 6.72E-03 1.08E-02 0.90% 913 2311 3.44E+05

2.13E-02 286.04 1.16E-03 3.62E-05 1.18E-02 1.17E-02 0.77% 780 2383 2.02E+05

2.48E-02 286.04 1.16E-03 4.22E-05 1.28E-02 1.59E-02 0.72% 728 2411 1.88E+05

2.67E-02 286.04 1.16E-03 4.54E-05 1.39E-02 1.87E-02 0.72% 732 2409 1.74E+05

Table 4.2 Calculation of Rov at different Qgas with inlet DP = -15ºC at 24ºC

in in in in out out Qgas pp cp np J Qwater Xwater pp Δp Rov

mol/s Pa mol/m3 mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

6.80E-04 191.02 7.73E-02 7.73E-05 1.27E-03 1.07E-03 2.55% 2586 1147 9.01E+05

1.36E-03 191.02 7.73E-02 1.55E-04 2.16E-03 1.82E-03 2.18% 2205 1509 6.98E+05

2.04E-03 191.02 7.73E-02 2.32E-04 2.98E-03 2.51E-03 2.00% 2031 1646 5.52E+05

2.72E-03 191.02 7.73E-02 3.09E-04 3.88E-03 3.26E-03 1.96% 1984 1681 4.33E+05

69 3.40E-03 191.02 7.73E-02 3.87E-04 4.69E-03 3.94E-03 1.89% 1918 1729 3.69E+05

7.08E-03 191.02 7.73E-02 8.05E-04 6.83E-03 5.73E-03 1.33% 1349 2102 3.08E+05

1.54E-02 191.02 7.73E-02 1.75E-03 1.09E-02 9.13E-03 0.98% 995 2309 2.12E+05

2.10E-02 191.02 7.73E-02 2.39E-03 1.31E-02 1.10E-02 0.86% 875 2376 1.82E+05

3.15E-02 191.02 7.73E-02 3.58E-03 1.70E-02 1.42E-02 0.75% 757 2440 1.44E+05

3.54E-02 191.02 7.73E-02 4.02E-03 1.82E-02 1.53E-02 0.72% 725 2457 1.35E+05

Table 4.3 Calculation of Rov at different Qgas with inlet DP = -20ºC at 24ºC

in in in in out out Qgas pp cp np J Qwater Xwater pp Δp Rov

mol/s Pa mol/m3 mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

6.80E-04 125.29 5.07E-02 5.07E-05 1.29E-03 1.09E-03 2.72% 2752 943 7.28E+05

1.36E-03 125.29 5.07E-02 1.01E-04 2.22E-03 1.86E-03 2.35% 2385 1373 6.18E+05

2.04E-03 125.29 5.07E-02 1.52E-04 3.11E-03 2.61E-03 2.21% 2241 1502 4.83E+05

2.72E-03 125.29 5.07E-02 2.03E-04 4.16E-03 3.49E-03 2.22% 2245 1498 3.60E+05

70 3.40E-03 125.29 5.07E-02 2.54E-04 5.12E-03 4.30E-03 2.18% 2214 1524 2.98E+05

8.16E-03 125.29 5.07E-02 6.09E-04 7.95E-03 6.68E-03 1.47% 1487 2044 2.57E+05

9.19E-03 125.29 5.07E-02 6.85E-04 8.44E-03 7.09E-03 1.39% 1412 2091 2.48E+05

1.02E-02 125.29 5.07E-02 7.61E-04 9.06E-03 7.61E-03 1.35% 1368 2118 2.34E+05

1.18E-02 125.29 5.07E-02 8.77E-04 9.72E-03 8.17E-03 1.27% 1284 2169 2.23E+05

1.36E-02 125.29 5.07E-02 1.01E-03 1.08E-02 9.04E-03 1.22% 1234 2199 2.04E+05

1.78E-02 125.29 5.07E-02 1.33E-03 1.28E-02 1.08E-02 1.12% 1136 2257 1.76E+05

Table 4.4 Calculation of Rov at different Qgas with inlet DP = -60ºC at 24ºC

in in in in out out Qgas pp cp np J Qwater Xwater pp Δp Rov

mol/s Pa mol/m3 mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

6.62E-04 1.90 7.67E-04 7.67E-07 1.38E-03 1.16E-03 2.84% 2881 688 4.97E+05

1.32E-03 1.90 7.67E-04 1.53E-06 2.61E-03 2.19E-03 2.68% 2718 1026 3.94E+05

1.99E-03 1.90 7.67E-04 2.30E-06 3.71E-03 3.12E-03 2.55% 2584 1202 3.24E+05

2.65E-03 1.90 7.67E-04 3.07E-06 4.97E-03 4.17E-03 2.56% 2595 1189 2.39E+05

71 3.31E-03 1.90 7.67E-04 3.84E-06 6.08E-03 5.11E-03 2.51% 2544 1248 2.05E+05

5.03E-03 1.90 7.67E-04 5.83E-06 7.22E-03 6.07E-03 1.97% 1998 1738 2.41E+05

7.42E-03 1.90 7.67E-04 8.59E-06 8.84E-03 7.42E-03 1.64% 1664 1977 2.24E+05

Take the data in Table 4.1 as an example. The gas flow rate used for these calculations was measured using an Omega volumetric flow meter sensor. The molar flow rate of inlet

in air (Qgas ) was converted by multiplying volumetric flow rate with the air density and

in then dividing with the air molecular weight. The inlet gas partial pressure (pp ) was derived from the Goff-Gratch correlations (Equation (36), (37), (38)) with the measured

in dew point and the molar concentration of water in the inlet gas (cp ) calculated using the

in ideal gas law. The molar flow rate of water np was measured by time volume collection.

The water flux (J) was measured by the change of mass of feed reservoir converted to a

out molar water flux (Qwater). The water molar fraction of the outlet (Xwater ) was calculated from the molar water flux and molar flow rate of the inlet gas. The water partial pressure

out of the outlet pp was calculated as the product of the water mole fraction and atmosphere pressure. The pressure difference (Δp) was calculated according to the inlet and outlet water vapor pressure of both permeate and liquid side based on Equation (45). Hence, the overall resistance (Rov) can be evaluated from the log mean pressure difference and water flux.

The membrane permeability is the product of diffusivity and solubility. One expects the diffusivity of water to increase with water content in the membrane. Thus, the water permeability also should increase with water concentration. Additionally, if the membrane mass transfer resistance is dominant, the overall mass transfer coefficient should increase with water concentration. However, the experimental measurements indicate the overall mass transfer coefficient 1/Rov decreases as the water concentration increases. Hence, the membrane mass transfer resistance (l/DH'') must be smaller than

n the gas phase resistance (RT/βQgas ) and Equation (45) can be approximately expressed as:

p RT 1  n  n (46) J QG QG

By taking the natural log on both sides of Equation (46) and performing of linear regression of ln(Δp/J) versus QG, the best fit values for β and n can be determined. The calculated parameters for different inlet dew points are listed in Table 4.5.

Table 4.5 Parameters of mass transfer coefficient in gas phase with different inlet dew points at 24ºC

DP n α β

ºC (mol/s)1-n·m-2n/pa (mol/s)1-n·m3-2n/mol

-10 0.462 2.75E-05 6.80E-02

-15 0.474 3.57E-05 8.82E-02

-20 0.429 3.17E-05 7.83E-02

-60 0.366 3.14E-05 7.75E-02

The results in Table 7 can be used to evaluate mass transfer coefficients at different flow rates within the dew point range considered.

4.2.2.2 Effect of Temperature on Mass Transfer Coefficient in Gas Phase

As mass transfer coefficient is a function of the viscosity and diffusivity of the bulk system which are functions of temperature. Viscosity of gas will increase with temperature. Hence, the increase of temperature might have a negative impact on the mass transfer coefficient in the gas phase due to higher viscosity. However, the diffusivity of gas will increase with temperature which could result in decrease of the mass transfer resistance.

Figure 4-2 shows the water flux test for pure water using nitrogen as sweep, the water and the cell are maintained at 40ºC.

Figure 4-2 Water flux vs. gas flow rates with N2 sweep at 40ºC

Compare the water flux at 40°C (Figure 4-2) with 24°C (Figure 4-1), we can find there is a large increment of water flux dependent on temperature. As water flux is a function of water pressure difference through the membrane and the mass transfer coefficient, this elevation could be from the contribution of the water vapor pressure increment due to higher temperature. The contribution of mass transfer coefficient needs to be analyzed further.

Table 4.6 Calculation of Rov at different Qgas with inlet dew point of -60ºC at 40ºC

in in in in out out Qgas pp cp np J Qwater Xwater pp Δp Rov

mol/s Pa mol/m3 mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

1.48E-02 1.90 7.67E-04 1.66E-05 1.24E-02 1.04E-02 1.16% 1180 2302 1.90E+05

1.88E-02 1.90 7.67E-04 2.12E-05 1.38E-02 1.16E-02 1.02% 1036 2380 1.75E+05

2.29E-02 1.90 7.67E-04 2.59E-05 1.50E-02 1.26E-02 0.91% 924 2435 1.62E+05

2.42E-02 1.90 7.67E-04 2.72E-05 1.53E-02 1.28E-02 0.88% 891 2450 1.59E+05

75 2.74E-02 1.90 7.67E-04 3.08E-05 1.71E-02 1.43E-02 0.87% 879 2475 1.49E+05

The mass transfer resistance of gas phase in at 40ºC is calculated following the process mentioned in previous section. The calculated data are shown in Table 4.6.

As discussed in previous section, regression of Rov vs. Qgas gives parameters in Table 4-7.

Table 4.7 Parameters of mass transfer coefficient in gas phase with nitrogen sweep at

40ºC

DP n α Β

ºC (mol/s)1-n·m-2n/pa (mol/s)1-n·m3-2n/mol

-60 0.382 2.65E-05 6.54E-02

Comparing the parameters for mass transfer coefficient at 40ºC with 24ºC, we can find there is small difference between them. This is reasonable as the diffusivity and viscosity of gas is not highly sensitive to temperature. The difference of the water flux with temperature is due to the water vapor pressure increase with temperature.

4.3 Calculation for Mass Transfer Resistance in Liquid Phase and Membrane

(Pervaporation of Aqueous IL)

4.3.1 Theory

The impact of mass transfer resistance in gas phase is discussed in previous section. In order to evaluate the contribution of mass transfer resistance in the liquid phase and membrane, pervaporation of aqueous IL needs to be conducted.

From Equation (29), the overall mass transfer resistance can be calculated based on the water flux experimental data and the vapor pressure difference in bulk feed and bulk

permeate. The bulk feed water vapor pressure can be predicted from the water vapor pressure test for aqueous IL in Chapter 2.

Based on equation (30) and (31), the overall mass transfer resistance can be expressed as a function of aqueous IL flow rate as Equation (47):

1 1 1 Rov  R fB  Rm  R pB  m   (47) ' H 'QIL ' H ''(D / l) k pB / RT

As RpB can be derived from the equation and parameters discussed previously, the remaining RfB and Rm can be calculated by deduct RpB from Rov. If do regression of (RfB+

Rm) vs. QIL, the parameters related to RfB and Rm can be analyzed.

4.3.2 Results

The pervaporation dehydration of aqueous IL was operated using the set up as Figure 3-3.

The system was isolated in a box to maintain the temperature. A certain concentration of

IL is pumped into the system and water flux was calculated by measuring the weight loss during a period of time. Experiments were performed for a broud range of gas and liquid flow rate. Figure 4-3 illustrates the measured water flux with a nitrogen sweep (DP = -

60ºC) for a feed IL mole concentration of 14.0% at room temperature.

Figure 4-3 Jwater vs. QIL with different nitrogen flow rates (DP=-60ºC) at room temperature for 14.0 mole% of IL.

The mass transfer coefficients in the liquid side and membrane for 14.0 mole% aqueous

IL are then calculated based on the water flux data shown in Figure 4-3. The calculation process for overall mass transfer resistance with different gas flow rates for fixed liquid flow rate of 60 ml/min, 90 ml/min, 120 ml/min, 150 ml/min and 180 ml/min are shown in

Table 4.8, 4.9, 4.10, 4.11 and 4.12 respectively.

Table 4.8 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 60 ml/min

in in in in out out QN2 QN2 pp np J Qwater Xwater pp Δp Rov

L/min mol/s Pa mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

0.5 3.43E-04 1.90 3.85E-07 1.26E-04 1.05E-04 0.51% 516 836 6.65E+06

1 6.86E-04 1.90 7.69E-07 2.26E-04 1.89E-04 0.46% 465 867 3.83E+06

2 1.37E-03 1.90 1.54E-06 4.08E-04 3.41E-04 0.41% 418 894 2.19E+06

3 2.06E-03 1.90 2.31E-06 6.14E-04 5.14E-04 0.41% 420 893 1.45E+06

79 4 2.74E-03 1.90 3.08E-06 7.27E-04 6.07E-04 0.37% 373 921 1.27E+06

5 3.43E-03 1.90 3.85E-06 50.266 6.47E-04 0.31% 318 952 1.23E+06

Table 4.9 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 90 ml/min

in in in in out out QN2 QN2 pp np J Qwater Xwater pp Δp Rov

L/min mol/s Pa mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

0.5 3.43E-04 1.90 3.85E-07 1.28E-04 1.07E-04 0.52% 526 829 6.47E+06

1 6.86E-04 1.90 7.69E-07 2.30E-04 1.92E-04 0.47% 472 862 3.75E+06

2 1.37E-03 1.90 1.54E-06 4.15E-04 3.47E-04 0.42% 425 890 2.15E+06

3 2.06E-03 1.90 2.31E-06 6.21E-04 5.19E-04 0.42% 425 891 1.43E+06

80 4 2.74E-03 1.90 3.08E-06 7.33E-04 6.12E-04 0.37% 376 919 1.25E+06

5 3.43E-03 1.90 3.85E-06 7.85E-04 6.56E-04 0.32% 322 950 1.21E+06

Table 4.10 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 120 ml/min

in in in in out out QN2 QN2 pp np J Qwater Xwater pp Δp Rov

L/min mol/s Pa mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

0.5 3.43E-04 1.90 3.85E-07 1.30E-04 1.08E-04 0.52% 532 826 6.37E+06

1 6.86E-04 1.90 7.69E-07 2.32E-04 1.94E-04 0.47% 476 860 3.70E+06

2 1.37E-03 1.90 1.54E-06 4.21E-04 3.52E-04 0.43% 432 886 2.10E+06

3 2.06E-03 1.90 2.31E-06 6.28E-04 5.25E-04 0.42% 429 888 1.41E+06

81 4 2.74E-03 1.90 3.08E-06 7.43E-04 6.21E-04 0.38% 381 916 1.23E+06

5 3.43E-03 1.90 3.85E-06 7.98E-04 6.66E-04 0.32% 327 947 1.19E+06

Table 4.11 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 150 ml/min

in in in in out out QN2 QN2 pp np J Qwater Xwater pp Δp Rov

L/min mol/s Pa mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

0.5 3.43E-04 1.90 3.85E-07 1.31E-04 1.10E-04 0.53% 538 822 6.26E+06

1 6.86E-04 1.90 7.69E-07 2.34E-04 1.96E-04 0.47% 480 858 3.67E+06

2 1.37E-03 1.90 1.54E-06 4.27E-04 3.57E-04 0.43% 438 883 2.07E+06

3 2.06E-03 1.90 2.31E-06 6.33E-04 5.29E-04 0.43% 433 886 1.40E+06

82 4 2.74E-03 1.90 3.08E-06 7.45E-04 6.23E-04 0.38% 382 916 1.23E+06

5 3.43E-03 1.90 3.85E-06 8.05E-04 6.73E-04 0.33% 330 945 1.17E+06

Table 4.12 Calculation of Rov at different QN2 with inlet dew point of -60ºC for at 24ºC for 14.0 mole% of IL with QIL = 180 ml/min

in in in in out out QN2 QN2 pp np J Qwater Xwater pp Δp Rov

L/min mol/s Pa mol/min mol/(s·m2) mol/min Pa Pa Pa·s·m2/mol

0.5 3.43E-04 1.90 3.85E-07 1.33E-04 1.12E-04 0.54% 546 816 6.13E+06

1 6.86E-04 1.90 7.69E-07 2.38E-04 1.99E-04 0.48% 488 853 3.59E+06

2 1.37E-03 1.90 1.54E-06 4.35E-04 3.64E-04 0.44% 446 878 2.02E+06

3 2.06E-03 1.90 2.31E-06 6.33E-04 5.29E-04 0.43% 433 886 1.40E+06

83 4 2.74E-03 1.90 3.08E-06 7.57E-04 6.33E-04 0.38% 388 912 1.21E+06

5 3.43E-03 1.90 3.85E-06 8.14E-04 6.80E-04 0.33% 334 943 1.16E+06

Take the calculation of mass transfer resistance in Table 4.8 as an example. The gas flow rate used for these calculations was measured using an Omega volumetric flow meter

in sensor. The molar flow rate of inlet air (QN2 ) was converted by multiplying volumetric flow rate with the nitrogen density and then dividing with the nitrogen molecular weight.

in The inlet gas partial pressure (pp ) was derived from the Goff-Gratch correlations

(Equation (36), (37), (38)) with the measured dew point. The molar concentration of water in the inlet gas was calculated using the ideal gas law and the molar flow rate of

in water np was measured by time volume collection. The water flux (J) was measured by the change of mass of feed reservoir converted to a molar water flux (Qwater). The water

out molar fraction of the outlet (Xwater ) was calculated from the molar water flux and molar

out flow rate of the inlet gas. The water partial pressure of the outlet pp was calculated as the product of the water mole fraction and atmosphere pressure. The pressure difference

(Δp) was calculated according to the inlet and outlet water vapor pressure of both permeate and liquid side based on Equation (45). Moreover, the water vapor pressure of aqueous IL with certain concentration can be derived from the measurement data in

Chapter 2. Hence, the overall resistance (Rov) can be evaluated from the log mean pressure difference and water flux.

As the parameters for gas side mass transfer coefficient are derived in section 4.2.2, the remaining resistances in membrane and liquid side can be calculated by deducting RpB from Rov. The equation (47) can be expressed as:

1 1 Rov  R pB  R fB  Rm  m  (48) ' H'QIL ' H''(D / l)

H' in Equation (48) stands for the Henry’s law constant which is the ratio at equilibrium of concentration of water to the partial pressure of water for aqueous IL system (data in

Figure 2-3). The calculated Henry’ law constant is shown in Figure 4-4:

Figure 4-4 Henry’s law constant for aqueous IL vs. IL mole concentrations at 24°C and

40°C

Do regression of (RfB + Rm) with QIL for certain QN2. The parameters in Equation (48) can be derived using Polymath. Table 4-13 shows the calculated results at room temperature:

Table 4.13 Parameters for mass transfer resistance in membrane and liquid side at 24ºC

in QN2 (L/min) parameters 0.5 1 2 3 4 5 2 Rm(Pa·s·m /mol) 4.4E+06 2.7E+06 1.3E+06 9.4E+05 8.1E+05 7.5E+05 α’((min/ml)0.33m/s) 2.1E-07 4.8E-07 6.2E-07 1.6E-06 1.9E-06 1.5E-06 m 0.33 0.33 0.33 0.33 0.33 0.33

As can be seen both Rm and RfB are decreasing with gas flow rates. The mass transfer resistance in the membrane and liquid side for whole range of IL concentrations can be analyzed following such process.

The dependence of mass transfer resistances of the overall, liquid side, membrane and gas side on IL concentration with QN2 of 5 L/min and QIL of 180 ml/min at 24ºC and

40ºC are shown in Figure 4-5 and 4-6.

Figure 4-5 Mass transfer resistance vs. IL mole concentration at 24ºC with QN2 = 5

L/min, and QIL = 180 ml/min

Figure 4-6 Mass transfer resistance vs. IL mole concentration at 40ºC with QN2 = 5

L/min, and QIL = 180 ml/min

It can be seen that the main mass transfer resistance is in the membrane. Rov, Rm and RfB are increasing with the IL concentrations. Comparing the data at 24ºC with 40ºC, the overall mass transfer resistance is decreasing with temperature which is mainly due to the decrement in Rm. The impact of temperature, gas flow rate and IL concentrations on Rm will be discussed in the next section.

4.4 Mass Transfer Coefficient in Membrane

4.4.1 Introduction

The mass transfer coefficient in the membrane is analogous to the permeance for a given component diffusing through a membrane of a given thickness. The permeance (Pm) is defined as the ratio of the permeability coefficient (P) to the membrane thickness (l). The permeability of component i is a product of two terms as equation (47):

Pi  S i  Di (49) where Si is the sorption (or partition) coefficient which is an equilibrium ratio of the concentration of the component i in a fluid phase to the concentration in the membrane polymer phase, and Di is the diffusion coefficient in the membrane which reflects the effect of the surrounding environment on the molecular motion of component i.

Based on Equation (30), the membrane resistance can be expressed as follows:

l R  m D  H'' (50)

Membranes used in modules so far working in industrial pervaporation plants are generally of composite type. They are prepared by coating a porous support of definite structure with a thin, dense layer of permselective polymer. During its formation, the superimposed skin is slightly crosslinked to reduce its ability to swell. The pores in the support layer are supposed to be wide enough to avoid undesirable pressure drop in the permeate stream.

The pervaporation membrane in this research is RO AK polyamide membrane purchased from GE labstore which consists of three different parts: polyamide dense layer,

asymmetric polysulfone UF layer and polyester nonwoven support layer. The selective layer is the dense layer and contributes to the main resistance of the membrane. Figure 4-

7 shows the schematic view of the RO AK membrane:

Figure 4-7 Schematic view of the RO AK membrane

The parameter l in Equation (50) stands for the thickness of the dense layer as shown in

Figure 4-7. The solvent molecules can diffuse into the polymer to produce a swollen gel.

If the polymer-polymer intermolecular forces are high, the membrane could not change shape due to high crosslinking, crystallinity or strong hydrogen bonding. However, if the introduction of strong polymer-solvent interactions can overcome the polymer-polymer interactions, the dissolution of the polymer by the solvent would take place. The polymer in the membrane can be expanded and as a consequence the thickness of dense layer would increase along with the increment of membrane resistance according to Equation

(50). Therefore, the fluid contacting the membrane may change the structure of the dense layer and affect mass transport through the membrane.

The self-diffusion coefficient of water in the membrane increases with water content and temperature which will result in decrease of mass transfer resistance in membrane.

H'' in Equation (50) is the partition coefficient which is the ratio of partial pressure of water in the membrane to the partial pressure of water at the interface outside the membrane. Partition coefficient indicates the extent to which the component will partition into the medium of interest and may be concentration dependent. This parameter is important to analyze and related to the mass transfer resistance in membrane.

A careful choice of the membrane material is crucial for an efficient transport process.

The pervaporation membrane in this research is RO AK polyamide membrane which is hydrophilic. Hydrophilic membrane has good attraction with water molecules and can help the transport of water from aqueous IL to membrane polymers.

4.4.2 Results

Based on experiment data for overall mass transfer resistance calculation with pure water as feed (Table 4.1, 4.2, 4.3, 4.4) and the results of mass transfer resistance estimation

(Table 4.5) in previous section, the mass transfer resistance in membrane can be simply derived by subtracting the gas phase resistance from the overall mass transfer resistance and using regression to obtain the parameters in Equation (48).

Figure 4-8 and 4-9 shows the calculated data for the mass transfer resistance in the membrane for different flow rates of sweep gas at room temperature and 40ºC respectively:

Figure 4-8 Mass transfer coefficient in membrane at 24ºC

Figure 4-9 Mass transfer coefficient in membrane at 40ºC

As can be seen in Figure 4-8, the mass transfer resistance in membrane is increasing with

IL concentration especially at high concentrations. The diffusivity of water in the membrane may decrease as the IL concentration rises due to swelling of the membrane by IL and the dependence of water diffusivity on IL concentration which would be consistent with the observed increase in resistance.

The dependence of partition coefficient of water on the IL concentrations in liquid phase also needs to be investigated to evaluate its contribution to Rm. The flow rate of the gas also has a big impact on membrane resistance and Rm is larger at lower gas flow rates. As sweep gas flow rate decreases, the water concentration in the permeate side will be larger.

Water in the permeate side can swell the membrane and the membrane thickness will

increase. Hence, the resistance in the membrane can increase due to lower gas flow rate.

Compare

Figure 4-8 with 4-9 indicate increasing the temperature decreases membrane resistance.

As the diffusivity of water in the membrane can be highly dependent on temperature, the mass transport of molecules can be accelerated by temperature increment.

4.4 Conclusion

The mass transfer resistance for pervaporation is mainly in membrane. Increasing temperature can highly increase the water flux due to the contribution of increased water vapor pressure difference and also mass transfer coefficients. The mass transfer resistance in the membrane can increase with the decrease of gas flow rate. A new membrane with better mass transport performance is demanded.

Chapter 5

Energy Consumption of Vacuum Evaporation vs. Pervaporation for Aqueous IL Dehydration

5.1 Introduction

As ionic liquids are non-volatile, evaporation is one of the possible processes for the recovery of ILs from the mixture. However, conventional evaporation is an energy- intensive and time-consuming method for the removal of water or organic solvents.

Membrane based separation process is a low-cost and energy-efficient technology [64, 65,

66].

This chapter will compare the process costs for evaporation with pervaporation for aqueous ILs.

5.2 Vacuum Evaporation for Aqueous IL

5.2.1 Introduction

Evaporation is a process used to concentrate a solution consisting of a non-volatile solute and a volatile solvent. Water is the overwhelming majority solvent of evaporation system.

Evaporation differs from distillation in that the vapor phase usually is a single component, and even when the vapor is a mixture, no attempt is made in the evaporation step to separate the vapor into fractions. In most of the evaporation operations, the thick liquor is

the valuable product, and the vapor is condensed and discarded. As ionic liquids have non-measurable vapor pressure, the diluted aqueous ILs can be concentrated with evaporation technique. Recent studies apply vacuum in evaporation process to purify ILs from different solvents due to the extremely low vapor pressure of IL containing system especially at high concentrations [67, 68].

Most evaporators are heated with steam condensing on the outside of metal tubes.

Usually the liquid to be concentrated is under vacuum. The vacuum pressure needs to be chosen below the vapor pressure of the aqueous IL to evaporate the water at a certain temperature. The temperature of the system also needs to be carefully controlled as the IL will degrade at high temperatures. The temperature difference between the steam and the liquid has two parts of contributions: heating the liquid to the temperature which evaporation could happen at operated vacuum; heat of water evaporation from aqueous

IL.

Multiple stages of evaporators can be used in the system to reduce the energy cost. The vapor pressure of aqueous IL is highly dependent on IL concentration and the concentration of the product in the effect determines the vacuum applied in the system.

Applying different vacuum pressures to multiple stages can reduce the energy cost.

Smaller pressure difference is needed for the evaporation of low IL concentration ranges.

Less amount of water needs to be concentrated in the following effect of higher IL concentration range. Consequently, the total energy cost for the vacuum for all the stages can be minimized.

5.2.2 Theory

5.2.2.1 Single-effect Evaporator

Figure 5-1 shows diagrammatically a vertical-tube, single-effect evaporator.

Figure 5-1 Material and enthalpy balances in evaporator

The rate of steam flow and condensate is ms, that of feed is mf, and that of the thick liquor is m. The rate of vapor flow to the condenser is mf - m. Let Tf to be the temperature of the feed, T the temperature of the liquid at which water evaporation starts in the evaporator under vacuum, and Ts the condensing temperature of the steam. The material to be

evaporated flows inside under vacuum and the vapor from the heated liquid is condensed and discarded. The steam entering the steam chest may be superheated, and the condensate usually leaves the steam chest subcooled below its boiling point. It is acceptable to neglect the superheat and the subcooling of the condensate in making an enthalpy balance as they are small. Hence, the difference between the enthalpy of the steam and condensate is λs which is the latent heat of condensation of the steam. The enthalpy balance for the steam line is

qs  ms (H s  H c )  mss (51) where

qs – rate of heat transfer through heating surface from steam,

Hs – specific enthalpy of steam,

Hc – specific enthalpy of condensate,

λs – latent heat of condensation of steam,

ms – rate of flow of steam,

The enthalpy balance for the liquor side is

q  (m f  m)Hv  m f H f  mH (52) where

q – rate of heat transfer from heating surface to liquid,

Hv – specific enthalpy of vapor,

Hf – specific enthalpy of feed,

H – specific enthalpy for thick liquor,

Suppose no heat loss, the heat transferred from the steam to the tubes equals that transferred from the tubes to the liquor. Combine Equation (51) and (52):

q  mss  (m f  m)Hv  m f H f  mH (53)

The liquor-side enthalpy Hv, Hf , and H depend on the characteristics of the solution being concentrated. The dependence of the rate of heat transfer q through the heating surface of an evaporator on the area of the heat-transfer surface A, the overall heat-transfer coefficient U, and the overall temperature drop ΔT can be expressed as:

q UAT (54)

5.2.2.4 Enthalpy for Aqueous 1-ethyl-3-methylimidazolium Acetate

The enthalpy of aqueous IL dependent on temperature is demanded to evaluate the energy cost during a thermal process.

The enthalpy change of aqueous IL with evaporation at constant temperature can be measured using TA Instruments SDT-Q600 Simultaneous TGA/DSC. The enthalpy of vaporization of water can be found from previous research [69]. The mass enthalpy change of aqueous IL can be calculated by deducting the mass heat of vaporization of pure water from the mass enthalpy measured by the TGA/DSC.

5.2.2.5 Energy Consuming from Compressor and Vacuum

Most of the energy cost of evaporation is from compressing the steam and vacuum. The theoretical power for compressors, fans, blowers and vacuum pumps can be expressed as

Equation (55).

(55)

where T1 is inlet temperature, R is gas constant, z1 is compressibility, m is molar flow rate,

P2 is outlet pressure of compressor or upstream pressure of vacuum, P1 is inlet pressure

of compressor or downstream pressure of vacuum, a = (k-1)/k, k = Cp/Cv, and Cp and Cv are the heat capacity of compressed gas at constant pressure and volume respectively. For steam, k = 1.32. The actual power needed by a compressor or vacuum is increased as the efficiency is less than 1. The actual power can be calculated as:

(56) where η is the efficiency of the compressor or vacuum.

5.2.3 Results

5.2.3.1 Enthalpy of Aqueous IL

The enthalpy of aqueous IL with different temperatures is tested by TA Instruments SDT-

Q600 Simultaneous TGA/DSC. The data is shown in Figure 5-3.

Figure 5-3 Enthalpy of aqueous IL at different temperatures

From Figure 5-3 we can see that the enthalpy is increasing with IL concentrations especially for high concentration. This can be easily understood that when adding water to high concentration of aqueous IL the system gives large amounts of heat. Increasing the temperature also increases the enthalpy of the system.

5.2.3.2 Energy Cost Calculation for Evaporation

5.2.3.2.1 Single-effect Evaporator

Assume a single-effect evaporator is used to concentrate 44 lb/day (20kg/day) of 40 wt% aqueous IL with 104 ºF (40 ºC) to 95 wt%. The pressure of the steam is set to be 247.1 psi which is corresponding to temperature of 400 ºF. The temperature of solution at

100

which evaporation starts needs to be carefully chosen. Higher temperature of solution gives higher vapor pressure and lower energy cost for the vacuum. However, the temperature is recommended to operated at no more than 302 ºF (150 ºC) to prevent the degradation of IL. Hence, the optimum temperature chosen for the liquid to evaporate could be 302 ºF. Also, the vacuum pressure can be selected slightly lower than the vapor pressure of aqueous IL with certain concentration at 302 ºF. For single-effect evaporator, the vapor pressure of 95 wt% IL at 302 ºF is 2.79 kPa as calculated by extrapolation from the previous work [70] and the vacuum can be chosen as 2.7 kPa. The overall coefficient is estimated to be 350 Btu/ft2·h·ºF (1987 W/m2·ºC) [71]. The amount of steam consumed, the energy cost, and the heating surface area can be determined.

The amount of water evaporated is calculated from the mass balance. The flow rate of the thick liquor (IL) is 0.77 lb/h. The rate of water evaporation is 1.06 lb/h. The enthalpies of the feed at 40°C and thick liquor at 150°C can be found from Fig. 5-3: Hf = 305.9

Btu/lb and H = 1129.4 Btu/lb.

The enthalpy of the vapor leaving the evaporator is found from steam tables [71]. The enthalpy of superheated water vapor Hv at 302 ºF and 0.39 psi is 1196.8 Btu/lb based on

Table of Superheated Steam [72]. The heat of vaporization of steam λs at a pressure of

89.6 psi is 826.8 Btu/lb.

The rate of heat transfer and the steam consumption can be found from Equation (53):

101

The condensation temperature of the steam is 400 ºF. The heating area required can be calculated from Equation (54).

The cost of the steam can be easily calculated based on the amount of steam needed in the process and the unit cost of the steam. The cost is 8.65 $/1000kg for high pressure steam [73]. The total steam cost for evaporation is 0.18 $/day.

Suppose the steam compressibility is 1, the efficiency of the vacuum is 0.75 running 24 hours/day and the cost of electricity is 12¢/ kW·h. Based on Equation (55) and (56), the process of energy costs calculation for vacuum can be listed in Table 5.1.

Table 5.1 Energy cost for vacuum of single-effect evaporator

m P2 P1 T1 W Wactual cost

mole/s psi psi ºF kW·h kW·h $/day

0.007 14.70 0.39 302 0.15 0.20 0.58

The pumping cost for the solution is negligible. The total energy costs of evaporation from the cost of the steam supply and vacuum are calculated to be 0.76 $/day.

5.2.3.2.2 Double-effect Evaporator

As discussed in previous section, the water from aqueous IL can be vacuum evaporated to concentration the mixture. The vapor pressure of 58 wt% IL at 150 °C is 15.3 psi which is still above atmosphere pressure. To reduce the energy cost of evaporator, the 40

102

wt% feed can be heated directly to 58 wt% and use only one vacuum pump for the next stage to concentrate aqueous IL from 58 wt% to 95 wt%.

A double-effect evaporator is designed. Effect I is to concentrate 20 kg 40 wt% aqueous

IL to 58 wt% without vacuum. Effect II is to concentrate 58 wt% aqueous IL from effect

I to 95 wt% under vacuum.

In effect I, the operation temperature is recommended to be 302°F. Saturated steam at

400°F is used for heating. The flow rate of the feed is 20 kg/day (1.83 lb/hr) and the product is 13.8 kg/day (1.27 lb/hr). The enthalpies of the feed at 40°C and thick liquor at

302°F are found from Fig. 5-3: Hf = 305.9 Btu/lb and H = 617.5 Btu/lb. The enthalpy of superheated vapor at 302°F under 1 atm is 1193.5 Btu/lb. The overall heat transfer coefficient for 40 wt% to 58 wt% is estimated to be 400 Btu/ft2·h·ºF (2271 W/m2·ºC).

Hence, the heat needed, steam flow rate and heat transfer area can be calculated:

Based on the flow rate of steam, the total steam cost for evaporation is 0.10 $/day.

The product from effect I at 302 °F is sent to effect II under vacuum pressure of 0.39 psi

(2.7 kPa) to concentrate 58 wt% aqueous IL to 95 wt%. Saturated steam at 400°F is used for heating. The flow rate of feed and product is 1.27 lb/hr and 0.77 lb/hr. The enthalpy of feed and product is 617.5 Btu/lb and 1129.4 Btu/lb. The enthalpy of superheated vapor

103

at 302F under pressure of 0.39 psi is 1196.8 Btu/lb. The overall heat transfer coefficient for 58 wt% to 95 wt% is estimated to be 350 Btu/ft2·h·ºF (1987 W/m2·ºC). The heat needed, steam flow rate and heat transfer area for effect II can be calculated:

The cost of the steam in effect II can be calculated to be 0.08 $/day.

The process of calculation of energy costs for vacuum in effect II can be listed in Table

5.2:

Table 5.2 Energy cost for vacuum of effect II evaporator

m P2 P1 T1 W Wactual cost

mole/s psi psi ºF kW·h kW·h $/day

0.003 14.70 0.39 302 0.07 0.09 0.27

The total energy costs in both effects are 0.45 $/day. Compare with the costs in single effect, double effect can save about 40% of the energy. The huge amount of energy saving in double-effect is due to heating aqueous IL at high water content under atmosphere instead of vacuum.

104

5.3 Pervaporation

5.3.1 Introduction

The processing cost for pervaporation includes the cost of the membrane and the energy cost for the operation. The membrane area is calculated based on the feed and product conditions through the process and the conditions of the dry gas. The energy cost is mainly from providing a dry sweep gas.

5.3.2 Membrane Area and Energy Cost Calculation

5.3.2.1 Theory

Counter current pervaporation process is designed to concentrate 20 kg/day of 40 wt% aqueous IL to 95 wt% using one stage of membrane separation. As the effect of liquid flow rate on the water flux through membrane is negligible, the flow rate of feed can be set to be coordinated to the requirement of the goal. The dry gas can be produced from a desiccant compressed air dryer. Figure 5-2 shows the counter current pervaporation set up for concentration of aqueous IL.

Figure 5-2 Pervaporation of aqueous IL

105

The mass transfer balance in the pervaporation system can be expressed as equation (57):

(57) where is the aqueous IL mass flow rate, is IL weight fraction in the liquid, is the mass transfer coefficient based on certain operation conditions, is water partial pressure in liquid side, is water partial pressure in gas side, is the width of the membrane and is the length of the membrane. The membrane area can be calculated by the product of and .

In equation (57), the mass transfer coefficient and water partial pressure in aqueous IL

is a function of IL concentration . As water permeates and accumulates into the permeate side, the water partial pressure in the gas side is also dependent on the IL concentration , the flow rate of the gas and length of the membrane . The conditions of aqueous IL liquid and dry gas in the product end is known and the membrane area can be calculated by integration of Equation (57) from the product IL concentration backward to the feed side until reaching the point of the initial feed concentration.

5.3.2.2 Results

Take the pervaporation experiment at 40°C with QIL = 60 ml/min and Qgas = 0.5 L/min in this research as an example and use the mass transfer coefficient under such conditions to calculate the membrane area needed. The whole process is operated in one sheet of membrane to concentrate 20kg of 40 wt% aqueous IL to 95 wt% in one day.

The relationship of the water partial pressure in aqueous IL and IL wt% at 40°C can be seen in Figure 5-3. The regression expression equation can be generated.

106

Figure 5-3 Regression of water partial pressure vs. IL weight fraction at 40°C

The relationship of mass transfer coefficient K0 and the IL weight concentration can be derived as Figure 5-4:

107

Figure 5-4 Regression of overall mass transfer coefficient vs. IL weight fraction at 40°C

The regressions of water partial pressure and overall mass transfer coefficient vs. IL weight fraction are generated respectively for IL concentration from 40 wt% - 80 wt% and >80 wt% to achieve high accuracy of the energy cost estimation.

Suppose the inlet dry gas has dew point of -40°C and the water partial pressure can be calculated with Goff-Gratch water dew point correlations. To calculate the membrane area using Equation (51), use the known conditions of the product end to integrate by small IL concentration intervals ( ) until the IL concentration reaches to 40 wt%. Take the integration in the first 0.1 wt% interval (95.00 wt% to 94.9 wt%) as an example, and the calculation process is shown as below:

108

Table 5.3 Calculation process of integration in the first concentration interval

l Cwater in (wt %) 5.1 %

l Cwater out (wt %) 5.0 %

Qwater (mole/s) 5.7E-05

Qdry gas (mole/s) 2

2 K0 (g/s/m /pa) 1.09E-06

l pwater (pa) 20.7

g pwater in (pa) 18.9

Qwater accumulated (mole/s) 3.79E-04

2 Jwater (g/s/m ) 1.98E-06

W×dz (m2) 51.87

g pwater out (pa) 19.18

In Table 5.3, apply Equation (51) to integrate from the starting point of the concentration

l l interval Cwater in of 5.1 wt% to the ending point Cwater out of 5.0 wt%. The amount of water permeated to the gas side Qwater can be calculated based on the raw feed condition

(20 kg/day of 40 wt% IL). The flow rate of the gas Qdry gas is a trial number which is chosen small to cut down the cost of the gas and also high enough to guarantee the water concentration in the gas side to be smaller than the liquid side. The mass transfer coefficient K0 is calculated based on the regression equation in Figure 5-4. The partial

l pressure of water in the liquid side pwater is derived from Figure 5-3. Suppose the initial

g dry gas has dew point of -40°C and the water partial pressure pwater in is calculated based

109

on Goff-Gratch water dew point correlations. The amount of water accumulated Qwater accumulated is the sum of the amount of water passing through the membrane and the water in the dry gas. The water flux through the membrane Jwater is calculated by the product of mass transfer coefficient and the water partial pressure difference between the liquid and gas side. The membrane area W×dz can be derived from the ratio of the amount of water permeated and the water flux through the membrane. The partial pressure of water

g coming out of the concentration interval pwater out is calculated based on the water partial pressure coming in, the amount of water permeated and the flow rate of the dry gas. The

g air coming out of the first concentration interval with water partial pressure pwater out is introduced into the second concentration interval as the initial dry gas. Apply the same process to calculate the membrane area in the following concentration intervals until the

IL concentration reaches to 40 wt%.

Sum of the membrane area calculated in every interval from 95 wt% to 40 wt% of aqueous IL and the total membrane area can be estimated to be 548.5 m2. In the experiment of the research, the gas flows at 0.5 L/min in a channel with 94.5 mm width and 0.4 mm height. To keep the same velocity of the gas, the 2 mole/s (2688 L/min) of air needs to flow in a channel with 508.0 m width and the same height. The length of the membrane can be calculated to be 1.1 m.

The partial pressure of water in the gas side is smaller than the liquid side for every interval and the membrane area is not changing with higher numbers of intervals. Hence, the total area should be accurate enough.

110

Van Air HL Series heatless regenerative desiccant compressed air dryer [74] can be used to generate dry gas for pervaporation. HL Series air dryers with 115V/1PH/60HZ supply power provide a low -40° or -100°F pressure dew point between 200 and 5000 SCFM.

The optimum amount of air needed for concentrate 20 kg of 40 wt% aqueous IL is chosen to be 2 mole/s (95.0 SCFM). Suppose using the desiccant dryer to generate 2000 SCFM of air with DP of -40°F, it can concentrate 2000/95 = 21 times of the setting amount of feed. The power of the desiccant dryer is 1 PH (0.735KW). Suppose the system is working 24 hr/day and the electricity cost is 0.12 $/kWh. The cost for concentrate 20 kg of 40 wt% to 95 wt % aqueous IL is 0.10 $/day.

5.4 Conclusion

Both vacuum evaporator and membrane pervaporation can be used to concentrate aqueous IL. Heating the dilute IL without vacuum as first effect in double-effect evaporator can save a big amount of energy compared with single effect evaporator.

Membrane pervaporation is more energy saving compared with evaporators. Instead of using vacuum pump to generate water vapor pressure difference between the liquid and gas phase, dry gas can be less costly.

111

Chapter 6

Conclusions and Recommendation for Future Work

6.1 Conclusions

This dissertation reports the studies of recovery of 1-ethyl-3-methylimidazolium acetate from aqueous solution. The physico-chemical properties of ionic liquid including vapor pressure, viscosity, conductivity and diffusivity are investigated for full range of IL concentrations with different temperatures. Two membrane processes (reverse osmosis and pervaporation) are discussed to concentrate the aqueous IL. The mass transport mechanisms of water flux through pervaporation process are. The economies of membrane pervaporation and evaporation are calculated and compared to concentrate dilute IL to 95 wt%. The following conclusions can be drawn from this work:

1. The physico-chemical properties of aqueous IL are highly dependent on the IL

concentrations especially at high concentrations. Water molecules tend to

combine with anions at low water content with hydrogen bonding. The interaction

of water and anion results in smaller diffusivity of anion compared with larger

cation and extremely low vapor pressure at high IL concentrations. The removal

of water with low water content becomes challenging.

112

2. Both straight RO and pervaporation can dehydrate aqueous IL. Straight RO

process needs high energy even at low IL concentration and the concentration rate

is low with IL lost into the permeate side. The pervaporation is more promising as

the IL can be concentrated to more than 95 wt%. The water flux is increasing with

system temperature.

3. The mass transfer resistance for pervaporation is mainly in membrane. Increasing

temperature can highly increase the water flux due to the contribution of increased

water vapor pressure difference and also mass transfer coefficients.

4. Recovery of IL from aqueous solution using membrane pervaporation is less

energy consuming than comparing with vacuum evaporation. The extremely low

vapor pressure of aqueous IL at high concentrations limits the energy efficiency

of the vacuum evaporation. The dry gas from high IL concentration stage in the

pervaporation system can be recycled to low IL concentration stage to decrease

the energy cost.

6.2 Future Work

Temperature plays an important role to increase the water flux of pervaporation which is mainly due to the water vapor pressure increment. The membrane in this research is polyamide RO membrane from GE osmonics labstore which can only withstand up to

50ºC operation temperature. An inert, chemical resistant and high temperature membrane is demanded to enhance the mass transport rate of water through the membrane. The mainly resistance for mass transport lays in the membrane, hence, a membrane with proper materials having better mass transfer performance on aqueous IL system needs to be investigated. Highly hydrophilic material is recommended for future investigation to

113

reduce the membrane resistance for water transport. As a sequence, the energy cost of pervaporation can be decreased dramatically.

114

Reference

1 . C. L. Hussey, Room Temperature molten salt systems. Adv. Molten Salt Chem., 5

(1983) 185-230.

2 . C. L. Hussey, Room-temperature haloaluminate ionic liquids-novel solvents for transition-metal solution chemistry. Pure Appl. Chem., 60 (1988) 1763-1772.

3 . K. R. Seddon, Ionic Liquids for Clean Technology, J. Chem. Tech. Biotechnol. 68

(1997) 351-356.

4 . J. Robinson, R. A. Osteryoung, An electrochemical and spectroscopic study of some aromatic hydrocarbons in the room temperature molten salt system aluminum chloride-n- butylpyridinium chloride. J. Am. Chem. Soc., 101 (1979) 323-327.

5 . S. Zhu, Y. Wu, Q. Chen, Z. Yu, C. Wang, S. Jin, Y. Ding, G Wu, Dissolution of cellulose with ionic liquids and its application: a mini-review, Green Chem., 8 (2006)

325-327.

6 . M. J. Earle, K. R. Seddon, Ionic liquids. Green solvents for the future, Pure Appl.

Chem., 72 (2000) 1391-1398.

7 . M. J. Earle, J. M. S. S. Esperanca, M. A. Gilea, J. N. C. Lopes, L. P. N. Rebelo, J.

W. Magee, K. R. Seddon, J. A. Widegren, The distillation and volatility of ionic liquids, nature, 439 (2006) 831-834.

8 . G. Reginald, J. W. David, Process for the transalkylation of dealkylation of alkyl aromatic hydrocarbons, U.S. Patent 4,599,470, 1986.

115

9 . J. K. Baranyai, G. B. Deacon, D. R. MacFarlane, J. M. Pringle, J. L. Scott,

Thermal Degradation of Ionic Liquids at Elevated Temperatures, Aust. J. Chem., 57

(2004) 145 – 147.

10 . G. Ciardelli, L. Corsi, M. Marcucci, Membrane separation for wastewater reuse in the textile industry, Resour., Conserv. Recycling, 2 (2001) 189-197.

11 . M. T. Ravanchi, T. Kaghazchi, A. Kargari, Application of membrane separation processes in petrochemical industry: a review, Desalination, 235 (2009) 199-244.

12 . M. Mulder, Basic Principles of Membrane Technology, 2nd Edition, Kluwer

Academic, Dordrecht, The Netherlands, 2000.

13. H. Ohno, Electrochemical Aspects of Ionic Liquids, Wiley-Interscience, Hoboken,

(2005).

14 . F. van Rantwijk, R. M. Lau, R. A. Sheldon, Biocatalytic transformations in ionic liquids, Trends Biotechnol, 21 (2003) 131-138.

15 . R. E. Baltus, R. M. Counce, B. H. Culbertson, H. Luo, D. W. DePaoli, S. Dai, D.

C. Duckworth, Examination of the Potential of Ionic Liquids for Gas Separations, Separ.

Sci. Technol., 40 (2005) 525-541.

16 . P. N. Tshibangu, S. N. Ndwandwe, E. D. Dikio, Density, Viscosity and

Conductivity Study of 1-Butyl-3-Methylimidazolium Bromide, Int. J. Electrochem. Sci.,

6 (2011) 2201-2213.

17 . Y. Chen, F. Ke, H. Wang, Y. Zhang, D. Liang, Phase Separation in Mixtures of

Ionic liquids and Water, ChemPhysChem, 13 (2011) 160-167.

116

18 . J. G. Huddleston, A. E. Visser, W. M. Reichert, H. D. Willauer, G. A. Broker, R.

D. Rogers, Characterization and comparison of hydrophilic and hydrophobic room temperature ionic liquids incorporating the imidazolium cation, Green Chem., 3 (2001)

156-164.

19 . K. R. Seddon, A. Stark, M. J. Torres, Influence of chloride, water, and organic solvents on the physical properties of ionic liquids, Pure Appl. Chem, 72 (2000) 2275-

2287.

20 . M. Eckstein, M. Sensing, U. Kragl, P. Adlercreutz, At low water activity α- chymotrypsin is more active in an ionic liquid than in non-ionic organic solvents,

Biotechnol Lett., 24 (2002) 867-876.

21 . J. A. Berberich, J. L. Kaar, A. J. Russell, Use of salt hydrate pairs to control water activity for enzyme catalysis in ionic liquids. Biotechnol Prog., 19 (2003) 1029-1032.

22 . K. R. Seddon, A. Stark, Selective catalytic oxidation of benzyl alcohol and alkylbenzenes in ionic liquids, Green Chem., 4 (2002) 119-123.

23 . M. Bier, S. Dietrich, Vapour pressure of ionic liquid, Mol. Phys., 2 (2010) 211-214.

24 . W. Jiang, Y. Wang, G. A. Voth, Molecular Dynamics Simulation of Nanostructural

Organization in Ionic Liquid/Water Mixtures, J. Phys. Chem. B., 111 (2007) 4812-4818.

25 . L. Cammarata, S. G. Kazarian, P. A. Salter, T. Welton, Molecular states of water in room temperature ionic liquids, Phys. Chem. Chem. Phys., 3 (2001) 5192-5200.

26 . M. S. Kelkar, E. J. Maginn, Effect of Temperature and Water Content on the Shear

Viscosity of the Ionic Liquid 1-Ethyl-3-methylimidazolium

117

Bis(trifluoromethanesulfonyl)imide As Studied by Atomistic Simulations, J. Phys. Chem.

B., 111 (2007) 4867-4876.

27 . C. G. Hanke, R. M. Lynden-Bell, A Simulation Study of

Water−Dialkylimidazolium Ionic Liquid Mixtures, J. Phys. Chem. B., 107 (2003) 10873-

10878

28 . J. Bowers, C. P. Butts, P. J. Martin, M. C. Vergara-Gutierrez, Aggregation

Behavior of Aqueous Solutions of Ionic Liquids, Langmuir, 20 (2004) 2192-2198.

29 . G. Reginald, J. W. David, Process for the transalkylation of dealkylation of alkyl aromatic hydrocarbons, U.S. Patent 4,599,470, 1986.

30 . H. B. Zhao, J. E. Holladay, H. Brown, Z. C. Zhang, Metal chlorides in ionic liquid solvents convert sugars to 5-hydroxymethylfurfural, Science, 316 (2007) 1597-1600.

31 . Y. Fukaya, K. Hayashi, M. Wada, H. Ohno, Cellulose dissolution with polar ionic liquids under mild conditions: Required factors for anions, Green Chem., 10 (2008) 44-

32 . S. H. Lee, T. V. Doherty, R. J. Linhardt, J. S. Dordick, Ionic liquid-mediated selective extraction of lignin from wood leading to enhanced enzymatic cellulose hydrolysis, Biotchnol. Bioeng., 102 (2009) 1368-1376.

33 . H. Zhao, L. Jackson, Z. Y. Song, A. Olubajo, Using ionic liquid

[EMIM][CH3COO] as an enzyme-'friendly' co-solvent for resolution of amino acids,

Tetrahedron: Asymmytry, 17 (2006) 2491-2498.

34 . M. B. Shiflett, A. Yokozeki, Phase Behavior of Carbon Dioxide in Ionic Liquids:

[emim][Acetate], [emim][Trifluoroacetate], and [emim][Acetate] +

[emim][Trifluoroacetate] Mixtures, J. Chem. Eng. Data., 54 (2009) 108-114.

118

35 . K. J. Baranyai, G. B. Deacon, D. R. MacFarlane, J. M. Pringle, J. L. Scott,

Thermal Degradation of Ionic Liquids at Elevated Temperatures, Aust. J. Chem., 57

(2004) 145-147.

36 . B. Cornils, W. A. Herrmann, I. T. Horvath, W. Leitner, S. Mecking, H. Olivier-

Bourbigou, D. Vogt (Eds.), Multiphase Homogeneious Catalysis, Wiley-VCH,

Weinheim, 2005.

37 . D. J. Cole-Hamilton, R. P. Tooze (Eds.), Catalyst Separation, Recovery and

Recycling: Chemistry and Process Design, Springer, Dordrecht, 2006.

38 . G. F. Mara, M. N. B. F. S. Luís, M. F. Ana, A. P. C. João, M. M. Isabel, An overview of the mutual solubilities of water-imidazolium-based ionic liquids systems,

Fluid Phase Equilibr., 261 (2007) 449-454.

39 . M. Yoshizawa, W. Xu, C. A. Angell, Ionic liquids by proton transfer: Vapor pressure, conductivity, and the relevance of ∆pKa from aqueous solutions, J. Am. Chem.

Soc., 125 (2003) 15411-15419.

40 . U. P. Kreher, A. E. Rosamilia, C. L. Raston, J. L. Scott, C. R. Strauss, Self- associated, “Distillable” Ionic Media, Molecules, 9 (2004) 387–393.

41 . S. Tanimura, S. Nakao, S. Kimura, Ethanol-selective membrane for reverse osmosis of ethanol/water mixture, AICHE J., 36 (1990) 1118-1120.

42 . R. W. Backer, Membrane Technology and Applications, Willey, (2004).

43 . S. Bhattacharya, S. T. Hwang, Concentration polarization, separation factor, and

Peclet number in membrane processes, J. Membr. Sci., 132 (1997) 73-90.

119

44 . C. Chiappe, D. Pieraccini, Ionic liquids: solvent properties and organic reactivity, J.

Phys. Org. Chem., 18 (2005) 275-297.

45 . A, Stoppa, J. Hunger, R. Buchner, Conductivities of Binary Mixture of Ionic

Liquids with Polar Solvents, J. Chem. Eng. Data., 54 (2009) 472-479.

46 . A. L. Zhu, J. J. Wang, L. J. Han, M. H. Fan, Measurements and Correlation of

Viscosities and Conductivities for the Mixtures of Imidazoium Ionic Liquids with

Molecular Solutes, Chem. Eng. J., 147 (2009) 27-35.

47 . G. Y. Zhao, T. Jiang, W. Z. Wu, X. B. Han, Z. M. Liu, H. X. Gao, Elctro- oxidation of Benzyl Alcohol in a Biphasic System Consisting of Supercritical CO2 and

Ionic Liquids, J. Phy. Chem. B., 108 (2004) 13052-43057.

48 . W. Zhao, F. Leroy, B. Heggen, S. Zahn, B. Kirchner, S. Balasubramanian, F.

Müller-Plathe, Are There Stable Ion-Pairs in Room-Temperature Ionic Liquids?

Molecular Dynamics Simulations of 1-n-Butyl-3-methylimidazolium

Hexafluorophosphate, J. AM. CHEM. SOC., 131 (2009) 15825-15833.

49 . http://www.vernier.com/experiments/cwv/10/vapor_pressure_of_liquids/

50 . J. A. Goff, S. Gratch, Low-pressure properties of water from -160º to 212ºF, Trans.

Amer. Sco. Heat Vent. Eng., 52 (1946) 95-122.

51 . J. Vila, E. Rilo, L. Segade, O. Cabeza, L. M. Varela, Electrical conductivity of aqueous solutions of aluminum salts, Phys. Rev. E., 71 (2005) 031201.

52 . P. Claes, Y. Loix, Maximum De Conductivity Des Solutions Aqueuses

D’electrolytes, J. Glibert, Electrochim. Acta., 28 (1983) 421-427.

120

53 . J. O’M. Bockris, A. K. N. Reddy, Modern Electrochemistry, vol. 1, Plenum Press,

New York, 1998.

54 . A. Heintz, D. Klasen, J. K. Lehmann, Excess molar volumes and viscosities of binary mixtures of methanol and the ionic liquid 4-methyl-N-butylpyridinium tetrafluoroborate at 25, 40, and 50ºC, J. Sol. Chem., 31 (2002) 467-476.

55 . W. S. Price, Pulsed-Field Gradient Nuclear Magnetic Resonance as a Tool for

Studying Translational Diffusion: Part 1. Basic Theory, Concepts Magn Reson., 9 (1997)

299-336.

56 . G. A. Webb, C. J. Jameson, L. Fukui, K. K. Trela, Nuclear Magnetic Resonance,

Royal Society of Chemistry, (2004).

57 . D. T. Bowron, C. D'Agostino, L. F. Gladden, C. Hardacre, J. D. Holbrey, M. C.

Lagunas, J. McGregor, M. D. Mantle, C. L. Mullan, T. G. A. Youngs, Structure and dynamics of 1-ethyl-3-methylimidazolium acetate via molecular dynamics and neutron diffraction. J. Phys. Chem. B., 114 (2010) 7760-7768.

58 . H. Tokuda, K. Hayamizu, K. Ishii, M. A. B. H. Susan, M. Watanabe,

Physicochemical Properties and Structures of Room Temperature Ionic Liquids. 2.

Variation of Alkyl Chain Length in Imidazolium Cation. J. Phys. Chem. B., 109 (2005)

6103-6110.

59 . AL. Rollet, P. Porion, M. Vaultier, I. Billard, M. Deschamps, C. Bessada, L.

Jouvensal, Anomalous Diffusion of Water in [BMIM][TFSI] Room-Temperature Ionic

Liquid. J. Phys. Chem. B., 111 (2007) 11888-11891.

121

60 . W. Jiang, Y. Wang, G. A. Voth, Molecular Dynamics Simulation of Nanostructural

Organization in Ionic Liquid/Water Mixtures. J. Phys. Chem. B., 111 (2007) 4812-4818.

61 . T. J. Wooster, K. M. Johanson, K. J. Fraser, D. R. MacFarlane, J. L. Scott,

Thermal Degradation of Cyano Containing Ionic Liquids, Green Chem., 8 (2006) 691-

696.

62 . J. G. Huddleston, A. E. Visser, W. M. Reichert, H. D. Willauer, G. A. Broker, R.

D. Rogers, Characterization and Comparison of Hydropilic and Hydrophobic Room

Temperature Ionic Liquids Incorporating the Imidazolium Cation. Green Chem., 3 (2001)

156-164.

63 . R. D. Noble, S. A. Stern, Membrane Separations Technology: Principles and

Applications, Elsevier, (1995).

64 . T. Tsuru, Inorganic porous membranes for liquid phase separation, Sep. Purif.

Methods., 30 (2001) 191-200.

65 . S. C. George, S. Thomas, Transport phenomena through polymeric systems, Prog.

Polym. Sci., 26 (2001) 985-1017.

66 . H. Strathmann, Membrane separation processes: Current relevance and future opportunities, AIChE J., 47 (2001) 1077-1087.

67 . S. Keskin, D. Kayrak-Talay, U˘. Akman, Ö. Hortaçsu, A review of ionic liquids towards supercritical fluid applications. J. Supercrit. Fluids., 43 (2007) 150-180.

68 . M. Zavrel, D. Bross, M. Funke, J. Büchs, A. C. Spiess, High-throughput screening for ionic liquids dissolving (lingo-) cellulose, Bioresour Technol., 100 (2009) 2580-2587.

122

69 . K. N. Marsh, Recommended Reference Materials for the Realization of

Physicochemical Properties, Blackwell, Oxford, 1987.

70 . K. T. Ponnaiyan, A. P. Dadi, S. Varanasi, Characterization of High Quality Lignin

Isolated From Ionic Liquid Pretreated Biomass: Residual Lignin V/s Re-Precipitated

Lignin, International Congress on Energy 2011, AICHE 2011 Annual Meeting

Proceeding.

71 . W. L. McCabe, J. C. Smith, P. Harriot, Unit Operations of Chemical Engineering,

McGraw-Hill, New York, 2001.

72 . J. M. Smith, H. C. Van Ness, M. Abbott, Introduction to Chemical Engineering

Thermodynamics, McGraw-Hill, New York, 2001.

73 . R. Turton, R. C. Bailie, W. B. Whiting, J. A. Shaeiwitz, Analysis, Synthesis, and

Design of Chemical Processes, Prentice-Hall, 1998.

74 . http://www.vanairsystems.com/products_regenerative_Heatless_large.html.

123