i

APPLICATION OF THE HLD AND NAC MODELS TO

THE FORMATION AND STABILITY OF EMULSIONS

By

Sumit Kumar Kiran

A thesis submitted in conformity with the requirements

for the degree of Doctor of Philosophy

Graduate Department of Chemical Engineering and Applied Chemistry

University of Toronto

© Copyright by Sumit Kumar Kiran (2013)

ii

APPLICATION OF THE HLD AND NAC MODELS TO THE

FORMATION AND STABILITY OF EMULSIONS

Sumit Kumar Kiran

Degree of Doctor of Philosophy

Graduate Department of Chemical Engineering and Applied Chemistry

University of Toronto

2013

ABSTRACT

This thesis explored how asphaltene and naphthenic amphiphile species influence the

formation (morphology and size) and stability of heavy crude oil (bitumen) emulsions. It was

experimentally shown that asphaltenes produce water-in-oil emulsions. Naphthenic amphiphiles

on the other hand flip the emulsion morphology to oil-in-water. It was further demonstrated that

the size and stability of these emulsions is influenced by physicochemical effects such as the

pH, solvent-bitumen-water ratios, solvent aromaticity, and temperature. In view of these

findings, the hydrophilic-lipophilic deviation (HLD) and net-average curvature (NAC) models

were looked at as potential means for predicting the formation and stability of emulsions. Owing

to the complexity of bitumen emulsions, however, the HLD and NAC models were instead

tested against well-defined sodium dihexyl sulfosuccinate-toluene-water emulsions. The

morphologies of these emulsions were predicted as a function of the formulation salinity

whereas corresponding droplet sizes were predicted as a function of the continuous phase

density and interfacial tension (γow). Emulsion stability trends were in turn predicted using a

collision-coalescence-separation assumption. From this assumption, emulsion stability was

iii

expressed as a function of the emulsion droplet collision frequency and activation energy. The

key parameters of the highly scrutinized activation energy term included the γow, interfacial

rigidity, and critical film thickness. In applying the same modeling approach to the stability of

other emulsions already published in the literature, it was found that the rigidity of adsorbed

multilayer/liquid crystal films cannot yet be fully accounted for. This shortcoming was the

reason for which only minimum stability times were reported for bitumen emulsions.

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ACKNOWLEDGEMENTS

I would first and foremost like to express my gratitude to my supervisor, Professor Edgar

J. Acosta, for his belief in me as a graduate student and as a leader of his laboratory. His

continuous guidance and support over the course of this work has allowed for me to publish in

several world renowned journals, present at various national and international conferences, and

collaborate with academic and industrial leaders in the field of colloid and formulation science.

He has further served as a close confidant and friend in helping me make tough personal and

professional decisions. I am also thankful to my committee members, Professors A.

Ramchandran, C.A. Mims, and C.M. Yip, for their constructive criticisms and insightful

suggestions to help improve the foundation of this work. Other university staff members that I

would like to recognize for their administrative and technical support include Dan Tomchyshyn,

Gorette Silva, Joan Chen, Julie Mendonca, Leticia Gutierrez, Paul Jowlabar, Pauline Martini,

and Phil Milczarek.

In addition to the above, I am also appreciative of the Natural Sciences and Engineering

Research Council of Canada (NSERC) and Syncrude Canada Ltd. for their financial support. I

would also like to thank the American Oil Chemists’ Society (AOCS) and the Society of

Cosmetic Chemists (SCC) for their generous student awards.

Finally, I would like dedicate this work to my parents, Ravi and Shashi Kiran, my elder

and wiser brother, Amit Kiran, and my best friend and wife, Roshni Patel. All of my

experienced achievements are a reflection of their infinite patience, unwavering love, and

continuous encouragement to never stop chasing my dreams.

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TABLE OF CONTENTS

ABSTRACT ................................................................................................................................. ii

ACKNOWLEDGEMENTS ....................................................................................................... iv

TABLE OF CONTENTS ............................................................................................................ v

LIST OF FIGURES ................................................................................................................... xii

CHAPTER 1: THESIS OVERVIEW ........................................................................................ 1

1.1 OVERVIEW .................................................................................................................. 2

1.2 REFERENCES ............................................................................................................. 8

CHAPTER 2: STUDY OF SOLVENT-BITUMEN-WATER RAG LAYERS .................... 13

2.1 ABSTRACT ................................................................................................................ 14

2.2 INTRODUCTION ...................................................................................................... 14

2.3 MATERIALS AND METHODS ............................................................................... 16

2.3.1 Materials .............................................................................................................. 16

2.3.2 Formulation of Rag Layers .................................................................................. 16

2.3.3 Microscopy .......................................................................................................... 17

2.3.4 Material Balances ................................................................................................ 18

2.3.5 Asphaltene Losses ............................................................................................... 19

2.3.6 Interfacial Tension Measurements ....................................................................... 21

2.4 RESULTS .................................................................................................................... 22

2.4.1 Phase Separation of Solvent-Bitumen-Water Systems ........................................ 22

2.4.2 Development of Batch Emulsification-Separation Protocol................................ 24

2.4.3 Trends of Oil and Asphaltene Losses to the Rag Layer ...................................... 27

2.5 DISCUSSIONS ............................................................................................................ 30

2.5.1 Effect of Heptol 80/20 to Bitumen Dilution Ratio and Water Content on Rag

Layer Stability ..................................................................................................... 30

vi

2.5.2 Effect of Temperature and Solvent Aromaticity on Rag Layer Stability ............ 35

2.5.3 Surface Activity of Asphaltenes .......................................................................... 37

2.5.4 Correlation of Oil and Asphaltene Losses to the Rag Layer ............................... 41

2.6 CONCLUSIONS ......................................................................................................... 42

2.7 REFERENCES ........................................................................................................... 43

CHAPTER 3: IMPACT OF ASPHALTENES AND NAPHTHENIC AMPHIPHILES ON

THE PHASE BEHAVIOR OF SOLVENT-BITUMEN-WATER SYSTEMS .................... 49

3.1 ABSTRACT ................................................................................................................ 50

3.2 INTRODUCTION ...................................................................................................... 50

3.3 MATERIALS AND METHODS ............................................................................... 52

3.3.1 Materials .............................................................................................................. 52

3.3.2 Formulation Preparation ...................................................................................... 52

3.3.3 Microscopy and Material Balances ..................................................................... 53

3.3.4 Asphaltene Losses ............................................................................................... 53

3.3.5 Interfacial Tension Measurements ....................................................................... 54

3.3.6 Surface Pressure-Area Isotherms ......................................................................... 55

3.4 RESULTS .................................................................................................................... 56

3.4.1 Phase Behavior of Naphthenic Amphiphile Systems .......................................... 56

3.4.2 Transitions to the Rag Layer Morphology .......................................................... 59

3.4.3 Interfacial Tension Isotherms .............................................................................. 60

3.4.4 Impact of Naphthenic Acids on Asphaltene Film Properties .............................. 61

3.5 DISCUSSIONS ............................................................................................................ 63

3.5.1 Interfacial Co-Adsorption of Asphaltenes and Sodium Naphthenates ................ 63

3.5.2 Effect of Temperature and Solvent Aromaticity on Oil Recovery from Rag

Layers .................................................................................................................. 67

3.5.3 Effect of pH on Oil Recovery from Rag Layers .................................................. 68

3.6 CONCLUSIONS ......................................................................................................... 69

vii

3.7 REFERENCES ........................................................................................................... 70

CHAPTER 4: EVALUATING THE HYDROPHILIC-LIPOPHILIC NATURE OF

ASPHALTENIC OILS AND NAPHTHENIC AMPHIPHILES USING

MICROEMULSION MODELS ............................................................................................... 73

4.1 ABSTRACT ................................................................................................................ 74

4.2 INTRODUCTION ...................................................................................................... 74

4.3 MATERIALS AND METHODS ............................................................................... 80

4.3.1 Materials .............................................................................................................. 80

4.3.2 Asphaltene Precipitation ...................................................................................... 80

4.3.3 Formulation of Microemulsions with Test Oil and Toluene Mixtures ................ 81

4.3.4 Formulation of Microemulsions with Test Surfactant and SDHS Mixtures ....... 81

4.3.5 Interfacial Tension Measurements ....................................................................... 82

4.3.6 Asphaltene Partitioning at the Oil-Water Interface ............................................. 82

4.4 RESULTS .................................................................................................................... 82

4.4.1 Microemulsion Phase Behavior Scans................................................................. 82

4.4.2 Interfacial Tension of Test Oil and Toluene Mixtures ........................................ 83

4.4.3 EACN of Test Oils............................................................................................... 85

4.4.4 Interfacial Tension of Test Surfactant and SDHS Mixtures ................................ 88

4.4.5 Cc of Test Surfactants .......................................................................................... 90

4.5 DISCUSSIONS ............................................................................................................ 93

4.5.1 Analysis of the Hydrophilic-Lipophilic Nature of Asphaltenic Crude Oils ........ 93

4.5.2 Analysis of the Hydrophilic-Lipophilic Nature of Naphthenic Amphiphiles and

Asphaltene Aggregates ........................................................................................ 94

4.6 CONCLUSIONS ......................................................................................................... 95

4.7 REFERENCES ........................................................................................................... 95

CHAPTER 5: EXPERIMENTAL EVALUATION OF EMULSION STABILITY VIA

SURFACTANT-OIL-WATER PHASE BEHAVIOR SCANS ............................................ 100

5.1 ABSTRACT .............................................................................................................. 101

viii

5.2 INTRODUCTION .................................................................................................... 101

5.3 MATERIALS AND METHODS ............................................................................. 108

5.3.1 Materials ............................................................................................................ 108

5.3.2 Microemulsion Phase Behavior Scans and Emulsification ............................... 108

5.3.3 Interfacial Tension of Baseline Microemulsions ............................................... 109

5.3.4 Average Diameter of Baseline Emulsion Droplets ............................................ 110

5.3.5 Emulsion Phase Separation Profiles .................................................................. 110

5.4 RESULTS .................................................................................................................. 112

5.4.1 Interfacial Tension and Average Emulsion Droplet Diameter .......................... 112

5.4.2 Emulsion Phase Separation Profiles and Time Periods ..................................... 114

5.4.3 Effect of Temperature on Emulsion Phase Separation Time Periods ................ 117

5.5 DISCUSSIONS .......................................................................................................... 120

5.6 CONCLUSIONS ....................................................................................................... 127

5.7 REFERENCES ......................................................................................................... 128

CHAPTER 6: PREDICTING THE MORPHOLOGY AND VISCOSITY OF

MICROEMULSIONS USING THE NAC MODEL ............................................................ 133

6.1 ABSTRACT .............................................................................................................. 134

6.2 INTRODUCTION .................................................................................................... 134

6.3 DEVELOPMENT OF EXPRESSIONS FOR THE SHAPE-BASED NAC

MODEL AND MAXIMUM HYDRODYNAMIC RADIUS ................................. 140

6.4 MATERIALS AND METHODS ............................................................................. 142

6.4.1 Materials ............................................................................................................ 142

6.4.2 Microemulsion Phase Behavior Scans............................................................... 142

6.4.3 Oil-Water Solubilization .................................................................................... 142

6.4.4 Viscosity Measurements .................................................................................... 143

6.4.5 Dynamic Light Scattering .................................................................................. 143

6.4.6 Prediction of Oil and Water Solubilization with the NAC Model..................... 144

ix

6.4.7 Prediction of Small Angle Neutron Scattering Profiles ..................................... 146

6.5 RESULTS AND DISCUSSIONS ............................................................................. 147

6.5.1 Comparison of Spherical Viscosity Models and Experimental Measurements. 147

6.5.2 Comparison of Predicted and Experimental SANS Profiles for Type I and Type

II Microemulsions .............................................................................................. 150

6.5.3 Comparison of Maximum Predicted and Experimental Hydrodynamic Radii .. 152

6.5.4 Comparison of Non-Spherical Viscosity Models and Experimental

Measurements .................................................................................................... 153

6.6 CONCLUSIONS ....................................................................................................... 156

6.7 REFERENCES ......................................................................................................... 156

CHAPTER 7: MODELING THE SIZE AND STABILITY OF EMULSIONS AROUND

THE PHASE INVERSION POINT ....................................................................................... 164

7.1 ABSTRACT .............................................................................................................. 165

7.2 INTRODUCTION .................................................................................................... 166

7.3 DEVELOPMENT OF EMULSION STABILITY MODEL SOLUTION ........... 172

7.4 MATERIALS AND METHODOLOGIES ............................................................. 175

7.4.1 Materials ............................................................................................................ 175

7.4.2 NAC Modeling Methodology ............................................................................ 175

7.4.3 Density and Viscosity Measurements ................................................................ 178

7.5 RESULTS .................................................................................................................. 178

7.5.1 Experimental and Predicted SDHS-Toluene-Water Microemulsion Properties 178

7.5.2 Predicted Droplet Size of 0.1 M SDHS-Toluene-Water Emulsions ................. 180

7.5.3 Predicted Stability of 0.1 M SDHS-Toluene-Water Emulsions ........................ 181

7.5.4 Effect of Surfactant Concentration on the Predicted Stability of SDHS-Toluene-

Water Emulsions ................................................................................................ 184

7.5.5 Effect of Temperature on the Predicted Stability of SDHS-Toluene-Water

Emulsions .......................................................................................................... 185

7.6 DISCUSSIONS .......................................................................................................... 187

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7.6.1 Prediction of the Droplet Size and Stability of SDHS-Toluene-Water Emulsions187

7.6.2 Prediction of the Emulsion Droplet Size of Lin et al......................................... 190

7.6.3 NAC Prediction of Initial Emulsion Droplet Size and Stability for the Systems of

Binks et al. with AOT and 0.65 cSt Silicone Oil ............................................... 192

7.6.4 NAC Prediction of Emulsion Stability for the Systems of Kabalnov et al. with

C12E5 and Octane ............................................................................................... 194

7.6.5 NAC Prediction of Emulsion Stability for the Systems of Salager et al. with

SDS, Pentanol, and Kerosene ............................................................................ 195

7.7 CONCLUSIONS ....................................................................................................... 196

7.8 REFERENCES ......................................................................................................... 197

CHAPTER 8: REVIEW OF THE FORMATION AND STABILITY OF BITUMEN

EMULSIONS ........................................................................................................................... 204

CHAPTER 9: CONCLUSIONS ............................................................................................. 213

CHAPTER 10: RECOMMENDATIONS.............................................................................. 218

10.1 RECOMMENDATIONS ......................................................................................... 219

10.2 REFERENCES ......................................................................................................... 221

APPENDIX 1: THE REVISED NET CURVATURE .......................................................... 223

APPENDIX 2: DISJOINING PRESSURE EQUATIONS ................................................... 227

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LIST OF TABLES

Table 2.1: Compositions (wt%) of tested heptol (H)-bitumen (B)-water (W) systems. ............. 17

Table 2.2: Material balance closure for systems prepared with 50 wt% water (W) at various

heptol 80/20 (H) to bitumen (B) dilution ratios and 25°C. .......................................................... 24

Table 3.1: Relative asphaltene and NA compositions in formulations tested. ........................... 62

Table 3.2: Measurements of ε for mixed asphaltene and NA films............................................ 62

Table 4.1: Calculated EACN of asphalt, bitumen, naphthalene, and deasphalted bitumen ....... 88

Table 4.2: Cc (* a

cC ) of NAs, NaNs, and asphaltene aggregates. ................................................ 91

Table 6.1: Reference viscosities () and refractive indices (n). ............................................... 144

Table 6.2: HLD and NAC modeling parameters for SDHS-toluene-water µEs. ...................... 145

Table 6.3: HLD and NAC modeling parameters for the non-ionic µEs of Leaver and Olsson28

.

................................................................................................................................................... 145

Table 7.1: Required HLD and NAC parameters for predicting the properties of the emulsions of

Kiran et al. (Chapter 5), Kabalnov et al., Salager et al., Binks et al., and Lin et al.16,18,19,20

. .... 177

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LIST OF FIGURES

Figure 1.1: Number of publications per year for emulsions across various industrial fields of

interest............................................................................................................................................ 2

Figure 1.2: Commercial hot water flotation process used in oil sands operations. ...................... 4

Figure 2.1: Calibration of asphaltene absorbance at a wavelength of 450 nm and as a function

of the bitumen to toluene ratio. .................................................................................................... 20

Figure 2.2: (i) Optical, (ii) cross-polarized, and (iii) fluorescent micrographs of samples of the

separated oil phase, aqueous phase, and rag layer for a system containing 68.2 wt% heptol

80/20, 6.8 wt% bitumen, and 25 wt% water at 25°C. .................................................................. 23

Figure 2.3: (a) Oil and (b) asphaltene losses to the rag layer as a function of the mixing time for

systems prepared with either 25 wt% or 75 wt% water and the balance oil with a heptol 80/20 to

bitumen dilution ratio of 3 at 25°C. ............................................................................................. 25

Figure 2.4: Oil losses to the rag layer as a function of the g force × time for systems prepared

with (a) 12.5 wt% heptol 80/20, 12.5 wt% bitumen, and 75 wt% water, (b) 20 wt% heptol 80/20,

5 wt% bitumen, and 75 wt% water, (c) 37.5 wt% heptol 80/20, 37.5 wt% bitumen, and 25 wt%

water, and (d) 60 wt% heptol 80/20, 15 wt% bitumen, and 25 wt% water. ................................ 26

Figure 2.5: Oil and asphaltene losses to the rag layer for systems prepared with (a and b) heptol

80/20 at 25°C, (c and d) heptol 80/20 at 80°C, and (e and f) heptol 50/50 at 25°C. ................... 27

Figure 2.6: Ternary phase diagrams for systems prepared with (a) heptol 80/20 at 25°C, (b)

heptol 80/20 at 80°C, and (c) heptol 50/50 at 25°C..................................................................... 29

Figure 2.7: (a) E

dd , (b) sample E

id distribution at 50 wt% water, and (c) normalized Aow of

heptol 80/20-diluted bitumen droplets at 25°C. ........................................................................... 33

Figure 2.8: Morphology of emulsions for systems prepared with (a) 10 wt% heptol 80/20, 15

wt% bitumen, and 75 wt% water, (b) 12.5 wt% heptol 80/20, 12.5 wt% bitumen, and 75 wt%

xiii

water, (c) 10 wt% heptol 50/50, 15 wt% bitumen, and 75 wt% water, and (d) 12.5 wt% heptol

50/50, 12.5 wt% bitumen, and 75 wt% water. ............................................................................. 37

Figure 2.9: The measured γow as a function of solvent-bitumen-water ratios for systems

prepared with (a) heptol 80/20 at 25°C, (b) heptol 80/20 at 80°C, and (c) heptol 50/50 at 25°C.

..................................................................................................................................................... 38

Figure 2.10: The measured γow (against water) of asphaltenes diluted in heptol 80/20 and heptol

50/50 at 25°C. .............................................................................................................................. 39

Figure 2.11: Correlations of oil and asphaltene losses to the rag layer for systems prepared with

heptol 80/20 at (a) 25°C and (b) 80°C. ........................................................................................ 41

Figure 3.1: Oil and asphaltene losses to the rag layer for 0 wt% (baseline) and 3 wt% NaN

systems prepared at pH 7.5 with (a and b) heptol 80/20 at 25°C, (c and d) heptol 80/20 at 80°C,

and (e and f) heptol 50/50 at 25°C. .............................................................................................. 57

Figure 3.2: Oil losses to the rag layer for 0 wt% (baseline) and 3 wt% NA systems as a function

of the heptol 80/20 to bitumen dilution ratio and water content. All systems are evaluated at pH

4 and 25°C. .................................................................................................................................. 58

Figure 3.3: Fluorescent micrographs of the effect of (a) NaNs (pH 7.5) and (b) NAs (pH 4) on

the morphology of heptol 80/20-bitumen-water rag layers. (c) Cross-polarized images of rag

layers containing NAs at pH 4. .................................................................................................... 59

Figure 3.4: Baseline γow isotherms as well as those for (a) NaN systems at pH 7.5 and (b) NA

systems at pH 4 as a function of the heptol 80/20 to bitumen dilution ratio. The temperature is

maintained at 25°C. ..................................................................................................................... 60

Figure 3.5: πp-Asp isotherms of asphaltene (A) and NA surfactant mixtures. ............................ 61

Figure 3.6: γow of bitumen diluted with heptol 80/20 versus the added NaN concentration to the

aqueous phase at pH 7.5 and 25°C. The included fluorescent micrographs show a transition in

the rag layer morphology from w/o to o/w at the CMC of NaNs (~1 wt%). ............................... 64

xiv

Figure 3.7: (a) Bilayer model proposed by Wu and Czarnecki for the interfacial co-adsorption

of asphaltenes and NaNs at the oil-water interface1. (b) Co-adsorption mechanisms proposed by

Varadaraj and Brons for asphaltenes and NAs at the oil-water interface include (i) mixed

monolayers and (ii) mixed aggregates3........................................................................................ 65

Figure 4.1: The phase behavior and corresponding γow of SDHS-oil (20 wt% naphthalene and

80 wt% toluene)-water µEs as a function of the salinity. ............................................................ 77

Figure 4.2: µE phase behavior transitions as a function of the salinity for a (a) 30 wt% bitumen

and 70 wt% toluene oil phase mixture and (b) 20 mol% NAs and 80 mol% SDHS surfactant

mixture. ........................................................................................................................................ 83

Figure 4.3: The measured γow as a function of the salinity for (a) bitumen and toluene, (b)

asphalt and toluene, (c) naphthalene and toluene, and (d) deasphalted bitumen and toluene oil

phase mixtures. ............................................................................................................................ 84

Figure 4.4: Experimented and modeled shifts of S* for µEs composed of hexadecane and

toluene oil phase mixtures and 0.1 M of a 35 mol% SDHS and 65 mol% AOT surfactant

mixture. ........................................................................................................................................ 87

Figure 4.5: The measured γow as a function of the salinity for (a) NAs and SDHS, (b) NaNs and

SDHS, and (c) asphaltenes and SDHS surfactant mixtures at a total concentration of 0.1 M. ... 89

Figure 4.6: Salinity shift of (a) NAs and SDHS and (b) NaNs and SDHS surfactant mixtures as

a function of the test surfactant mole fraction within the 0.1 M surfactant mixture. .................. 90

Figure 4.7: (a) Asphaltene partitioning and (b) the shift in S* for optimum µEs formulated with

an asphaltenes and SDHS surfactant mixture. ............................................................................. 92

Figure 5.1: A Type IType IIIType II phase behavior scan for 0.1 M SDHS-toluene-water

µEs at 25°C. ............................................................................................................................... 102

Figure 5.2: Emulsion aggregation, settling, film thinning, and coalescence processes. .......... 103

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Figure 5.3: (a) Evolution of the observable phase separation stages versus time in a test tube

for emulsions produced in Type II µEs. (b) Normalized displacement of the µE and excess

aqueous phase separation fronts versus time and the extrapolated ap, dp, and cp time periods. 104

Figure 5.4: In-house multipoint turbidimeter used to track the net displacement of µE and

excess phase separation fronts at a high time resolution. .......................................................... 111

Figure 5.5: Measured (a) interfacial tension (γow) and (b) emulsion droplet diameter ( E

dd )

profiles for baseline (0.1 M SDHS) emulsions at 25°C. ........................................................... 113

Figure 5.6: Example of sample time-lapse images and ΔsµE(t) and Δsw(t) profiles versus time

for a Type II baseline (0.1 M SDHS) emulsion at HLD=0.9 and 25°C. ................................... 114

Figure 5.7: Net rate of advance of the µE and excess phase separation fronts as a function of the

HLD for baseline (0.1 M SDHS) emulsions at 25°C. ............................................................... 116

Figure 5.8: Compiled ap, dp, cp, and ep (ap+dp+cp) time periods for emulsions at (a) 0.01 M, (b)

0.1 M, and (c) 0.3 M SDHS and 25°C. ...................................................................................... 117

Figure 5.9: Compiled ap, dp, cp, and ep (ap+dp+cp) time periods for baseline (0.1 M SDHS)

emulsions at (a) T=7°C, (b) T=15°C, (c) T=35°C, and (d) T=44°C. ........................................ 118

Figure 5.10: (a) Sample W* fittings for the aggregation, drainage, and coalescence of Type I

baseline (0.1 M SDHS) emulsion at HLD=-1 and (b) their values as a function of the HLD at

298 K. ........................................................................................................................................ 119

Figure 5.11: Dependence of the ap time period for baseline (0.1 M SDHS) emulsions at 25°C on

E

dd . The solid line represents the fit of tdiff from Equation 9 (R2=0.84).................................... 121

Figure 5.12: Comparison of the originally measured E

dd for baseline (0.1 M SDHS) emulsion

droplets at 25°C after mixing and the estimated E

appdd , required for settling (Equation 10) and

film thinning (Equation 11) as a function of the HLD. ............................................................. 123

Figure 5.13: General correlation between all of the fitted W* and the measured γow for baseline

(0.1 M SDHS) emulsions at 25°C. ............................................................................................ 125

xvi

Figure 6.1: Experimental and predicted (a) E

IoV , and E

IIwV , as well as (b) E

d

as a function of

the HLD for SDHS-toluene-water µEs. ..................................................................................... 147

Figure 6.2: Experimental and NAC modeled (for dilute liquid spheres (Equation 2) and

concentrated hard spheres (Equation 3)) ηµE of SDHS-toluene-water µEs as a function of the

HLD. .......................................................................................................................................... 148

Figure 6.3: Predicted (a) E

dr and E

dl as well as (b) total aspect ratio ( E

dl /2 E

dr ) as a function

of the HLD for SDHS-toluene-water µEs using the NAC model. ............................................ 149

Figure 6.4: Predicted SANS profiles of SDHS-toluene-water µEs formulated at (a) HLD=-0.9,

(b) HLD=-0.6, (c) HLD=-0.4, (d) HLD=0.3, (e) HLD=0.4, and (f) HLD=0.6 using the NAC

model. ........................................................................................................................................ 151

Figure 6.5: Experimental E

hr and predicted E

hr

max, for Type I and Type II SDHS-toluene-water

µEs as a function of the HLD. ................................................................................................... 152

Figure 6.6: Experimental and NAC modeled (for dilute rigid rods (Equation 6) and prolate

ellipsoids (Equation 8)) ηµE of SDHS-toluene-water µEs as a function of the HLD. ............... 154

Figure 6.7: Experimental and predicted E / E

c

ratio for the (a) C12E4-hexadecane-water and

(b) C12E5-cyclohexane and hexadecane-water µEs of Leaver and Olsson28

. ............................ 155

Figure 7.1: Overview of the simplified CCS demulsification mechanism used to model

emulsion stability. ...................................................................................................................... 170

Figure 7.2: Cross-section of coalescing emulsion droplets showing the formation of the

nucleating hole and its characteristic dimensions. ..................................................................... 171

Figure 7.3: General algorithm used to predict the size and stability of emulsions................... 172

Figure 7.4: Experimental and NAC modeled (a) γow, (b) ρµE, and (c) ηµE of SDHS-toluene-water

µEs. ............................................................................................................................................ 180

xvii

Figure 7.5: Predicted E

dr of 0.1 M SDHS-toluene-water emulsions at 298 K using the NAC

model. Experimental data are obtained from the work presented in Chapter 5. ....................... 181

Figure 7.6: Predicted (a) ut and (b) tH,crit of 0.1 M SDHS-toluene-water emulsions at 298 K

using the NAC model. ............................................................................................................... 182

Figure 7.7: Predicted Eac of 0.1 M SDHS-toluene-water emulsions at 298 K using the NAC

model. Experimental data are obtained from the work presented in Chapter 5. ....................... 183

Figure 7.8: Predicted stability (solid lines) of 0.1 M SDHS-toluene-water emulsions at room

temperature with E

d =0.5 (initial volume fraction) and 0.74 (closely packed hard spheres). All

experimental data is taken from Chapter 5. ............................................................................... 184

Figure 7.9: Predicted stability (solid lines) of (a) 0.01 M and (b) 0.3 M SDHS-toluene-water

emulsions at 298 K and E

d =0.5. Dashed line in (a) corresponds to predicted values using

tH,crit=0. All experimental data is taken from Chapter 5. ........................................................... 185

Figure 7.10: Predicted stability (solid lines) of 0.1 M SDHS-toluene-water emulsions at

E

d =0.5 and (a) 280 K, (b) 288 K, (c) 308 K, and (d) 317 K. All experimental data was taken

from Chapter 5. .......................................................................................................................... 186

Figure 7.11: Predicted emulsion droplet size (solid line) for the nonylphenol ethoxylate-mineral

oil-water systems at 294 K and as a function of the wt% of polyethylene oxide (PEO) in the

surfactant. See Table 7.1 for additional simulation conditions. The experimental data is taken

from Lin et al. ............................................................................................................................ 191

Figure 7.12: Predicted (a) interfacial tension and (b) emulsion stability (solid lines) for the

AOT-0.65 cSt silicone oil system. See Table 7.1 for additional simulation conditions. The

dashed line in (b) represents the predicted stability considering an additional contribution to

tH,crit of 15 nm. The experimental data is taken from Binks et al. .............................................. 193

Figure 7.13: Predicted emulsion stability (solid lines) for the system (a) C12E5–octane and (b)

SDS-pentanol-kerosene. See Table 7.1 for additional simulation conditions. The experimental

xviii

data for the systems in (a) and (b) is obtained from Kabalnov et al. and Salager et al.

respectively. ............................................................................................................................... 195

Figure 8.1: The HLD as a function of the heptol to bitumen dilution ratio for asphaltene-

stabilized rag layers prepared with heptol 80/20 and heptol 50/50 at 25°C………………...…206

Figure 8.2: The effect of 3 wt% sodium naphthenates (NaNs) on the HLD of bitumen emulsions

at an oil to water ratio of 1 to 1 and as a function of the heptol to bitumen dilution ratio…….207

Figure 8.3: The effect of 3 wt% naphthenic acids (NAs) on the HLD of bitumen emulsions as a

function of the heptol to bitumen dilution ratio………………………………………………..208

Figure 8.4: Experimental and predicted γow as a function of the HLD for o/w bitumen emulsions

stabilized by a mixture of asphaltenes and NaNs……………………….……………………..210

Figure 8.5: Experimental and predicted E

dd as a function of the HLD for o/w bitumen emulsion

droplets stabilized by a mixture of asphaltenes and NaNs…………………………………….211

Figure 8.6: Predicted stability of o/w bitumen emulsions stabilized by a mixture of asphaltenes

and NaNs as a function of the HLD……………………………………………………………212

xix

LIST OF SYMBOLS

α = 0.6 for turbulent flow and 1 for laminar

flow

= Shear rate

γow = Interfacial tension

δtail = Surfactant tail length

ɛ = Elasticity

ɛmix = Energy dissipation during mixing

ηµE = Viscosity of µE

E

c

= Viscosity of µE’s continuous phase

E

d

= Viscosity of µE’s dispersed phase

ηoil = Viscosity of oil phase

ηwater = Viscosity of aqueous phase

θ = Scattering angle

θc = Contact angle

κ = Bending modulus

= Saddle-splay modulus

~ = Dimensionless chemical potential

µE = Microemulsion

πp = Surface pressure

Δρ = Density difference

ρµE = Density of µE phase

E

c = Density of emulsion’s continuous

phase

E

d = Density of emulsified phase

ρoil = Density of oil phase

ρwater = Density of aqueous phase

σ = Surface tension in presence of surfactant

~ = Dimensionless excess free energy

σo = Surface tension of pure solvent

τ(t) = Measured turbidity

E

d

= Volume fraction of µE droplets

E

d

max, = Maximum E

d

E

wo

= Volume fraction of oil (water) in the

µE phase

E

d = Volume fraction of emulsion droplets

ϕ(A) = Co-surfactant effect on the HLD of

nonionic surfactants

ω = Rotational velocity

E

ella = Major axis of µE ellipsoids

ap = Aggregation time period

ai = Surface area per molecule of surfactant

species i

aT = Pre-factor of temperature effect on the

HLD of ionic surfactants

E

dA = Surface area of emulsion droplet

AH = Effective Hamaker constant

As = Interfacial area of surfactant

Asp = Area of spreading phase

Aow = Area of oil-water interface

ACN = Alkane carbon number

AOT = Sodium dioctyl sulfosuccinate

A(q) = Scattering amplitude

b = Pre-factor of salt effect on the HLD of

nonionic surfactants

∆b = Scattering length difference between

the µE’s continuous and dispersed phases

xx

E

ellb = Minor axis of µE ellipsoids

cµc = Critical µE concentration

cg = Concentration of µE rigid rods

cp = Coalescence time period

csi = Concentration of surfactant species i

C1 = Fitting constant (~0.5)

C2 = Fitting constant (~0.25)

Cc = Characteristic curvature of ionic

surfactant

Cc,mix = Characteristic curvature of ionic

surfactant mixture

Cn = Characteristic curvature of nonionic

surfactant

a

cC = Apparent characteristic curvature of

ionic surfactant

ref

cC = Characteristic curvature of reference

ionic surfactant

test

cC = Characteristic curvature of test ionic

surfactant

CSA,avg = Surface area-averaged curvature

CAC = Critical aggregation concentration

CCS = Collision-coalescence-separation

mechanism

CMC = Critical micelle concentration

CTC = Critical transition concentration

E

dd = Average emulsion droplet diameter

E

appdd , = Apparent average emulsion droplet

diameter

dH = Hole diameter

E

id = Emulsion droplet of diameter i

dp = Drainage time period

E

rodsd = Diameter of µE rigid rods

D = Diffusion coefficient

Dimp = Impeller diameter

ep = Equilibrium time period

E1 = Energy of hole formation

E2 = Energy of hole opening

Er = Interfacial rigidity

Eac = Activation energy of coalescence

EACN = Equivalent alkane carbon number

EACNmix = EACN of oil mixture

f = Pre-exponential factor

fc = Fitting constant for E1

fratio = Ratio of film radius to Rc

f(ϕ) = Entropic function of E

d

f(A) = Co-surfactant effect on the HLD of

ionic surfactants

F = Force exerted by Du Noüy ring

E

iF = Number frequency of E

id

g = Gravitational acceleration constant (9.81

m/s2)

ΔhµE(t) = Net displacement of µE phase

separation front

ΔhµE,tot = Total displacement of µE phase

separation front at equilibrium

Δho(t) = Net displacement of excess oil

phase separation front

Δho,tot = Total displacement of excess oil

phase separation front at equilibrium

Δhw(t) = Net displacement of excess

aqueous phase separation front

xxi

Δhw,tot = Total displacement of excess

aqueous phase separation front at

equilibrium

Ha = Average curvature

Hn = Net curvature

H’n = Revised net curvature = Hn/2

HLD = Hydrophilic-lipophilic deviation

HLDmix = HLD of an oil or surfactant

mixture

HLDref = HLD of reference oil or surfactant

HLDtest = HLD of test oil or surfactant

I(q) = Scattering intensity

k = Debye length at salinity S

k* = Debye length at optimal salinity S

*

kB = Boltzmann constant (1.38×10-23

J/K)

K = Pre-factor of Nc,o effect on the HLD of

ionic and nonionic surfactants

l = Characteristic diffusion length

E

dl = Length of µE’s cylindrical neck

E

rodsl = Length of µE rigid rods

L = Length scaling parameter

Lp = Laser path length

nE = Number concentration of emulsion

droplets

nEo = Initial number concentration of

emulsion droplets

ntoluene = Refractive index of toluene

nwater = Refractive index of water

N = Number density of µE rigid rods

Nc,o = General descriptor for ACN/EACN

ref

ocN , = General descriptor for the

ACN/EACN of a reference oil

test

ocN , = General descriptor for the

ACN/EACN of a test oil

Nimp = Impeller speed of mixing

NAs = Naphthenic acids

NAC = Net-average curvature

NaN(s) = Sodium naphthenates

o/w = Oil-in-water

p2 = Polydispersity

P~

= Dimensionless osmotic pressure

Pcons = Power consumption

Pe = Peclét number

q = Scattering vector

E

dr = Radius of µE’s hemispherical end cap

E

dr = Average emulsion droplet radius

E

hr = Hydrodynamic radius of µE droplets

E

hr

max, = Maximum E

hr

rH = Hole radius

E

wor = Sphere-equivalent radius of the oil

(aqueous) phase

rring = Probe ring radius

rwire = Probe wire radius

Rc = Radius of curvature

Re = Reynolds number

ΔsµE(t) = Normalized displacement of µE

phase separation front

Δso(t) = Normalized displacement of excess

oil phase separation front

xxii

Δsw(t) = Normalized displacement of excess

aqueous phase separation front

S = Aqueous phase salinity

S* = Optimal S (where HLD=0)

*

mixS = S* with reference to HLDmix

SDHS = Sodium dihexyl sulfosuccinate

SDS = Sodium dodecyl sulfate

td = Falling ball’s time of descent

tdiff = Characteristic diffusion time

tH = Film thickness

tH,crit = Critical film thickness at which hole

formation occurs

tp = Phase separation time of interest

T = Temperature

∆T = Temperature difference from 298 K

uosc = Oscillatory free energy contribution

to ut

ut = Free energy penalty of expelling µE

droplet layers during film thinning

uvdW = van der Waals free energy

contribution to ut

Uc = Rate of emulsion droplet coalescence

vrel = Relative mixing velocity

vs = Settling rate

VµE = Volume of µE phase

E

dV = Volume of emulsion droplet

VE = Volume of total emulsion

Vmax = Maximum V(t)

pulloV , = Volume of oil pulled through

interface

E

woV = Volume of oil (water) in the µE

phase

E

woV = Volume of emulsified oil (water)

V(t) = Voltage

w = Width of expanded oil droplet

w/o = Water-in-oil

W* = Apparent activation energy

We = Weber number

xi = Mol (volume) fraction of a given

surfactant (oil)

xref = Mol (volume) fraction of a reference

surfactant (oil)

xtest = Mol (volume) fraction of a test

surfactant (oil)

1

CHAPTER 1:

THESIS OVERVIEW

2

1.1 OVERVIEW

Emulsions are thermodynamically unfavorable mixtures of surfactant-encapsulated

droplets (>0.4 µm) of one medium (oil or water) dispersed throughout the other1. According to

the query conducted using SciVerse Scopus in Figure 1.1, the study of emulsions has

commanded a growing amount of attention across various applications. While a large part of

this growth has been experienced in developing new and innovative paint and coating, foodstuff,

and pharmaceutical products, it is projected that the role of emulsions as an undesirable process

intermediate in the oil sands sector will soon become a more prominent area of focus. The

reason for such is that billions of dollars are being spent to help expand the bitumen (i.e. heavy

crude oil) production capacity of oil sands operations in north-eastern Alberta (Canada)2,3,4

. In

view of this investment, it is expected that the bitumen output of 1,000,000 barrels/day will at

least triple over the next decade.

Paints and Coatings

Lubricants

0

300

600

900

1200

1500

1950 1955 1960 1965 1970 1975 1980 1985 1990 1995 2000 2005 2010 2015

Year

# P

ub

lic

ati

on

s

Foodstuffs

Cosmetics

Pharmaceuticals

Crude Oil

Pesticides

Cleaning

Figure 1.1: Number of publications per year for emulsions across various industrial fields of

interest.

3

An overview of the commercial hot water flotation process used in oil sands operations

is outlined in Figure 1.25. In this process, mined oil sands are first added along with hot water to

a rotating horizontal tumbler where their fragmentation and dispersion facilitates the liberation

of bitumen from quartz sand particles. To further help separate bitumen from water in

downstream process units, this mixed slurry is treated with sodium hydroxide (NaOH) and

steam so that it exits the tumbler at a conditioned pH (8), temperature (80°C), and degree of

aeration (30%). The resulting effluent next passes through a series of vibrating screens to

remove all oversized solid particles and undissociated oil sand lumps prior to being diluted with

excess hot water and pumped to a primary separation vessel. Within this vessel, the aerated

bitumen droplets rise to the top and are removed as a primary froth whereas most of the fine

solids collect at the bottom and are removed as a concentrated tailings solution. The middlings

region is on the other hand subjected to additional aeration steps to produce a secondary froth

that separates out from a secondary separation vessel at a similar composition as that of the

primary froth (60% bitumen, 30% water, and 10% fine solids). As a result of its large viscosity

(10,000 cP at 25°C) and similar density to that of water, the combined froth is diluted with a

solvent (naphtha) prior to passing through an energy-intensive mechanical separation cycle

where formed oil-in-water (o/w) and water-in-oil (w/o) emulsion droplets are broken and a final

and cleaner bitumen product ready for upgrading into synthetic crude oil is produced. The

formation and stability of these emulsions as “rag layers” is a major operational issue and

therefore serves as the main focus of this thesis.

4

Mined Oil Sands

Tumbler

Vibrating Screens Solid

Rejects

Hot Water + NaOH

+ Steam

Excess

Hot Water

S

e

p

a

r

a

t

i

o

n

Tailings

MIddlingsAeration

S

e

p

a

r

a

t

i

o

n

Recycle

1 Froth

2 Froth

Mechanical

Separation Cycle

Upgrading

Figure 1.2: Commercial hot water flotation process used in oil sands operations.

The formation and stability of rag layers are governed by surfactant-like species found

naturally within bitumen. An example of one such component, which makes up close to 15% of

bitumen, is asphaltenes6. Asphaltenes may be classified as highly aromatic structures that

solubilize in aromatic hydrocarbons (e.g. benzene and toluene) and precipitate upon mixing with

paraffinic solvents (e.g. pentane and heptane)6,7

. The surface activity of asphaltenes varies

according to their solubility. As has been demonstrated by several previous researchers,

asphaltenes having undergone a monomer-to-aggregate transition under partially soluble

conditions show an enhanced affinity for the oil-water interface6,8,9

. The resulting adsorbed

asphaltene skins, which are characteristically lipophilic, promote the formation of w/o emulsion

droplets and act as a rigid energy barrier that they must overcome in order to coalesce10,11,12

.

Naphthenic amphiphiles, which make up no more than 4% of bitumen, are another class

of surfactants that notably impact rag layer properties13

. These species, which have a pKa of ~6,

may be categorized as either naphthenic acids (NAs) or naphthenate salts (e.g. sodium

5

naphthenates (NaNs)) depending on the pH of the system14

. Under acidic conditions, where the

pH is less than the pKa of naphthenic amphiphiles, NAs are the dominant surrogate. This term

encompasses all alkyl-substituted cycloaliphatic carboxylic acids (R-COOH) present within

bitumen15

. Like asphaltenes, NAs are oil-soluble and therefore promote the formation of w/o

emulsion droplets14,15

. Under alkaline conditions, where the pH is greater than the pKa of

naphthenic amphiphiles, NaNs are produced via the association of carboxylic acid anions

(COO-) in the oil phase with sodium cations (Na

+) dissolved in the aqueous phase

16. These

metallic soaps aid in emulsification by significantly reducing the oil-water interfacial tension17

.

Being water-soluble, they also induce a shift in the rag layer morphology from w/o to o/w18

. It

has been proposed that the principal mechanism by which naphthenic amphiphiles stabilize

emulsion droplets is by forming lamellar liquid crystals19,20

. These mesomorphic phases, which

are typically produced under concentrated surfactant regimes, spontaneously spread across the

oil-water interface and reduce its mobility and bending ability. Contributions of electrostatic

repulsion are also relevant for the case of NaNs21

.

A lot of work has evidently already gone into understanding the individual influence of

asphaltenes and naphthenic amphiphiles on rag layers. Despite these advancements, very few

efforts have been aimed at practically relating the observed findings to actual commercial hot

water flotation process variables. This highlighted gap is partially addressed in Chapter 2 by

studying the effects of solvent-bitumen-water ratios, solvent aromaticity, and temperature on

asphaltene-stabilized emulsions. A more complete picture emerges in Chapter 3 where the

added effect of naphthenic amphiphiles at different concentrations is tested. From this last set of

data, the theoretically proposed makeup of co-adsorbed asphaltene and naphthenic amphiphile

films is reviewed and new insights are offered.

6

An important lesson learned from Chapters 2 and 3 is that changes to the

physicochemical environment of formed w/o and o/w emulsions leads to changes in their

stability. In knowing so, the next logically posed question is how to predictively model this

relationship. The most widely used approach aimed at doing just so is the hydrophilic-lipophilic

balance (HLB) model22,23,24

. The general premise behind this model is that the experimental

emulsification behavior of a surfactant is related to the difference of its hydrophilic and

lipophilic group contributions. At HLB<10 (HLB>10), stable w/o (o/w) emulsion droplets are

formed. Closer to the balanced (or phase inversion) point (HLB=10), w/o and o/w emulsion

droplets phase separate almost instantaneously. The selection of a surfactant for a given

application is governed by the overall HLB requirement. A notable shortcoming of such is that

this parameter is based purely on the nature of the oil phase. Consequentially, the additional

effect of other physicochemical parameters is ignored.

The hydrophilic-lipophilic deviation (HLD) model has more recently been suggested as

a way of helping overcome the limitations of the HLB model25,26

. The HLD model was

originally devised to track the equilibrium phase behavior of microemulsions (µEs). By

scanning the salinity, oil phase/surfactant hydrophobicity, co-surfactant type/concentration,

and/or temperature to induce a HLD<0HLD=0HLD>0 transition, a shift from o/w µEs in

equilibrium with an excess oil phase (Type I)bicontinuous µEs in equilibrium with excess oil

and aqueous phases (Type III)w/o µEs in equilibrium with an excess aqueous phase (Type II)

is observed. It wasn’t realized until later that upon dispersing these excess phases as emulsion

droplets throughout their continuous µE phase that Type I and Type II stability maxima are

separated by a Type III stability minimum at the phase inversion point. To relate the formation

and stability of heavy crude oil emulsions (such as the above rag layers) using the HLD model,

7

the hydrophobicity of asphaltenic oils (bitumen, deasphalted bitumen, asphalt, and naphthalene)

as well as that of surface-active asphaltene aggregates and naphthenic amphiphiles is evaluated

in Chapter 4. A key finding is that bitumen and toluene share a similar hydrophobicity. This

finding is consistent with the fact that these oils are mutually soluble in one another. As a result

of this, toluene is instead used as the model oil phase of interest going forward because of its

simpler makeup and well-defined properties27,28

.

Despite the advanced usefulness of the HLD model compared to the HLB model in

predicting the stability of emulsions as a function of their formation, it can unfortunately only do

so qualitatively. A better grasp of how the HLD model relates to the kinetic processes common

to demulsification is required for a more quantitative prediction to be achieved. The objective of

Chapter 5 is to experimentally explore this relationship for anionic surfactant (sodium dihexyl

sulfosuccinate (SDHS))-toluene-water emulsions at different surfactant concentrations and

temperatures. In doing so, the total demulsification time is categorized into aggregation,

drainage (settling + film thinning), and coalescence time periods. The advantage of studying

these other emulsions is that it is suspected that the net-average curvature (NAC) model can be

employed to bridge together their HLD and corresponding experimental stability. This suspicion

is based on the fact that the NAC model has been used before to successfully predict numerous

other SDHS-based µE phase behavior properties29

.

The next chapters are thus aimed at predicting the stability of SDHS-toluene-water

emulsions with the aid of the NAC model. To do so, the NAC model is first used to predict the

shape of µE droplets in Chapter 6. These predicted µE droplet shapes are validated via Dr.

Edgar J. Acosta’s efforts to predict their small angle neutron scattering (SANS) profiles. From

these predicted µE droplet shapes, estimates of the continuous µE phase viscosity (ηµE) are

8

obtained. By being able to model ηµE on top of other continuous µE phase properties, emulsion

stability is finally predicted in Chapter 7 using an Arrhenius-like expression. The pre-

exponential part of this fundamental expression reflects the collision frequency of emulsion

droplets whereas the exponential part is indicative of their activation energy barrier. The

applicability of this modeling approach to the stability of other emulsions already published in

the literature is further presented.

In Chapter 8, the HLD and NAC models are both relayed back to the original formation

and stability of rag layers. The calculated HLD of these rag layers is compared against their

experimental morphology at different solvent-bitumen-water ratios, asphaltene to naphthenic

amphiphile ratios, solvent aromaticities, and temperatures. The NAC model is more specifically

applied to rag layers stabilized by a mixture of asphaltenes and NaNs. A discussion of the

experienced shortcomings and challenges that still lay ahead before a complete prediction of rag

layer stability can be obtained is provided towards the end of this chapter.

The conclusions chapter (Chapter 9) summarizes all of the major contributions that can

be drawn from this thesis. A final recommendations chapter (Chapter 10) is also included that

highlights proposed areas of related research that should be carried out to further substantiate

some of the above conclusions.

1.2 REFERENCES

1 Israelachvili, J. The Science and Applications of Emulsions-An Overview. Colloids Surf., A

1994, 91, 1-8.

9

2 Kasoff, M.J. East Meets West in the Canadian Oil Sands. Am. Rev. Can. Stud. 2007, 37, 177-

183.

3 Scales, M.; Werniuk, J. Oil Sands Changing the Face of Canada. Can. Min. J. 2008, 129, 16-

25.

4 Owen, N.A.; Inderwildi, O.R.; King, D.A. The Status of Conventional World Oil Reserves-

Hype or Cause for Concern? Energy Policy 2010, 38, 4743-4749.

5 Shaw, R.C.; Schramm, L.L.; Czarnecki, J. Suspensions in the Hot Water Flotation Process for

Canadian Oil Sands. Adv. Chem. Ser. 1996, 251, 639-675.

6 Yang, X.; Hamza, H.; Czarnecki, J. Investigation of Subfractions of Athabasca Asphaltenes

and Their Role in Emulsion Stability. Energy Fuels 2004, 18, 770-777.

7 Taylor, S.D.; Czarnecki, J.; Masliyah, J. Disjoining Pressure Isotherms of Water-in-Bitumen

Emulsion Films. J. Colloid Interface Sci. 2002, 252, 149-160.

8 Yarranton, H.W.; Hussein, H.; Masliyah, J. Water-in-Hydrocarbon Emulsions Stabilized by

Asphaltenes at Low Concentrations. J. Colloid Interface Sci. 2000, 228, 52-63.

9 Rondón, M.; Pereira, J.C.; Bouriat, P.; Graciaa, A.; Lachaise, J.; Salager, J.-L. Breaking of

Water-in-Crude Oil Emulsions. 2. Influence of Asphaltene Concentration and Diluent Nature on

Demulsifier Action. Energy Fuels 2008, 22, 702-707.

10

10 Sztukowski, D.M.; Jafari, M.; Alboudwarej, H.; Yarranton, H.W. Asphaltene Self-Association

and Water-in-Hydrocarbon Emulsions. J. Colloid Interface Sci. 2003, 265, 179-186.

11 Asekomhe, S.O.; Chiang, R.; Masliyah, J.H.; Elliott, J.A.W. Some Observations on the

Concentration Behavior of a Water-in-Oil Drop with Attached Solids. Ind. Eng. Chem. Res.

2005, 44, 1241-1249.

12 Jiang, T.; Hirasaki, G.; Miller, C.; Moran, K.; Fleury, M.; Diluted Bitumen Water-in-Oil

Emulsion Stability and Characterization by Nuclear Magnetic Resonance (NMR)

Measurements. Energy Fuels 2007, 21, 1325-1336.

13 Barrow, M.P.; Headley, J.V.; Peru, K.M.; Derrick, P.J. Data Visualization of the

Characterization of Naphthenic Acids within Petroleum Samples. Energy Fuels 2009, 23, 2592-

2599.

14 Havre, T.E.; Ese, M.-H.; Sjöblom, J.; Blokhus, A.M. Langmuir Films of Naphthenic Acids at

Different pH and Electrolyte Concentrations. Colloid Polym. Sci. 2002, 280, 647-652.

15 Havre, T.E.; Sjöblom, J. Emulsion Stabilization by Means of Combined Surfactant Multilayer

(D-Phase) and Asphaltene Particles. Colloids Surf., A 2003, 228, 131-142.

16 Brandal, Ø.; Sjöblom, J. Interfacial Behavior of Naphthenic Acids and Multivalent Cations in

Systems with Oil and Water. II: Formation and Stability of Metal Naphthenate Films at Oil-

Water Interfaces. J. Dispersion Sci. Technol. 2005, 26, 53-58.

11

17 Havre, T.E.; Sjöblom, J.; Vindstad, J.E. Oil/Water-Partitioning and Interfacial Behavior of

Naphthenic Acids. J. Dispersion Sci. Technol. 2003, 24, 789-801.

18 Acosta, E.J.; Yuan, J.S.; Bhakta, A.S. The Characteristic Curvature of Ionic Surfactants. J.

Surfactants Deterg. 2008, 11, 145-158.

19 Horváth-Szabó, G.; Czarnecki, J.; Masliyah, J.H. Sandwich Structures at Oil-Water Interfaces

under Alkaline Conditions. J. Colloid Interface Sci. 2002, 253, 427-434.

20 Häger, M.; Ese, M.-H.; Sjöblom, J. Emulsion Inversion in an Oil-Surfactant-Water System

Based on Model Naphthenic Acids under Alkaline Conditions. J. Dispersion Sci. Technol. 2005,

26, 673-682.

21 Arla, D.; Sinquin, A.; Palermo, T.; Hurteyent, C.; Graciaa, A.; Dicharry, C. Influence of pH

and Water Content on the Type and Stability of Acidic Crude Oil Emulsions. Energy Fuels

2007, 21, 1337-1342.

22 Griffin, W.C. Classification of Surface Active Agents by HLB. J. Soc. Cosmet. Chem. 1949,

1, 311-326.

23 Kruglyakov, P.M. Hydrophile-Lipophile Balance of Surfactants and Solid Particles-

Physicochemical Aspects and Applications, 1st ed.; Elsevier Science B.V.: Amsterdam, 2000.

24 Rosen, M.J. Surfactants and Interfacial Phenomena, 3

rd ed.; John Wiley & Sons, Inc.: New

Jersey, 2004.

12

25 Salager, J.L.; Antón, R.; Briceno, M.I.; Choplin, L.; Márquez, L.; Pizzino, A.; Rodriguez,

M.P. The Emergence of Formulation Engineering in Emulsion Making-Transferring Know-How

from Research Laboratory to Plant. Polym. Int. 2003, 52, 471-478.

26 Rondón, M.; Bouriat, P.; Lachaise, J.; Salager, J.-L. Breaking of Water-in-Crude Oil

Emulsions. 1. Physicochemical Phenomenology of Demulsifier Action. Energy Fuels 2006, 20,

1600-1604.

27 Scott, D.W.; Guthrie, G.B.; Messerly, J.F.; Todd, S.S.; Berg, W.T.; Hossenlopp, I.A.;

Mccullough, J.P. Toluene: Thermodynamic Properties, Molecular Vibrations, and Internal

Rotation. J. Phys. Chem. 1962, 66, 911-914.

28 Linde, B.B.J.; Skrodzka, E.B.; Lezhnev, N.B. Vibrational Relaxation in Several Derivatives

of Benzene. Int. J. Thermophys. 2012, 33, 664-679.

29 Acosta, E.; Szekeres, E.; Sabatini, D.A.; Harwell, J.H. Net-Average Curvature Model for

Solubilization and Supersolubilization in Surfactant Microemulsions. Langmuir 2003, 19, 186-

195.

13

CHAPTER 2:

STUDY OF SOLVENT-BITUMEN-WATER RAG LAYERS

This chapter is derived from the following published manuscript:

Kiran, S.K.; Acosta, E.J.; Moran, K. Study of Solvent-Bitumen-Water Rag Layers. Energy

Fuels 2009, 23, 3139-3149.

14

2.1 ABSTRACT

In this chapter, the stability of water-in-oil rag layers was evaluated as a function of

solvent-bitumen-water ratios, solvent aromaticity, and temperature using a combination of

microscopic imaging and spectroscopic techniques. With the aid of these techniques, it was

possible to obtain via material balances an estimate of the amount of oil, water, and asphaltenes

in the rag layer and excess phases. It was observed that when bitumen was diluted with a

paraffinic (or poor) solvent, such as heptol 80/20 (80 vol% heptane and 20 vol% toluene),

asphaltenes in solution tended to preferentially adsorb/segregate at the exposed oil-water

interface and stabilize rag layers. Diluting similar systems with a more aromatic solvent (heptol

50/50, 50 vol% heptane and 50 vol% toluene) reduced the surface activity of asphaltenes and

hence their ability to stabilize rag layers. It was further observed that an increase in temperature

minimized rag layer stability. This effect was explained by the lower viscosity of the oil, which

resulted in its improved drainage from the rag layer.

2.2 INTRODUCTION

A number of studies aimed at illustrating the role that asphaltenes play in stabilizing

water-in-oil (w/o) rag layers have already been published in the literature. Yan et al. showed that

the emulsifying capacity of deasphalted bitumen is minimal1. Yarranton et al. later showed that

asphaltene molecules at a low concentration (≤2 kg/m3) act as surfactant monomers capable of

stabilizing w/o emulsions2. Rondón et al. further demonstrated that a critical aggregation

concentration of asphaltenes exists (~1000 ppm in their system) where the oil-water interface is

saturated3. These researchers additionally hypothesized that the observed reduction in the rate of

demulsification in the presence of excess asphaltenes (beyond the critical aggregation

15

concentration) is due to the formation of micellar aggregates as well as a type of gel phase

adjacent to the interfacial layer. By deflating such asphaltene-stabilized emulsions using

micropipette techniques, Yeung et al. and Tsamantakis et al. proved that excess asphaltenes

adsorb at the oil-water interface as rigid multilayer skins4,5

. Additional results acquired by Yang

et al. showed that asphaltenes are most able to form rigid skins under partially soluble

conditions6. Wu finally revealed that the extent to which asphaltene skins form is enhanced at an

increased oil-water interfacial area (Aow)7. This was confirmed by Gelin et al. who analyzed the

effect of water concentration on the behavior of asphaltenes within reservoir hydrocarbon

fluids8. Obtained results indicate that the concentration of emulsified water significantly affects

the amount and physical properties of asphaltenes lost to the rag layer. While it is not typical for

asphaltenes to adsorb as liquid crystals, there is some evidence of their existence9,10,11

.

Despite these studies with model systems that point to the surface activity of asphaltenes

in stabilizing w/o rag layers, there is a lack of literature characterizing the effectiveness of heavy

crude oil processing variables on the stabilization potential of produced asphaltene skins. One of

the goals of this chapter is to introduce a batch emulsification-separation protocol and method of

analysis to study solvent-bitumen-water rag layers. The second goal of this chapter is to

subsequently study the effect that variables such as solvent-bitumen-water ratios, solvent

aromaticity, and temperature have on the surface activity of asphaltenes and the stability of

formed rag layers. These variables are evaluated by measuring the fraction of oil and asphaltene

losses to the rag layer and the interfacial tension (γow) of all the systems considered.

16

2.3 MATERIALS AND METHODS

2.3.1 Materials

All chemicals were used as supplied: anhydrous toluene (product #244511, 99.8%) was

purchased from Sigma-Aldrich Corp. (Oakville, ON, Canada), reagent grade heptane (product

#5400-1-10) was purchased from Caledon Laboratory Chemicals (Georgetown, ON, Canada),

and coker feed bitumen (16.3% saturates, 39.8% aromatics, 28.5% resins, and 14.7%

asphaltenes according to SARA analysis) was donated by Syncrude Canada Ltd. (Edmonton,

AB, Canada)12

. A salt water solution was prepared by dissolving 25 mmol/L NaCl, 15 mmol/L

NaHCO3, 2 mmol/L Na2SO4, 0.3 mmol/L CaCl2, and 0.3 mmol/L MgCl2 in deionized water to

simulate typical water compositions in oil field operations13

. The pH of this salt water solution

was ~7.5.

2.3.2 Formulation of Rag Layers

The oil phase was prepared by mixing heptol 80/20 (80 vol% heptane and 20 vol%

toluene) and bitumen together at various ratios in individual glass jars overnight using a wrist-

action shaker with a stroke length and frequency of 1.5 inches and 180 strokes/minute

respectively. These mixing parameters were used to ensure that a homogeneous oil phase was

obtained. The diluted bitumen solutions were next added to prescribed amounts of the aqueous

phase in 15 mL glass centrifuge tubes. These formulations were mixed (150 W VWR analog

vortex mixer) at 3200 rpm for 2 minutes and then centrifuged (IEC clinical centrifuge) at 500 g

for 1 minute at 25°C. The development of these emulsification and separation protocols will be

discussed later. The compositions of solvent-bitumen-water systems tested in this work are

presented in Table 2.1.

17

Table 2.1: Compositions (wt%) of tested heptol (H)-bitumen (B)-water (W) systems.

Heptol/Bitumen

Ratio 1/1.5 1/1 3/1 4/1 10/1

W H B H B H B H B H B

75 10 15 12.5 12.5 18.8 6.3 20 5 22.7 2.3

50 20 30 25 25 37.5 12.5 40 10 45.5 4.5

25 30 45 37.5 37.5 56.3 18.8 60 15 68.2 6.8

9.1 36.4 54.5 45.5 45.5 68.2 22.7 72.7 18.2 82.6 8.3

0 40 60 50 50 75 25 80 20 90.9 9.1

The same systems described in Table 2.1 were studied at 80°C by placing the test tubes

in a hot water bath prior to mixing and centrifuging. Another set of phase behavior studies was

also carried out using heptol 50/50 (50 vol% heptane and 50 vol% toluene) as the solvent.

2.3.3 Microscopy

The oil phase, aqueous phase, and rag layer produced in all the formulations after mixing

and centrifuging were sampled and analyzed according to slight variations in the procedure

developed by Varadaraj et al.14

. To sample the excess oil phase (top phase), a small volume of

oil was slowly extracted using a 5 mL pipet. When sampling the rag layer (middle phase) and

aqueous phase (bottom phase), a small positive pressure was applied during the insertion and

extraction of the pipet to ensure that no additional phases were sampled along with the phase of

interest. Three distinct sampling points were used for each of the phases to guarantee that the

sample was representative. All extracted samples were imaged using an Olympus C-7070 wide

zoom digital camera at 4× magnification mounted on top of an Olympus BX-51 microscope set

at 50× magnification. Three different microscope configurations were used: optical (transmitted)

light (to differentiate between solid and liquid phases), cross-polarized light (to detect liquid

crystals), and fluorescent light (to differentiate between oil (green) and aqueous (black) phases).

18

2.3.4 Material Balances

Because of the variable composition and density of the rag layer in all the systems

studied, material balances were performed by tracking the volume of each formulation

component in all the phases produced after mixing and centrifuging. The volume of each phase

produced was calculated by measuring the relative heights of the separated oil phase, aqueous

phase, and rag layer. The separated oil phase was considered a recoverable and useful product in

which only minor traces of water were dispersed throughout. This was confirmed by measuring

its water content using Karl Fischer titration (<0.2 vol% for all systems) and the microscopy

techniques in Section 2.3.3. The purity of the separated aqueous phase was confirmed using

only micrographs to track suspended or dissolved oil. To determine the volume of oil and water

within the rag layer, their volume fractions were first calculated by analyzing the captured

fluorescent micrographs using Scion Image Software15

:

1) A drop of the rag layer sample (20 µL) was deposited on top of a glass slide and then

carefully covered with a coverslip. This forced the rag layer sample to evenly spread and

form a thin film of 10 µm. This film thickness was determined by dividing the volume of

the deposited rag layer sample by the area of the glass slide-coverslip assembly.

2) The threshold used to transform a fluorescent micrograph into a binary map of clear (oil)

and dark (water) areas was adjusted such that the boundary between the oil and water areas

was located at the middle of their fuzzy interface. This fuzzy interface was caused by the

curvature of the walls of the squeezed water droplets. Non-squeezed water droplets with a

diameter <10 µm were not commonly observed.

19

3) The area of the rag layer image was calculated along with that of the oil and water phases

using measurement tools within the software package.

4) Knowing the area of both the oil and water phases (pixels2) as well as the area of the entire

image (pixels2), the fractional areas (and hence volume fractions) of oil and water within

the rag layer were calculated.

From these calculated volume fractions, the volumes of oil and water within the rag

layer were determined. The measured amount of oil and water in the excess phases and within

the rag layer in a given test tube agreed within ±1 vol% of the known volume of oil and water

initially added (this is the material balance closure). Experiments with larger deviations were

rejected and repeated. The standard deviation of oil losses to the rag layer in replicate systems

was ≤5 vol%. It needs to be clarified that the above material balance approach is only useful in

exploring oil-water phase separation trends. Its accuracy has yet to be compared against more

established methods (i.e. gravimetric techniques).

2.3.5 Asphaltene Losses

Yang et al. developed a UV-vis spectroscopic technique to analyze the concentration of

asphaltenes in heavy crude oils6. This technique was adapted here to quantify asphaltene losses

to the rag layer (as a fraction of their initial concentration in the oil phase). These losses include

interfacially adsorbed/segregated asphaltenes and precipitated aggregates. To this end, a

calibration curve of asphaltene absorbance (optical density) at a wavelength of 450 nm and as a

function of its concentration (expressed in terms of a bitumen dilution ratio) was generated to

establish a baseline asphaltene absorbance in the oil phase under soluble conditions. The optical

density of all the solutions was measured using an Ocean Optics spectrophotometer (model

20

#HR2000). Toluene was used as the solvent because it completely solubilizes asphaltenes. The

relationship between the optical density (absorbance) and concentration of asphaltenes is shown

in Figure 2.1.

y = 2.08x + 0.09 ( 0.03)R² = 0.99 0.09

0

0.5

1

1.5

2

2.5

3

3.5

0.0 0.5 1.0 1.5

Ab

so

rba

nc

e

Bitumen to Toluene Ratio (x103)

Figure 2.1: Calibration of asphaltene absorbance at a wavelength of 450 nm and as a function

of the bitumen to toluene ratio.

To determine the fraction of asphaltene losses from the separated oil phase of

experimental systems after mixing and centrifuging, a sample (which already includes heptol)

was further diluted with a known volume of toluene and its resulting optical density was

measured. The following difference between the baseline and experimental absorbances

represents the fraction of asphaltene losses to the rag layer:

100%

AbsorbanceBaseline

AbsorbancealExperimentAbsorbanceBaselineLossesAsphaltene (Eq. 1)

It is important to clarify that the fraction of asphaltene losses is not used in the material

balance of oil and it does not necessarily depend, at least from the material balance point of

view, on the fraction of oil losses to the rag layer. For example, one could have a significant

21

fraction of oil losses to the rag layer and yet no asphaltene losses if the concentration of

asphaltenes in the separated oil phase is the same as that of the original oil. The fraction of

asphaltene losses only reflects the change in the concentration of asphaltenes in the separated oil

phase with respect to the original oil.

2.3.6 Interfacial Tension Measurements

To measure the γow of the systems analyzed, a spinning drop tensiometer manufactured

by Temco Inc. (model #500) was used. In this technique, a borosilicate glass tube was first

completely filled with the separated aqueous phase that was recovered from the formulations

after mixing and centrifuging. A droplet (~5 µL) of the separated oil phase was then inserted

into the aqueous phase and the glass tube was spun at increasing rpm values until the oil droplet

expanded sufficiently such that its length was 4 times greater than its width. Once the oil droplet

expansion reached an equilibrium value at a given rpm, the γow was calculated as follows:

4

32wow

(Eq. 2)

In this equation, Δρ is the difference in density between the heavy (water) and light (oil)

phases, ω is the rotational velocity, and w is the width of the expanded oil droplet. Through the

use of a thermocouple and temperature controller, the operating temperature of the device was

easily adjusted.

22

2.4 RESULTS

2.4.1 Phase Separation of Solvent-Bitumen-Water Systems

The separated oil phase, aqueous phase, and rag layer of each of the mixed and

centrifuged formulations are analyzed using optical, cross-polarized, and fluorescence

microscopy. Micrographs taken for a system containing 68.2 wt% heptol 80/20, 6.8 wt%

bitumen, and 25 wt% water at 25°C are displayed in Figure 2.2. The images of Figure 2.2

illustrate that in these experiments, significant w/o emulsions occur throughout the rag layer

whereas the excess oil and aqueous phases contain insignificant traces of water and oil

respectively. As a result, the assumption of pure recoverable oil and aqueous phases when

performing material balances is reasonable. These results are consistent for all other

formulations analyzed. Although liquid crystals are sometimes observed in systems prepared at

25°C with heptol 80/20 and heptol 50/50 as the diluents, their relative concentration within the

rag layer (<1%) suggests that they play a minor role in promoting emulsion stability. No liquid

crystal formation occurred in the systems prepared at 80°C.

23

250 µm

RAG

LAYER

OIL

PHASE

AQUEOUS

PHASE

i) ii)

iii)

ii)

i)

iii)

i)

ii)

iii)

Figure 2.2: (i) Optical, (ii) cross-polarized, and (iii) fluorescent micrographs of samples of the

separated oil phase, aqueous phase, and rag layer for a system containing 68.2 wt% heptol

80/20, 6.8 wt% bitumen, and 25 wt% water at 25°C.

When the phase heights and micrographs are analyzed in greater detail, a material

balance closure is obtained in which the volumes of the separated oil and aqueous phases and

the volumes of oil and water within the rag layer are estimated. Table 2.2 displays calculated

values for each of these parameters for systems prepared with 50 wt% water at various heptol

80/20 to bitumen dilution ratios and 25°C. The columns labeled as % error refer to the

difference between the total volume of each component (oil or water) initially added and that

calculated using the material balance.

24

Table 2.2: Material balance closure for systems prepared with 50 wt% water (W) at various

heptol 80/20 (H) to bitumen (B) dilution ratios and 25°C.

Composition

(cm3) Phases

Phase

Height

(mm)

Phase

Volume

(cm3)

Rag Layer

Image

(Scion)

Water to

Oil Ratio

(vol/vol)

%

Asphaltene

Losses

%

Error

Oil

%

Error

Water W H B

5 2.8 3

Oil 34 4.9

5.6 8 0.1 -0.1 Rag 40 5.7

Water 1 0.1

5 3.4 2.5

Oil 35 5

5.5 10 -0.12 0.2 Rag 39.5 5.7

Water 1.5 0.2

5 5.2 1.2

Oil 37 5.2

2.6 13 -0.3 0.4 Rag 30 4.2

Water 14 2

5 5.5 1

Oil 37 5.2

3.2 17 -0.3 0.4 Rag 39 5.5

Water 6 0.8

5 6.3 0.5

Oil 36 5.1

2.7 27 0 0 Rag 41 5.9

Water 5 0.7

On the basis of these results, the material balance closure is satisfactory, which validates

the procedures used to analyze the systems produced with the batch emulsification-separation

protocol.

2.4.2 Development of Batch Emulsification-Separation Protocol

To determine the appropriate mixing conditions for solvent-bitumen-water systems that

reflect the rag layers produced in industry, various formulations are mixed at 3200 rpm from 20

seconds to 5 minutes. The samples are then centrifuged for 20 minutes at 1000 g and the fraction

of oil and asphaltene losses to the rag layer are determined according to the procedures

described above. The data is presented in Figures 2.3 (a) and (b).

25

(a)

Oil

Lo

sses

to R

ag

La

yer

(V

ol

%)

Asp

ha

lten

e L

oss

es t

o R

ag

La

yer

(%

)

Mixing Time (Minutes)

56.25 wt% Heptol 80/20-18.75 wt%

Bitumen-25 wt% Water

18.75 wt% Heptol 80/20-6.25 wt%

Bitumen-75 wt% Water

(b)

Figure 2.3: (a) Oil and (b) asphaltene losses to the rag layer as a function of the mixing time for

systems prepared with either 25 wt% or 75 wt% water and the balance oil with a heptol 80/20 to

bitumen dilution ratio of 3 at 25°C.

The trends observed in Figures 2.3 (a) and (b) suggest that mixing conditions have a

significant impact on rag layer stability and the fate of asphaltenes. Furthermore, the data shows

from the time required for oil and asphaltene losses to plateau for the different formulations that

a steady state condition is reached at ~2 minutes. This time is used as a standard mixing time

going forward.

To select the appropriate separation conditions for these rag layers, various formulations

are mixed for 2 minutes at 3200 rpm and then either left to settle on their own (1 g) for up to

5000 minutes or arbitrarily centrifuged at 35 g for up to 143 minutes. The fraction of oil losses

26

to the rag layer at different times is then determined from the measured phase volumes (heights).

In Figures 2.4 (a)-(d), the fraction of oil losses to the rag layer is presented for different

formulations as a function of the g force × time. According to common practice and recent

models, the separation of a given suspension is expected to be a function of the g force × time as

it is proportional to the settling velocity × time16

.

g Force x Time (g x Minutes)

Oil L

osses t

o R

ag

Layer

(Vo

l %

)

(a) (b)

(c) (d)

1 g, 5000 Minutes 35 g, 143 Minutes

Figure 2.4: Oil losses to the rag layer as a function of the g force × time for systems prepared

with (a) 12.5 wt% heptol 80/20, 12.5 wt% bitumen, and 75 wt% water, (b) 20 wt% heptol 80/20,

5 wt% bitumen, and 75 wt% water, (c) 37.5 wt% heptol 80/20, 37.5 wt% bitumen, and 25 wt%

water, and (d) 60 wt% heptol 80/20, 15 wt% bitumen, and 25 wt% water.

The data in Figures 2.4 (a)-(d) confirms that g force × time is indeed the governing

parameter that describes the phase separation efficiency of a given system. In this work, 500 g ×

27

min is selected as the condition to evaluate the separation of rag layers. This parameter is

consistent with that published in the literature17

.

2.4.3 Trends of Oil and Asphaltene Losses to the Rag Layer

Oil and asphaltene losses to the rag layer are evaluated as a function of the solvent-

bitumen-water ratios, solvent aromaticity, and temperature. A summary of the obtained results is

presented in Figures 2.5 (a)-(f).

(b)

(d)

(e) (f)

Oil

Lo

sses

to

Ra

g L

ay

er (

Vo

l%)

Asp

ha

lten

e L

oss

es t

o R

ag

La

yer

(%

)

Heptol 80/20 to Bitumen Dilution Ratio Heptol 80/20 to Bitumen Dilution Ratio

Heptol 50/50 to Bitumen Dilution Ratio Heptol 50/50 to Bitumen Dilution Ratio

(c)

(a)

75 wt% Water 50 wt% Water 25 wt% Water 9.1 wt% Water 0 wt% Water

Figure 2.5: Oil and asphaltene losses to the rag layer for systems prepared with (a and b) heptol

80/20 at 25°C, (c and d) heptol 80/20 at 80°C, and (e and f) heptol 50/50 at 25°C.

28

To facilitate the presentation of the data in Figures 2.5 (a)-(f) and to enhance the visual

clarity of the obtained trends, only the upper or lower half of the error bars is shown to prevent

overlapping of the results. For systems prepared with heptol 80/20 as the solvent, increasing the

bitumen dilution ratio or water content produces an increase in the fraction of oil and asphaltene

losses to the rag layer. At 80°C, however, the oil losses are approximately half of that observed

at 25°C, whereas the fraction of asphaltene losses is only slightly reduced. For systems prepared

with heptol 50/50 as the solvent at 25°C, increasing the bitumen dilution ratio reduces the

amount of oil and asphaltene losses to the rag layer and excellent phase separation is obtained

for most systems, with the exception of those prepared with high water content and low bitumen

dilution ratios. Although traces of insoluble asphaltene aggregates are physically observed in

formulations prepared with heptol 80/20 and heptol 50/50 in the presence of 0 wt% water, the

absolute error in the UV-vis method used to estimate asphaltene losses is close to 3%,

suggesting that precipitation is minimal in these systems. To estimate the onset of asphaltene

precipitation, the fraction of asphaltene losses from the oil is determined as a function of the

heptane concentration in heptol at a dilution ratio of 10 and in the absence of water. It is found

that the fraction of asphaltene losses from the oil phase increases from 5% to 52% (forming a

substantial amount of precipitate at the bottom of the vial) as the percentage of heptane in heptol

is increased from 80 vol% to 85 vol%. This observation confirms that heptol 80/20 can be

considered a “poor” solvent because its composition is very close to that of the onset of

asphaltene precipitation. On the other hand, heptol 50/50 is more aromatic than heptol 80/20 and

therefore a more suitable solvent to dissolve asphaltenes.

From a processing perspective, the ternary phase diagrams presented in Figures 2.6 (a)-

(c) display transitions in achievable oil recovery under a given set of operational conditions.

29

Gradient Scale

(Oil Losses to Rag Layer (Vol%)

wt% Bitumen

(a) (b)

wt% Bitumen

(c)

wt% Bitumen

Figure 2.6: Ternary phase diagrams for systems prepared with (a) heptol 80/20 at 25°C, (b)

heptol 80/20 at 80°C, and (c) heptol 50/50 at 25°C.

It should be noted that, although the trends of oil and asphaltene losses to the rag layer

using coker feed bitumen are relevant to the bitumen recovery process, one cannot forget the

fact that, in producing coker feed bitumen, the composition of bitumen may have changed to a

certain degree. It is important to keep this in mind when trying to extrapolate these results and

ensuing discussions to other bitumen sources or heavy crude oils.

30

2.5 DISCUSSIONS

2.5.1 Effect of Heptol 80/20 to Bitumen Dilution Ratio and Water Content on

Rag Layer Stability

As shown in Figures 2.5 (a)-(d), an increase in the heptol 80/20 to bitumen dilution ratio

results in increased oil and asphaltene losses to the rag layer. This effect can be explained on the

basis that heptol 80/20 is enriched in heptane to a point that it is almost at the asphaltene

precipitation threshold and it is therefore an environment not favorable for asphaltenes18,19

. It is

believed, according to Gawrys et al., that asphaltenes become less soluble in the oil phase at

dilution ratios up to 10 because their aggregate size continuously increases20

. The shift in the

equilibrium partitioning of asphaltenes within solvent-bitumen-water systems can be

represented using the following simple relation2:

(Eq. 3)

To interpret the data of Figures 2.5 (a)-(d) in light of Equation 3, it is necessary to

highlight that, in the absence of water, asphaltene losses, which are typically ≤5%, are only

associated with precipitation. However, in the presence of water, a significant fraction of

asphaltene losses are associated with the formation of asphaltene aggregates that segregate near

the oil-water interface. By promoting the formation of interfacially adsorbed asphaltene

aggregates at increased heptol 80/20 to bitumen dilution ratios, rigid asphaltene skins are

produced, which enhance emulsion stability and lead to an increased fraction of oil losses to the

rag layer. This explanation is consistent with the observations of Sztukowski et al. and Moran et

al. that, when the solubility of asphaltenes is reduced at increased bitumen dilution ratios in the

presence of water, emulsion stability is enhanced21,22

. These researchers suggested that

Interfacial

Asphaltenes

Soluble

Asphaltenes Precipitated

Asphaltenes

31

asphaltene aggregates adhere to the primary layer of adsorbed asphaltenes at the oil-water

interface, preventing bridging between adjacent droplets. There is little known about the

molecular interactions that lead to the formation of interfacial asphaltene aggregates. Tan et al.,

however, proposed that hydrogen bonding may induce their association at the oil-water

interface23

.

These findings suggest that it is important to quantify Aow and how it is produced. As

evidenced in Figures 2.3 (a) and (b), the method of emulsification (such as mixing time)

influences the fraction of oil and asphaltene losses to the rag layer. These observations are

consistent with previous reports24

. A simplistic way of understanding the process of

emulsification is through the use of the Weber (We) number. The We number relates the inertial

and interfacial forces of emulsion droplets as follows25,26

:

ow

rel

E

d

E

c vdWe

2

(Eq. 4)

In this equation, E

c is the density of the emulsion’s continuous oil phase, E

dd is the

average emulsion droplet diameter, and vrel is the relative mixing velocity of the continuous oil

phase and emulsified aqueous phase. As the We number of an emulsified water droplet

increases above its critical value (~1-2), it becomes unstable and breaks up into smaller droplets

until E

dd becomes small enough to offset any further shearing. According to Figures 2.3 (a) and

(b), the process of emulsion droplet break-up takes about 2 minutes, which is considerably lower

than the amount of time required for γow to reach equilibrium (~16-40 minutes depending on the

asphaltene concentration)27

. This difference in time scale can be explained by the formation of

asphaltene skins. It is important to clarify that in industrial practice the process of emulsification

can be quite different, depending upon the flow dynamics inherent to it.

32

The value of Aow can be estimated as the product of the volume of the emulsified

aqueous phase ( E

wV ) and the average surface area ( E

dA ) to volume ( E

dV ) ratio of the spherical

water droplets as follows:

E

d

E

dE

wowV

AVA (Eq. 5)

The average E

dA to E

dV ratio of emulsified water droplets can be expressed in terms of

E

dd as follows2:

E

d

E

d

E

d

dV

A 6 (Eq. 6)

To calculate E

dd , the diameters of the squeezed cylindrical droplets obtained through

image analysis are first transformed into sphere-equivalent diameters. The calculated sphere-

equivalent water droplet diameters are then converted into E

dd as follows2:

2

3

E

ii

E

iiE

d

dF

dFd (Eq. 7)

In this equation, E

iF is the number frequency of water droplets of sphere-equivalent

diameter E

id . Equations 6 and 7 can be incorporated into Equation 5 to yield the following

alternative expression for Aow:

2

3

6

E

ii

E

ii

E

w

ow

dF

dF

VA (Eq. 8)

33

Figures 2.7 (a) and (b) present calculated estimates of E

dd as a function of the heptol

80/20 to bitumen dilution ratio and water content as well as a sample E

id distribution at 50 wt%

water. Figure 2.7 (c) presents estimates of Aow normalized by the volume of bitumen.

0

5

10

15

20

25

30

35

40

0 2 4 6 8 10

Sau

ter

Mean

Dia

mete

r (µ

m)

Heptol (80/20) to Bitumen Dilution Ratio

75 wt% SW

50 wt% SW

25 wt% SW

9.09 wt% SW

0

2

4

6H/B=1/1.5

H/B=1/1

H/B=3/1

H/B=4/1

H/B=10/1

FiE

diE (µm)

0

2000

4000

6000

8000

10000

12000

14000

16000

18000

0 2 4 6 8 10

Are

a/V

olu

me o

f B

itu

men

(1/c

m)

Heptol (80/20) to Bitumen Dilution Ratio

75 wt% SW

50 wt% SW

25 wt% SW

9.09 wt% SW

(a)

(b)

(c)

dd

E(µ

m)

No

rma

lize

d A

ow

(1/c

m)

Figure 2.7: (a) E

dd , (b) sample E

id distribution at 50 wt% water, and (c) normalized Aow of

heptol 80/20-diluted bitumen droplets at 25°C.

34

With respect to the calculated E

dd estimates in Figure 2.7 (a), larger droplet sizes are

obtained at increased heptol 80/20 to bitumen dilution ratios, which, as will be discussed later,

can be attributed to increased γow. With respect to the normalized Aow, Figure 2.7 (c) shows that

increasing the water content from 0 to 50 wt% increases Aow. For systems containing 25 wt%

and 75 wt% water, Aow are similar. These results suggest that Aow is proportional to the product

of the oil and water volume fractions in the test tube. Furthermore, one may also conclude by

analyzing the asphaltene losses presented in Figure 2.5 (b) in conjunction with the normalized

Aow results that the asphaltene surface coverage, calculated by assuming that the asphaltene

losses are due to adsorption only, for most heptol 80/20 systems at 25°C is 40-50 mg/m2. The

asphaltene surface coverage is ~1 order of magnitude larger for systems prepared with 9.1 wt%

water. This is likely due to the smaller Aow produced with these systems and the assumption that

all asphaltene losses correspond to adsorption. These observations support the hypothesis that

the formation of rag layers and their stability is highly dependent upon the amount of Aow

created during the different process operations. The average asphaltene surface coverage of 40-

50 mg/m2 is of the same order of magnitude but several folds larger than that obtained by

Gafonova and Yarranton using heptol 50/50 (~10 mg/m2)28

. For the heptol 50/50 systems

presented later, asphaltene surface coverage values of 6-12 mg/m2 are obtained at dilution ratios

≥3, corresponding to the range of asphaltene concentrations considered by Gafonova and

Yarranton.

The normalized Aow is also observed to increase with the heptol 80/20 to bitumen

dilution ratio. This suggests that diluting bitumen with such a poor solvent promotes the

segregation of asphaltenes near the oil-water interface, thus stabilizing more of the interface

35

created during mixing. This observation is consistent with the larger asphaltene losses at

increased heptol 80/20 to bitumen dilution ratios observed in Figure 2.5 (b).

2.5.2 Effect of Temperature and Solvent Aromaticity on Rag Layer Stability

According to Figures 2.5 (c) and (e), an improved separation of the oil phase can be

obtained by either increasing the temperature or the solvent aromaticity. With regards to an

increase in the temperature, it is important to note that this strategy reduces the fraction of oil

losses to the rag layer but does very little to the fraction of asphaltene losses. In this case, it

seems that increasing the temperature facilitates the drainage of oil from the rag layer, which

may be due to its reduction in viscosity and/or a weakening of the dipole-induced dipole or van

der Waals forces between the oil and asphaltene skins29,30,31

. This drainage effect at elevated

temperatures is believed to be furthermore assisted by the fact that the reduced viscosity of the

continuous oil phase and the reduced yield strength and elastic modulus of asphaltene skins

leads to more effective collisions of dispersed water droplets32

.

When Figures 2.5 (e) and (f) are compared to Figures 2.5 (a) and (b), one can appreciate

that the trends of oil and asphaltene losses to the rag layer with increasing bitumen dilution

ratios are opposite for heptol 80/20 and heptol 50/50. In the case of heptol 80/20, the large

fraction of heptane in the oil phase shifts the equilibrium of Equation 3 towards the formation of

interfacial asphaltene aggregates. These comments are consistent with the studies performed by

McLean et al. that show that asphaltenes stabilize w/o emulsions by collecting at the interface in

the form of aggregates33

. On the other hand, heptol 50/50 is an aromatic solvent, containing a

heptane concentration considerably lower than the threshold precipitation concentration of 80-

85 vol% of heptane. Thus, increasing the heptol 50/50 to bitumen dilution ratio reduces the

36

tendency of asphaltenes to segregate at the oil-water interface. This point will be discussed in

more detail in the next section.

A surprising element in Figures 2.5 (a) and (e) is that, at low bitumen dilution ratios

(≤1), heptol 50/50 significantly magnifies emulsion stability in comparison to heptol 80/20. A

likely explanation for this, according to the experimental proof Figures 2.8 (a)-(d), is that the

reduced droplet diameter in heptol 50/50 systems provides for an increased Aow that asphaltenes

can adsorb/segregate to. The Aow produced in heptol 80/20 and heptol 50/50 systems under such

processing conditions are 0.4 m2 and 1.5 m

2 respectively. Furthermore, the mechanical

properties of the adsorbed asphaltene skins and/or the kinetics of their formation (likely faster in

heptol 50/50 at low dilution ratios due to the enhanced diffusion of smaller aggregate sizes) may

be important factors to consider. Figures 2.8 (a)-(d) show that the water droplets in heptol 80/20

systems are separated but, for the case of heptol 50/50 systems, there is a network of smaller

droplets that are attached to each other.

37

Heptol 80/20 Systems

Heptol 50/50 Systems

a) b)

c) d)

250 µm

Figure 2.8: Morphology of emulsions for systems prepared with (a) 10 wt% heptol 80/20, 15

wt% bitumen, and 75 wt% water, (b) 12.5 wt% heptol 80/20, 12.5 wt% bitumen, and 75 wt%

water, (c) 10 wt% heptol 50/50, 15 wt% bitumen, and 75 wt% water, and (d) 12.5 wt% heptol

50/50, 12.5 wt% bitumen, and 75 wt% water.

2.5.3 Surface Activity of Asphaltenes

The γow of the separated oil and aqueous phases for the formulations in Figures 2.5 (a)-

(f) are measured to substantiate previously established phase behavior results and gain a better

understanding of the surface activity of asphaltenes and their role in promoting emulsion

stability. This data is presented in Figures 2.9 (a)-(c) as a function of the solvent-bitumen-water

ratios.

38

(b)

(c)

Heptol (50/50) to Bitumen Dilution Ratio

Heptol (80/20) to Bitumen Dilution Ratio

γo

w(m

N/m

)

(a)

(b)

(c)

75 wt% Water 50 wt% Water 25 wt% Water 9.1 wt% Water

Figure 2.9: The measured γow as a function of solvent-bitumen-water ratios for systems

prepared with (a) heptol 80/20 at 25°C, (b) heptol 80/20 at 80°C, and (c) heptol 50/50 at 25°C.

According to this data, the γow of all the systems increases with increasing heptol to

bitumen dilution ratios and water content. For the case of the heptol 80/20 systems in Figures

2.9 (a) and (b), the increased γow can be linked to the increased asphaltene losses from the

separated oil phase at an increased bitumen dilution ratio and water content. In the case of

39

heptol 50/50 systems, the increased γow can be explained on the basis of a dilution effect. When

the bitumen dilution ratio is increased, the concentration of asphaltenes in the recovered oil

phase falls below their critical transition concentration and their surface activity is compromised

(see discussion of Figure 2.10). The increased γow with the heptol to bitumen dilution ratios can,

according to the critical We number, be linked to the increase in E

dd .

To evaluate the surface activity of asphaltenes in heptol 80/20 and heptol 50/50,

asphaltenes are first extracted from bitumen (after a 13 to 1 dilution with heptane) and then

redissolved in each solvent at different concentrations. The measured γow as a function of the

asphaltene concentration is presented in Figure 2.10.

8

13

18

23

28

33

0.001 0.01 0.1 1 10

Asphaltene Concentration (%)

γow

(mN

/m)

Heptol 80/20

Heptol 50/50

Critical Aggregation Concentration

(CAC) for asphaltene monomers in

heptol 80/20

Critical Transition Concentration

(CTC) for asphaltene aggregates

in heptol 80/20

Figure 2.10: The measured γow (against water) of asphaltenes diluted in heptol 80/20 and heptol

50/50 at 25°C.

The difference between the initial and final (equilibrium) asphaltene concentrations are

estimated using known values of asphaltene surface coverage. At low asphaltene concentrations

(~1×10-3

%), one expects monolayer coverage (~1 mg/m2). Because a 5 µL droplet of

40

asphaltenes in heptol is introduced into the tensiometer, this represents an equivalent loss of

asphaltenes of ~2×10-4

% (or a 20% difference in concentration). At high asphaltene

concentrations (~6%), the asphaltene surface coverage is, at most, 50 mg/m2. This would

produce an equivalent loss of asphaltenes of ~1×10-2

% (or a 0.2% difference in concentration).

Therefore, the trends observed in Figure 2.10 should approach the trends observed if the

equilibrium concentrations were used instead of the initial concentrations. As the concentration

of asphaltenes in heptol 80/20 increases up to ~250 ppm, the γow decreases to a value near 17

mN/m. This transition is typically referred to as the critical aggregation concentration (or CAC)

where asphaltene aggregates begin to form. The values of γow leading up to the CAC and near

the CAC are consistent with those reported in most of the studies involving asphaltenes and

bitumen.

As the asphaltene concentration is increased beyond the CAC for heptol 80/20 systems,

γow remains nearly constant until the asphaltene concentration approaches 1%. At this point,

which will be referred to as the critical transition concentration (CTC), there is a second

transition in γow. The γow values leading up to the CTC and after the CTC are not as common as

those associated with the CAC. They have, however, been observed for select systems

containing asphaltenes with acidic or highly polar moieties34,35

. One cannot disregard the

possibility that these highly polar species are residual process aids in the bitumen. It is unlikely,

however, that they would have precipitated along with asphaltenes in the hexane dilution step.

This kind of transition associated with the CTC has been observed before in microemulsions

where it is commonly referred to as the critical microemulsion concentration (cµc)36

. The cµc

marks the formation of net-zero curvature films of surfactants adsorbed at oil-water interfaces. It

is possible, therefore, that the CTC marks the change in morphology of the asphaltene

41

aggregates segregated near the oil-water interface. Considering Figure 2.10, Figures 2.5 (b) and

(f), and that bitumen contains ~15% asphaltenes, one can infer that, for system above the CTC

(dilution ratios lower than 10 for heptol 80/20 and lower than 3 for heptol 50/50), asphaltenes

tend to segregate near the oil-water interface (the left side of Equation 3). This also reinforces

the fact that asphaltenes in heptol 50/50 are less surface-active than asphaltenes dissolved in a

“poor” solvent such as heptol 80/20.

2.5.4 Correlation of Oil and Asphaltene Losses to the Rag Layer

The general correlations of oil and asphaltene losses to the rag layer for systems

prepared with heptol 80/20 at 25°C and 80°C are presented in Figure 2.11.

(b)

(a)

Oil

Lo

sses

to

Ra

g L

ay

er (

Vo

l%)

Asphaltene Losses to Rag Layer (%)

Figure 2.11: Correlations of oil and asphaltene losses to the rag layer for systems prepared with

heptol 80/20 at (a) 25°C and (b) 80°C.

42

This data reinforces the hypothesis that asphaltene losses to the rag layer help stabilize

emulsions because a reduced concentration of asphaltenes is observed in the recovered oil phase

at increased oil losses to the rag layer. While one could argue that there is a linear correlation

between these variables, there is also substantial scatter in the data. This scatter is partially

explained by variations in the physical characteristics of monomeric asphaltenes37

. The resulting

packing constraints of asphaltenes affect their aggregation behavior at different dilution ratios

and ability to adsorb at the oil-water interface. As a result, the mechanical properties of

asphaltene skins produced differ, which results in different extents of rag layer formation and

stabilization. The real value of Figure 2.11 is to highlight the importance of the formulation

conditions because they affect the fraction of asphaltene losses that, in turn, influence the

fraction of oil losses to the rag layer. To find a numerical relation, the equilibrium between the

dissolved and interfacial asphaltenes needs to be better understood as well as the morphology

and mechanical properties of the asphaltene skins produced under different formulation

conditions.

2.6 CONCLUSIONS

In this work, a methodology involving phase volume measurements, optical, cross-

polarized, and fluorescence microscopy, and UV-vis spectroscopy was introduced to evaluate

the stability of solvent-bitumen-water rag layers and its relation to the surface activity of

asphaltenes. It was confirmed that the phase separation of these systems was severely impacted

by the formation of asphaltene aggregates that segregated near the oil-water interface. A major

finding in this study was that the water content in the formulation played a critical role on the

observed phase behavior because it impacted the amount of interfacial area produced that

asphaltenes could potentially segregate to. The tendency of asphaltenes to segregate or adsorb at

43

the oil-water interface depended upon the nature of the solvent used to dilute the bitumen. When

using a poor solvent, such as heptol 80/20, whose high heptane concentration is close to that of

the onset of asphaltene precipitation, an increase in the dilution ratio created a less suitable

environment for asphaltenes. This induced the segregation of asphaltenes at the oil-water

interface. Using a more aromatic solvent, such as heptol 50/50, reduced the tendency of

asphaltenes to segregate near the oil-water interface as the dilution ratio increased. In all cases,

the segregation of asphaltenes at the oil-water interface, as reflected by the asphaltene losses,

was correlated with increased oil losses to the rag layer. Another interesting finding was that the

increase in temperature did not seem to affect the segregation of asphaltenes. It did, however,

facilitate the drainage (and hence recovery) of the continuous oil phase from the rag layer.

Finally, the interfacial tension studies of precipitated and redissolved asphaltenes suggested that

asphaltenes may undergo different transitions with an increase in their concentration. The nature

of such transitions is a topic that should be further investigated.

2.7 REFERENCES

1 Yan, Z.; Elliott, J.A.W.; Masliyah, J. Roles of Various Bitumen Components in the Stability of

Water-in-Diluted Bitumen Emulsions. J. Colloid Interface Sci. 1999, 220, 329-337.

2 Yarranton, H.W.; Hussein, H.; Masliyah, J. Water-in-Hydrocarbon Emulsions Stabilized by

Asphaltenes at Low Concentrations. J. Colloid Interface Sci. 2000, 228, 52-63.

44

3 Rondón, M.; Pereira, J.C.; Bouriat, P.; Graciaa, A.; Lachaise, J.; Salager, J.-L. Breaking of

Water-in-Crude Oil Emulsions. 2. Influence of Asphaltene Concentration and Diluent Nature on

Demulsifier Action. Energy Fuels 2008, 22, 702-707.

4 Yeung, A.; Dabros, T.; Masliyah, J.; Czarnecki, J. Micropipette: A New Technique in

Emulsion Research. Colloids Surf., A 2000, 174, 169-181.

5 Tsamantakis, C.; Masliyah, J.; Yeung, A.; Gentzis, T. Investigation of the Interfacial

Properties of Water-in-Diluted Bitumen Emulsions using Micropipette Techniques. J. Colloid

Interface Sci. 2005, 284, 176-183.

6 Yang, X.; Hamza, H.; Czarnecki, J. Investigation of Subfractions of Athabasca Asphaltenes

and Their Role in Emulsion Stability. Energy Fuels 2004, 18, 770-777.

7 Wu, X. Investigating the Stability Mechanism of Water-in-Diluted Bitumen Emulsions

through Isolation and Characterization of the Stabilizing Materials at the Interface. Energy Fuels

2003, 17, 179-190.

8 Gelin, F.; Grutters, M.; Cornelisse, P.; Taylor, S. Asphaltene Precipitation from Live Oil

Containing Emulsified Water. Proceedings of the 5th

International Conference on Petroleum

Phase Behaviour and Fouling, Banff, Alberta, Canada, 2004.

9 Horváth-Szabó, G.; Czarnecki, J.; Masliyah, J. H. Sandwich Structures at Oil-Water Interfaces

under Alkaline Conditions. J. Colloid Interface Sci. 2002, 253, 427-434.

45

10 Yang, X.; Czarnecki, J. The Effect of Naphtha to Bitumen Ratio on Properties of Water in

Diluted Bitumen Emulsions. Colloids Surf., A 2002, 211, 213-222.

11 Häger, M.; Ese, M.-H.; Sjöblom, J. Emulsion Inversion in an Oil-Surfactant-Water System

Based on Model Naphthenic Acids under Alkaline Conditions. J. Dispersion Sci. Technol. 2005,

26, 673-682.

12 Akbarzadeh, K.; Alboudwarej, H.; Svrcek, W.Y.; Yarranton, H.W. A Generalized Regular

Solution Model for Asphaltene Precipitation from n-Alkane Diluted Heavy Oils and Bitumens.

Fluid Phase Equilib. 2005, 232, 159-170.

13 Allen, E.W. Process Water Treatment in Canada’s Oil Sands Industry: I. Target Pollutants and

Treatment Objectives. Journal of Environmental Engineering and Science 2008, 7, 123-138.

14 Varadaraj, R.; Brons, C. Molecular Origins of Crude Oil Interfacial Activity Part 3:

Characterization of the Complex Fluid Rag Layers Formed at Crude Oil-Water Interfaces.

Energy Fuels 2007, 21, 1617-1621.

15 Scion Corporation. Scion Imaging Software. Last Updated: December, 2007.

16 Leung, W. Separation of Dispersed Suspension in Rotating Test Tube. Sep. Purif. Technol.

2004, 38, 99-119.

46

17 Klasson, K.T.; Taylor, P.A.; Walker Jr., J.F.; Jones, S.A.; Cummins, R.L.; Richardson, S.A.

Investigation of a Centrifugal Separator for In-Well Oil Water Separation. Pet. Sci. Technol.

2004, 22, 1143-1159.

18 Spiecker, P.M.; Gawrys, K.L.; Trail, C.B.; Kilpatrick, P.K. Effects of Petroleum Resins on

Asphaltene Aggregation and Water-in-Oil Emulsion Formation. Colloids Surf., A 2003, 220, 9-

27.

19 Barré, L.; Simon, S.; Palermo, T. Solution Properties of Asphaltenes. Langmuir 2008, 24,

3709-3717.

20 Gawrys, K.L.; Spiecker, P.M.; Kilpatrick, P.K. The Role of Asphaltene Solubility and

Chemistry on Asphaltene Aggregation. ACS Div. Pet Chem. Preprints 2002, 47, 332-335.

21 Sztukowski, D.M.; Yarranton, H.W. Characterization and Interfacial Behavior of Oil Sands

Solids Implicated in Emulsion Stability. J. Dispersion Sci. Technol. 2004, 25, 299-310.

22 Moran, K.; Czarnecki, J. Competitive Adsorption of Sodium Naphthenates and Naturally

Occurring Species at Water-in-Crude Oil Emulsion Droplet Surfaces. Colloids Surf., A 2007,

292, 87-98.

23 Tan, X.; Fenniri, H.; Gray, M.R. Investigation of the Effects of Water on Aggregation of

Model Asphaltenes in Organic Solution. Proceedings of the 9th

Annual International Conference

on Petroleum Phase Behaviour and Fouling, Victoria, British Columbia, Canada, 2008.

47

24 Johansen, E.J.; Skjärvö, I.M.; Lund, T.; Sjöblom, J.; Söderlund, H.; Boström, G. Water-in-

Crude Oil Emulsions from the Norwegian Continental Shelf Part I. Formation, Characterization,

and Stability Correlations. Colloids Surf. 1988, 34, 353-370.

25 Walstra, P. Principles of Emulsion Formation. Chem. Eng. Sci. 1993, 48, 333-349.

26 Kocherginsky, N.M.; Tan, C.L.; Lu, W.F. Demulsification of Water-in-Oil Emulsions via

Filtration through a Hydrophilic Polymer Membrane. J. Membr. Sci. 2003, 220, 117-128.

27 Jeribi, M.; Almir-Assad, B.; Langevin, D.; Hénaut, I.; Argillier, J.F. Adsorption Kinetics of

Asphaltenes at Liquid Interfaces. J. Colloid Interface Sci. 2002, 256, 268-272.

28 Gafonova, O.; Yarranton, H. The Stabilization of Water-in-Hydrocarbon Emulsions by

Asphaltenes and Resins. J. Colloid Interface Sci. 2001, 241, 469-478.

29 Rosen, M.J. Surfactants and Interfacial Phenomena, 3

rd ed.; John Wiley & Sons, Inc.: New

Jersey, 2004.

30 Liu, J.; Zhang, L.; Xu, Z.; Masliyah, J. Colloidal Interactions between Asphaltene Surfaces in

Aqueous Solutions. Langmuir 2006, 22, 1485-1492.

31 Nour, A.H.; Suliman, A.; Hadow, M.M. Stabilization Mechanisms of Water-in-Crude Oil

Emulsions. J. App. Sci. 2008, 8, 1571-1575.

32 Rodríguez-Abreu, C.; Lazzari, M. Emulsions with Structured Continuous Phases. Curr. Opin.

Colloid Interface Sci. 2008, 13, 198-205.

48

33 McLean, J.D.; Kilpatrick, P.K. Effects of Asphaltene Solvency on Stability of Water-in-Crude

Oil Emulsions. J. Colloid Interface Sci. 1997, 189, 242-253.

34 Norgård, E.L.; Sjöblom, J. Model Compounds for Asphaltenes and C80 Isoprenoid Tetraacids.

Part I: Synthesis and Interfacial Activities. J. Dispersion Sci. Technol. 2008, 29, 1114-1122.

35 Acevedo, S.; Escobar, G.; Ranaudo, M.A.; Khazen, J.; Borges, B.; Pereira, J.C.; Méndez, B.

Isolation and Characterization of Low and High Molecular Weight Acidic Compounds from

Cerro Negro Extraheavy Crude Oil. Role of These Acids in the Interfacial Properties of the

Crude Oil Emulsions. Energy Fuels 1999, 13, 333-335.

36 Acosta, E.J.; Harwell, J.H.; Sabatini, D.A. Self-Assembly in Linker-Modified

Microemulsions. J. Colloid Interface Sci. 2004, 274, 652-664.

37 Victorov, A.I.; Smirnova, N.A. Thermodynamic Model of Petroleum Fluids Containing

Polydisperse Asphaltene Aggregates. Ind. Eng. Chem. Res. 1998, 37, 3242-3251.

49

CHAPTER 3:

IMPACT OF ASPHALTENES AND NAPHTHENIC

AMPHIPHILES ON THE PHASE BEHAVIOR OF SOLVENT-

BITUMEN-WATER SYSTEMS

This chapter is derived from the following published manuscript:

Kiran, S.K.; Ng, S.; Acosta, E.J. Impact of Asphaltenes and Naphthenic Amphiphiles on the

Phase Behavior of Solvent-Bitumen-Water Systems. Energy Fuels 2011, 25, 2223-2231.

50

3.1 ABSTRACT

The impact of asphaltene partitioning on oil-water phase separation was previously

evaluated as a function of solvent-bitumen-water ratios, solvent aromaticity, and temperature in

Chapter 2. In this chapter, the added effect of naphthenic amphiphiles at concentrations of 3

wt% and 10 wt% was assessed. The observed phase behavior of the resulting rag layers was

discussed in view of interfacial co-adsorption mechanisms proposed in the literature. A major

finding was that, under alkaline process conditions, a shift in the rag layer morphology from

water-in-oil (w/o) to oil-in-water (o/w) at increased sodium naphthenate (NaN) concentrations

limited oil-water phase separation as a result of an increase in the surface area to volume ratio of

emulsion droplets and interfacial asphaltene partitioning. Contrary to NaN-free systems, it was

also observed that both temperature and solvent aromaticity had a minimal effect on the phase

behavior of NaN systems. Furthermore, naphthenic acids (NAs) were capable of promoting the

separation of w/o rag layers under acidic formulation conditions.

3.2 INTRODUCTION

A detailed overview of the recent advances in understanding the individual role of

asphaltenes and naphthenic amphiphiles on the formation and stability of water-in-oil (w/o) and

oil-in-water (o/w) rag layers has been provided in Chapters 1 and 2. Their synergistic behavior,

however, still remains relatively unexplored. Insight into this matter for mixtures of asphaltenes

and sodium naphthenates (NaNs) was offered by Wu and Czarnecki using a thermodynamic

modeling approach1. These researchers proposed a bilayer structure to describe the competitive

adsorption of asphaltenes and NaNs at the oil-water interface. In this hypothesized structure,

NaNs occupy the primary adsorbed layer and therefore act to sufficiently reduce the interfacial

51

tension (γow) and promote an o/w rag layer morphology. The makeup of the secondary, or

“floating”, layer is predominantly asphaltenes. Because of the minimal interaction of this layer

with the oil-water interface, it mainly serves to enhance rag layer stability. Elements of this

interfacial model were justified experimentally in a follow-up study by Moran and Czarnecki2.

By comparing the γow isotherms of an aqueous solution of NaNs with a highly diluted bitumen

sample and synthetic solvent (heptol), they observed that NaNs completely displace asphaltenes

from direct oil-water interfacial adsorption at a concentration of 0.1 wt%. Furthermore, using

droplet interaction experiments, these researchers showed that emulsion droplets prepared using

diluted bitumen, instead of heptol, exhibit a more rigid interface because they are able to

withstand coalescence over prolonged contact periods. The nature of asphaltene and naphthenic

acid (NA) films is less well-defined. Upon analyzing their molecular adsorption characteristics,

Varadaraj and Brons concluded that these components likely adsorb at the oil-water interface as

either mixed aggregates or mixed monolayers3. The impact of the resulting surfactant mixture

on the stability of w/o emulsion droplets was assessed indirectly by Poteau et al. via coalescence

studies involving the addition of maltenes (containing NAs) to diluted asphaltene solutions4.

These authors observed that asphaltene and maltene mixtures considerably enhance emulsion

stability. In contrast to these indirect observations, Gao et al. suggested that NAs soften the oil-

water interface and promote emulsion coalescence5.

The aim of this chapter is to address how naphthenic amphiphiles impact the phase

behavior of solvent-bitumen-water rag layers under formulation conditions that are compatible

with bitumen extraction processes. It is hypothesized that the stability of these rag layers is

influenced by the presence of naphthenic amphiphiles when using formulation conditions that

promote their adsorption at the oil-water interface and that the nature of this influence may vary

52

depending on the interaction of the naphthenic amphiphiles with asphaltenes. To evaluate this

influence, oil-water phase separation, asphaltene partitioning, and rag layer properties will be

characterized as a function of the solvent-bitumen-water ratios, solvent aromaticity, pH, and

temperature at different naphthenic amphiphile concentrations. The resulting phase behaviors

will also be discussed in light of interfacial co-adsorption mechanisms currently available in the

literature.

3.3 MATERIALS AND METHODS

3.3.1 Materials

All materials were used as received: anhydrous toluene (product #244511, 99.8%) and

NAs (product #70340, technical-grade extract) were purchased from Sigma-Aldrich Corp.

(Oakville, ON, Canada), reagent grade heptane (product #5400-1-10) and hexane (product

#5500-1-10) were obtained from Caledon Laboratory Chemicals (Georgetown, ON, Canada),

NaNs (90%) were supplied by Eastman Kodak, HCl (6 N, product #CABDH3204-1) and NaOH

(10 N, product #CABDH3247-1) were acquired from VWR (Mississauga, ON, Canada), and

coker feed bitumen (16.3% saturates, 39.8% aromatics, 28.5% resins, and 14.7% asphaltenes

according to SARA analysis) was donated by Syncrude Canada Ltd. (Edmonton, AB, Canada)6.

A similar salt water solution as that prepared in Chapter 2 was used here.

3.3.2 Formulation Preparation

Homogeneous oil phases were prepared by mixing heptol 80/20 (80 vol% heptane and

20 vol% toluene) and heptol 50/50 (50 vol% heptane and 50 vol% toluene) together with

bitumen at the same mass ratios of 1 to 1.5, 1 to 1, 3 to 1, 4 to 1, and 10 to 1 and using the same

53

wrist-action shaking parameters described in Chapter 2. These oil phases were then added to

aqueous solutions consisting of 3 wt% and 10 wt% NaNs at mass ratios of 1 to 10, 1 to 3, 1 to 1,

and 3 to 1 in 15 mL flat-bottom glass centrifuge tubes. Using the previously established batch

emulsification-separation protocol, these formulations were mixed (150 W VWR analog vortex

mixer) at 3200 rpm for 2 minutes and then centrifuged (IEC clinical centrifuge) at 500 g for 1

minute at 25°C. The effect of temperature on oil-water phase separation was evaluated by

placing the above formulations in a hot water bath at 80°C prior to mixing and centrifuging. To

modify the formulation pH, 6 N HCl and 10 N NaOH were added dropwise to the aqueous

phase until the respective titration endpoints of pH 4 and pH 10 were detected using an Oakton

Benchtop pH/ion 510 meter. As a result of the precipitation of NaNs upon HCl addition, NAs

were instead pre-dissolved within the oil phase at concentrations of 3 wt% and 10 wt%.

3.3.3 Microscopy and Material Balances

Optical, cross-polarized, and fluorescent micrographs of the separated oil phase, aqueous

phase, and rag layer of all the formulations in Section 3.3.2 were taken according to the

procedure outlined in Chapter 2. Material balances were also similarly performed with the aid of

these micrographs.

3.3.4 Asphaltene Losses

The UV-vis spectroscopic methodology developed to track asphaltene losses to the rag

layer in Chapter 2 was employed here.

54

3.3.5 Interfacial Tension Measurements

The γow of NaN formulations, where γow<1 mN/m, was measured according to the

procedure for the spinning drop tensiometer manufactured by Temco Inc. (model #500) in

Chapter 2. The only difference in this application was that freshly prepared solutions of NaNs

and diluted bitumen were respectively used as the heavy (aqueous) and light (oil) phases

because of the poor phase separation results. A KSV Sigma 700 tensiometer equipped with a

platinum Du Noüy ring probe was alternatively used to measure the larger γow values (20-26

mN/m) of pH 4 systems. In this technique, the Du Noüy ring was initially immersed within 5

mL of a given oil phase that resided on top of 10 mL of acidified salt water. As the Du Noüy

ring was lowered through the oil-water interface, the resulting force exerted on it (F) was

calculated using the following relationship7:

pullogVF , (Eq. 1)

Here, Δρ is the difference in density between the heavy (water) and light (oil) phases, g

is the gravitational acceleration constant (9.81 m/s2), and Vo,pull is the volume of oil pulled

through the interface. The γow was subsequently determined as follows:

wire

ring

pullo

ring

ring

owr

r

V

rf

r

F,

4 ,

3

(Eq. 2)

Tabulated values of

wire

ring

pullo

ring

r

r

V

rf ,

,

3

, which is a dimensionless quantity that may be

expressed exclusively as a function of the probe’s ring (rring=9.545 mm) and wire (rwire=0.185

mm) radii as well as Vo,pull, are available throughout the literature8,9,10

. The Du Noüy ring was

rinsed with toluene, ethanol, and deionized water prior to reuse. Due to the large volumes of the

55

separated aqueous and oil phases required for this methodology, the effect of water to oil ratios

on γow could not be assessed in situ.

3.3.6 Surface Pressure-Area Isotherms

The effect of NAs on the collapse pressure and elasticity (ε) of asphaltene skins was

tested using a KSV Minitrough operated on a vibration-free table. Initially, asphaltenes were

precipitated from bitumen via hexane dilution at ratios greater than 40 to 1. Recovered

asphaltenes were next rinsed with additional quantities of hexane in order to minimize maltene

contamination11

. Surfactant mixtures of 100% asphaltenes, 75% asphaltenes and 25% NAs, 50%

asphaltenes and 50% NAs, 25% asphaltenes and 75% NAs, and 100% NAs were subsequently

diluted with toluene to a concentration of 0.1%. A volume of 45 µL of each of these mixed

surfactant solutions was then spread onto a subphase composed of pure deionized water with a

surface tension comparable to that of the salt water solution (~71 mN/m). After waiting 10

minutes to allow for solvent evaporation, the area of the spreading phase (Asp) was compressed

symmetrically at a constant rate of 8 mm/minute from 256 cm2 to 23 cm

2 using a pair of

interlinked surface barriers. The change in surface pressure (πp) during compression was

measured using a Wilhelmy plate as follows12

:

op (Eq. 3)

In this equation, σo and σ represent the respective surface tensions in the absence and

presence of surface-active molecules. The collapse pressure of mixed asphaltene and NA films

was interpreted from the resulting πp-Asp isotherms as the maximum attainable πp. Furthermore,

their ε, which is a measure of their resistance against compression, was calculated as follows13

:

56

sp

p

Aln

(Eq. 4)

Each compression test was followed by a cleaning cycle in which all of the

instrumentation was thoroughly rinsed with ethanol and deionized water. The cleanliness of the

subphase was verified by ensuring that πp<0.3 mN/m prior to depositing the spreading phase. In

addition, it was verified by spreading toluene alone onto pure deionized water that the solvent

had a negligible impact on πp. Although the πp-Asp isotherms generated in this study were not a

true reflection of those attainable at the oil-water interface, they provided an accurate depiction

of the trends in changes of the surfactant film stability for controlled asphaltene and NA

mixtures. This point was illustrated in an earlier study by Zhang et al. where similarities in the

trends of πp and interfacial pressure isotherms for asphaltene and demulsifier mixtures were

observed12

.

3.4 RESULTS

3.4.1 Phase Behavior of Naphthenic Amphiphile Systems

The phase behavior results presented in Figures 3.1 (a)-(f) for solvent-bitumen-water

systems composed of 3 wt% NaNs at pH 7.5 are overlaid on top of those for the baseline

systems at 0 wt% NaNs and 9.1 wt% and 50 wt% water taken from Chapter 2.

57

0

10

20

30

40

0 2 4 6 8 10

Asp

halt

en

es L

osses t

o R

ag

Layer

(%)

Heptol 80/20 to Bitumen Dilution Ratio

0

20

40

60

80

100

0 2 4 6 8 10

Oil

Lo

sses t

o R

ag

Layer

(Vo

l %

)

0

20

40

60

80

100

0 2 4 6 8 10

75 wt% Water (3 wt% NaNs) 50 wt% Water (3 wt% NaNs) 25 wt% Water (3 wt% NaNs)

9.1 wt% Water (3 wt% NaNs) 50 wt% Water (Baseline) 9.1 wt% Water (Baseline)

0

10

20

30

40

0 2 4 6 8 10

Heptol 50/50 to Bitumen Dilution Ratio

0

10

20

30

40

0 2 4 6 8 10

0

20

40

60

80

100

0 2 4 6 8 10

(a) (b)

(d)

(f)

(c)

(e)

Figure 3.1: Oil and asphaltene losses to the rag layer for 0 wt% (baseline) and 3 wt% NaN

systems prepared at pH 7.5 with (a and b) heptol 80/20 at 25°C, (c and d) heptol 80/20 at 80°C,

and (e and f) heptol 50/50 at 25°C.

For systems prepared with heptol 80/20 at 25°C (Figures 3.1 (a)), oil losses to the rag

layer are worsened at increased water contents. At a water content of 75 wt%, the entire oil

phase is emulsified within the rag layer. For systems containing ≤50 wt% water, reducing the

solvent to bitumen dilution ratio significantly increases the oil losses to the rag layer. Despite

58

observing similar trends in oil-water phase separation at 80°C (Figure 3.1 (c)) and when using

heptol 50/50 as the solvent (Figure 3.1 (e)), complete emulsification of the oil phase within the

rag layer is observed at a water content ≥50 wt%. As illustrated in Figure 3.2, acidifying

solvent-bitumen-water systems to pH 4 significantly improves oil recovery.

0

10

20

30

40

0 2 4 6 8 10

Oil

Lo

ss

es

to

Ra

g L

aye

r (V

ol %

)

Heptol 80/20 to Bitumen Dilution Ratio

75 wt% Water (3 wt% NAs)

50 wt% Water (3 wt% NAs)

25 wt% Water (3 wt% NAs)

9.1 wt% Water (3 wt% NAs)

50 wt% Water (Baseline)

9.1 wt% Water (Baseline)

Figure 3.2: Oil losses to the rag layer for 0 wt% (baseline) and 3 wt% NA systems as a function

of the heptol 80/20 to bitumen dilution ratio and water content. All systems are evaluated at pH

4 and 25°C.

With regards to the partitioning behavior of asphaltenes, Figures 3.1 (b) and (d) show

that an increase in the heptol 80/20 to bitumen dilution ratio promotes an increase in the

asphaltene losses to the rag layer. Whereas asphaltene losses are independent of temperature,

they are notably lower for heptol 50/50 systems (Figure 3.1 (f)) and decrease with increased

solvent to bitumen dilution ratios. In all cases, the presence of 3 wt% NaNs induces an increase

in asphaltene losses relative to the baseline formulations. Asphaltene losses are not evaluated for

systems at pH 4. Here, the presence of NAs changed the visible absorbance spectrum of diluted

59

bitumen. A similar shift observed by Östlund et al. in the near-IR spectrum of bitumen and NA

mixtures supports the idea of asphaltene and NA interactions14

.

3.4.2 Transitions to the Rag Layer Morphology

Changes to the rag layer morphology of solvent-bitumen-water systems produced upon

increasing the naphthenic amphiphile concentration from 0 wt% to 10 wt% are illustrated in

Figures 3.3 (a)-(c).

3 wt% NaNs 10 wt% NaNs

0 wt% NAs

(a)

Multiple Emulsions

0 wt% NaNs

250 µm

3 wt% NAs 10 wt% NAs(b)

10 wt% NAs3 wt% NAs0 wt% NAs(c)

Figure 3.3: Fluorescent micrographs of the effect of (a) NaNs (pH 7.5) and (b) NAs (pH 4) on

the morphology of heptol 80/20-bitumen-water rag layers. (c) Cross-polarized images of rag

layers containing NAs at pH 4.

For all NaN systems at pH ≥7.5, an increase in the surfactant concentration results in a

transition of the rag layer from w/o to o/w. This transition is accompanied by a major decrease

in the droplet size from 0 wt% to 3 wt% NaNs, whereas such changes are less dramatic from 3

60

wt% to 10 wt% NaNs. At pH 4, where NAs are dominant, a w/o rag layer is continuously

observed despite an increase in the surfactant concentration. Traces of multiple emulsions also

become evident. Of special interest for systems prepared at pH 4 is the increased observance of

liquid crystals at larger NA concentrations. These phases, which appear as bright spots in Figure

3.3 (c), do not form continuous films adsorbed at the oil-water interface as has been suggested

by other researchers15

. Instead, they appear as aggregates dispersed throughout the oil phase.

Differences in the resulting liquid crystal structures may be attributed to variations in the

emulsification protocol.

3.4.3 Interfacial Tension Isotherms

Baseline γow isotherms as well as those obtained for naphthenic amphiphile systems are

presented in Figures 3.4 (a) and (b).

0

10

20

30

40

0 2 4 6 8 10

γo

w(m

N/m

)

Heptol 80/20 to Bitumen Dilution Ratio

3 wt% NaN/NA 10 wt% NaN/NA 50 wt% Water (Baseline)

9.1 wt% Water (Baseline) pH 4 (Baseline)

0

10

20

30

40

0 2 4 6 8 10

(a) (b)

Figure 3.4: Baseline γow isotherms as well as those for (a) NaN systems at pH 7.5 and (b) NA

systems at pH 4 as a function of the heptol 80/20 to bitumen dilution ratio. The temperature is

maintained at 25°C.

61

At increased heptol 80/20 to bitumen dilution ratios, it is observed that NaNs

significantly reduce γow from 5-10 mN/m to 0.5-0.9 mN/m. Introducing NAs into similar

formulations at pH 4 results in an increase in γow from ~20-22 mN/m to ~23-26 mN/m. In both

of the above scenarios, increasing the naphthenic amphiphile concentration from 3 wt% to 10

wt% results in negligible changes to γow. Furthermore, varying the formulation temperature and

solvent aromaticity has little influence on all of the above isotherms. Equilibrium γow values are

obtained almost instantaneously in all of the above systems (on the order of seconds). This

finding is in agreement with that of Moran and Czarnecki who suggested that all dynamic

behavior is lost at such large asphaltene and naphthenic amphiphile concentrations because of

their relatively fast adsorption kinetics and high interfacial packing2.

3.4.4 Impact of Naphthenic Acids on Asphaltene Film Properties

The effect of NAs on the collapse pressure and ε of asphaltene skins may be interpreted

from the πp-Asp isotherms illustrated in Figure 3.5.

0

15

30

45

60

75

0 1 2 3 4 5 6

100% A

75% A-25% NAs

25% A-75% NAs

100% NAs

ln (Asp (cm2))

πp

(mN

/m)

50% A-50% NAs

Figure 3.5: πp-Asp isotherms of asphaltene (A) and NA surfactant mixtures.

62

According to these results, pure asphaltenes produce more stable films than pure NAs.

This is indicated by their larger collapse pressure. Mixtures of these components show an

intermediate behavior where the film stability is reduced at increased NA concentrations. As

shown in Table 3.1, such mixtures are a good representation of the spread in the relative bulk

phase asphaltene and NA compositions tested.

Table 3.1: Relative asphaltene and NA compositions in formulations tested.

Heptol 80/20 to Bitumen 3 wt% NAs 10 wt% NAs

% Asphaltenes % NAs % Asphaltenes % NAs

1 to 1.5 78 22 52 48

1 to 1 75 25 47 53

3 to 1 60 40 31 69

4 to 1 55 45 26 74

10 to 1 35 65 14 86

Values of ε for produced asphaltene and NA films are calculated at πp values ranging

from 0 mN/m to 10 mN/m as it is the common linear regime amongst all generated πp-Asp

isotherms. The results presented in Table 3.2 suggest that all such films are prone to

deformation. It is interesting to note that the ε of mixed asphaltene and NA films in the NA-rich

domain is lower than that of pure NAs.

Table 3.2: Measurements of ε for mixed asphaltene and NA films.

System ε (mN/m)

100% Asphaltenes 73

75% Asphaltenes and 25% NAs 34

50% Asphaltenes and 50% NAs 16

25% Asphaltenes and 75% NAs 16

100% NAs 23

The πp-Asp isotherms for NaN systems are also conducted by first spreading 45 µL of a

0.1% asphaltene solution onto deionized water. A concentrated NaN solution is then infused

within the subphase such that its final concentrations tested are 0.1%, 1%, and 3%. The resulting

63

measured collapse pressures are 25 mN/m, 9 mN/m, and 6 mN/m respectively. Although this

data illustrates the ability of NaNs to significantly weaken asphaltene films, it cannot be

extrapolated to phase behavior studies because a liquid-liquid trough is required to permit for

the formation of a secondary asphaltene layer on top of the primary adsorbed NaN layer. Gao et

al. successfully implemented the above recommendation for interfacial films composed of

asphaltenes and NaNs in a 1 to 1 volume ratio5. The resulting πp-Asp isotherm of the bilayer

structure showed an intermediate behavior compared to pure asphaltenes and pure NaNs with a

collapse pressure >30 mN/m.

3.5 DISCUSSIONS

3.5.1 Interfacial Co-Adsorption of Asphaltenes and Sodium Naphthenates

In a comparison of the asphaltene-controlled (baseline) and NaN-controlled phase

behavior results in Figures 3.1 (a)-(f), trends depicting increased oil losses to the rag layer at

increased water to oil ratios are maintained in addition to the partitioning behavior of

asphaltenes. Of special interest, however, is the increased magnitude of oil losses observed in

NaN systems. To understand this behavior, the effect of NaN concentration on emulsification

must first be evaluated. The first step in establishing this link is to discuss the relationship

between the γow and average emulsion droplet diameter ( E

dd )16

:

ow

E

dd (Eq. 5)

In this equation, α is 1 for laminar flow and 0.6 for turbulent flow. According to this

relationship, any reduction in γow produces a reduction in the drop size. The change in E

dd also

changes the surface area ( E

dA ) to volume ( E

dV ) ratio of emulsion droplets as follows:

64

E

d

E

d

E

d

dV

A 6 (Eq. 6)

From the above equations, it is apparent that γow is inversely related to the E

dA to E

dV

ratio of a given emulsion droplet. Figure 3.6 illustrates measured changes in γow for a given

diluted bitumen oil phase as a function of the added NaN concentration to the aqueous phase.

0

1

2

3

4

5

6

7

0 1 2 3NaN Concentration (wt%)

y1 = -6.20x1 + 7.38

y2 = -0.32x2 + 1.68

CMCNaN

γo

w(m

N/m

)

Figure 3.6: γow of bitumen diluted with heptol 80/20 versus the added NaN concentration to the

aqueous phase at pH 7.5 and 25°C. The included fluorescent micrographs show a transition in

the rag layer morphology from w/o to o/w at the CMC of NaNs (~1 wt%).

According to this data, γow decreases linearly up to a NaN concentration of 1 wt%

(Δγow/ΔNaN~-6.2 (mN/m)/wt% NaNs). By holding all parameters constant in Equation 5 other

than γow, Equation 6 reveals a potential increase in the E

dA to E

dV ratio of emulsion droplets by a

factor of 3 upon increasing the NaN concentration from 0 wt% to 1 wt%. Emulsion stabilization

is facilitated at such large oil-water interfacial areas (Aow) as the potential for asphaltene

adsorption is enhanced. Figure 3.1 (b) shows the increase in asphaltene losses to the rag layer.

65

At NaN concentrations >1 wt%, changes to the E

dA to E

dV ratio of emulsion droplets are less

pronounced (Δγow/ΔNaN~-0.3 (mN/m)/wt% NaNs). Furthermore, it is here where the dominant

morphology of the emulsion switches from w/o to o/w. This transition coincides with the fact

that the critical micelle concentration (CMC) of NaNs is 1 wt%. A similar CMC value was

previously reported by Moran and Czarnecki2. Beyond this concentration, the spontaneous

increase in the total number of dispersed droplets further increases the overall Aow and

asphaltene losses to the rag layer. Under interfacial saturation conditions, which occurs at NaN

concentrations >1 wt%, the bilayer model proposed by Wu and Czarnecki is believed to

uphold1. Here, asphaltene skins must adsorb as a secondary layer at the oil-water interface

because NaNs predominantly occupy the primary adsorbed layer. A modified schematic of this

model is provided in Figure 3. 7 (a).

(a)

WaterNaNs

Asphaltenes Oil

i) Mixed Monolayer ii) Mixed Aggregate

NAsAsphaltenes

Oil

Water

Mixed

aggregate

(b)

Figure 3.7: (a) Bilayer model proposed by Wu and Czarnecki for the interfacial co-adsorption

of asphaltenes and NaNs at the oil-water interface1. (b) Co-adsorption mechanisms proposed by

Varadaraj and Brons for asphaltenes and NAs at the oil-water interface include (i) mixed

monolayers and (ii) mixed aggregates3.

66

Of the mechanisms available for increasing Aow, an increase in the E

dA to E

dV ratio of

emulsion droplets is the primary factor leading to larger oil losses to the rag layer. This is

supported by the data in Figure 3.1 (a) for systems prepared with 0 wt% and 3 wt% NaNs.

Contributions from spontaneous emulsification are less significant because changes in oil losses

to the rag layer for systems prepared with 3 wt% and 10 wt% NaNs are marginal. It should be

noted that Δγow/ΔNaN in Figure 3.6 below the CMC of NaNs is impacted significantly by

formulation conditions. This is concluded from Figure 3.4 (a) for systems prepared with 0 wt%

NaNs where γow is observed to vary according to changes in the heptol 80/20 to bitumen dilution

ratio and water content as a result of modifications to the asphaltene partitioning within the rag

layer. Also depicted in Figure 3.4 (a) is that the impact of the above variables on γow at NaN

concentrations >1 wt% is less prominent.

The decrease in oil losses to the rag layer at the large bitumen dilution ratios described in

Figure 3.1 (a) for 3 wt% NaN systems at low water contents opposes the theoretical principles

relating asphaltene losses to emulsion stability for the NaN-free systems in Chapter 2. In the

presence of NaNs, however, the ratio of asphaltenes to NaNs decreases with an increase in the

dilution ratio, thus producing thinner, or “softer”, asphaltene skins. Emulsion coalescence

studies conducted by Gao et al. support this conclusion5. These researchers showed that a

critical asphaltene to NaN ratio exists below which asphaltene skins can no longer mask the

softening effect of NaNs and emulsion destabilization is promoted.

67

3.5.2 Effect of Temperature and Solvent Aromaticity on Oil Recovery from Rag

Layers

A more realistic depiction of the phase behavior of crude oil formulations that is in line

with actual processing conditions is to evaluate oil recovery from rag layers at 80°C. As was

previously outlined for baseline systems, increasing the temperature from 25°C to 80°C helps to

reduce oil losses to the rag layer from a maximum of 30 vol% to 10 vol% despite having a

negligible influence on the asphaltene partitioning characteristics. This effect is a direct

consequence of the rag layer being composed of water droplets dispersed throughout a

continuous oil phase. The reduced viscosity of the continuous oil phase at elevated temperatures

facilitates its drainage and hence overall recovery. For 3 wt% (and 10 wt%) NaN systems,

increasing the temperature from 25°C to 80°C is of negligible benefit to oil-water phase

separation because the γow isotherm presented in Figure 3.4 (a) remains unchanged. This

observation is supported by Acosta et al. who found that the temperature has little influence on

the interfacial activity of ionic surfactants17

. As a result, the hydrophilic nature of NaNs

continues to induce a transition in the rag layer morphology from w/o to o/w and therefore the

drainage of the aqueous phase is instead facilitated. It should be emphasized for NaN

formulations that asphaltene losses to the rag layer remain independent of the temperature.

The effect of solvent aromaticity on rag layer stability is tested by decreasing heptol’s

heptane to toluene ratio from 80/20 to 50/50. To understand the effect of such a change on the

observed phase behavior of baseline systems, the concept of critical transition concentration

(CTC) in Chapter 2 should be reviewed. In short, the CTC refers to the solvent to bitumen

dilution ratio below which the surface activity of asphaltene aggregates may be associated with

an increased tendency to stabilize rag layers. From γow measurements, the CTC of heptol 80/20

68

and heptol 50/50 systems is observed to correspond respectively to solvent to bitumen dilution

ratios of 10 and 2-3. In the presence of 3 wt% (and 10 wt%) NaNs, the effect of solvent

aromaticity becomes less relevant. As illustrated in Figures 3.1 (a) and (e) for the solvents

heptol 80/20 and heptol 50/50, respectively, oil losses to the rag layer are quite similar. This is a

result of emulsification being governed by NaNs, which are water-soluble molecules whose

surface activity is independent of the solvent type. It is surprising that, despite the reduced

surface activity of asphaltenes in heptol 50/50 systems (Figure 3.1 (f)), sufficient surface

coverage is still achieved to promote emulsion stability. The interfacial co-adsorption model

proposed by Wu and Czarnecki is a useful tool to understand this phenomenon1. These authors

speculated that the secondary adsorbed asphaltene layer is predominantly anchored into place

via polar point contacts with the oil-water interface instead of cross-linking with the primary

adsorbed NaN layer. Therefore, the smaller-sized asphaltene aggregates adsorb more efficiently

at the secondary layer of the interface. Such efficient adsorption promotes larger oil losses to the

rag layer for systems prepared with an aqueous phase content of 50 wt%.

3.5.3 Effect of pH on Oil Recovery from Rag Layers

As previously reported in Chapter 1, the pKa of naphthenic amphiphiles is ~6.

Increasing the pH of solvent-bitumen-water formulations containing 3 wt% NaNs from pH 7.5

to pH 10 is therefore expected to have no substantial impact on the observed phase behavior

because the ionized surfactant already exists in its dissociated state. Although the results at pH

10 are not presented, this hypothesis can indeed be experimentally validated using the

procedures outlined in Section 3.3.

69

As illustrated in Figure 3.2, reducing the pH of baseline solvent-bitumen-water

formulations results in improved oil-water phase separation. This behavior is unexpected

because asphaltenes should show an increased surface activity under such conditions as a result

of the protonation of their basic function groups, such as amines18

. Because of similarities in the

γow isotherms (Figure 3.4 (b)) and oil losses to the rag layer (Figure 3.2) for 0 wt% and 3 wt%

NA systems at pH 4, it is suspected that a considerable fraction of endogenous NAs remain in

the coker feed bitumen that, at low pH, are less active at the oil-water interface, promoting

higher γow and improved oil-water phase separation. This theory may also explain why lower γow

values are observed for solvent-bitumen-water systems under neutral conditions compared to

other values published in the literature2,5

. Although it cannot be confirmed from the air-liquid

πp-Asp isotherms presented in Figure 3.5, it is likely that NAs help to destabilize emulsion

droplets by weakening the asphaltene films adsorbed at the oil-water interface. The

experimental observations recently published by Gao et al. hint at a similar softening effect5.

Conflicting results obtained by Poteau et al. are likely a result of constituents other than NAs

present within the maltene solution added to the oil phase4. The possible mechanisms

highlighted by Varadaraj and Brons to describe the co-adsorption of asphaltenes and NAs at the

oil-water interface are provided in Figure 3.7 (b).

3.6 CONCLUSIONS

The results and ensuing discussions revealed that, consistent with the initial hypothesis,

the presence of naphthenic amphiphiles influenced the separation of solvent-bitumen-water rag

layers in formulations that promoted the adsorption of these species. Formulations with high

water content and low interfacial tensions (γow) produced smaller droplet sizes and increased oil-

water interfacial areas during emulsification, which magnified the effect of naphthenic

70

amphiphiles. For sodium naphthenates (NaNs), stable oil-in-water rag layers were produced in

formulations with high asphaltene and NaN contents (i.e. high bitumen and water contents). The

increased asphaltene losses in the presence of NaNs suggested that NaNs improved the

segregation of asphaltenes at the oil-water interface. Although indirect evidence suggested that

NaNs produced softer films, this apparent softening effect did not compensate for the decreased

γow and droplet size. In acidic environments, undissociated naphthenic acids (NAs) led to higher

γow that facilitated the separation of water-in-oil rag layers. Once again, indirect evidence at the

air-liquid interface suggested that NAs softened asphaltene films. This observation was

consistent with the improved separation obtained in NA-containing systems at pH 4 compared

with baseline systems at pH 7.5. Interfacial co-adsorption mechanisms proposed in the literature

were compatible with the results highlighted above. Although results obtained under acidic

conditions showed improved oil recovery, further studies are required to assess how NAs

modify the properties of asphaltene films adsorbed at the oil-water interface.

3.7 REFERENCES

1 Wu, X.A.; Czarnecki, J. Modeling of Diluted Bitumen-Water Interfacial Compositions using a

Thermodynamic Approach. Energy Fuels 2005, 19, 1353-1359.

2 Moran, K.; Czarnecki, J. Competitive Adsorption of Sodium Naphthenates and Naturally

Occurring Species at Water-in-Crude Oil Emulsion Droplet Surfaces. Colloids Surf., A 2007,

292, 87-98.

71

3 Varadaraj, R.; Brons, C. Molecular Origins of Heavy Crude Oil Interfacial Activity. Part 2:

Fundamental Interfacial Properties of Model Naphthenic Acids and Naphthenic Acids Separated

from Heavy Crude Oils. Energy Fuels 2007, 21, 199-204.

4 Poteau, S.; Argillier, J.-F.; Langevin, D.; Pincet, F.; Perez, E. Influence of pH on Stability and

Dynamic Properties of Asphaltenes and Other Amphiphilic Molecules at the Oil-Water

Interface. Energy Fuels 2005, 19, 1337-1341.

5 Gao, S.; Moran, K.; Xu, Z.; Masliyah, J. Role of Naphthenic Acids in Stabilizing Water-in-

Diluted Model Oil Emulsions. J. Phys. Chem. B 2010, 114, 7710-7718.

6 Akbarzadeh, K.; Alboudwarej, H.; Svrcek, W.Y.; Yarranton, H.W. A Generalized Regular

Solution Model for Asphaltene Precipitation from n-Alkane Diluted Heavy Oils and Bitumens.

Fluid Phase Equilib. 2005, 232, 159-170.

7 Huh, C.; Mason, S.G. A Rigorous Theory of Ring Tensiometry. Colloid Polym Sci. 1975, 253,

566-580.

8 Harkins, W.D.; Jordan, H.F. A Method for the Determination of Surface and Interfacial

Tension from the Maximum Pull on a Ring. J. Am. Chem. Soc. 1930, 52, 1751-1772.

9 Zuidema, H.H.; Waters, G.W. Ring Method for the Determination of Interfacial Tension. Ind.

Eng. Chem. 1941, 13, 312-313.

72

10 Fox, H.W.; Chrisman Jr., C.H. The Ring Method of Measuring Surface Tension for Liquids

of High Density and Low Surface Tension. J. Phys. Chem. 1952, 56, 284-287.

11 Yang, X.; Hamza, H.; Czarnecki, J. Investigation of Subfractions of Athabasca Asphaltenes

and Their Role in Emulsion Stability. Energy Fuels 2004, 18, 770-777.

12 Zhang, L.Y.; Xu, Z.; Masliyah, J.H. Langmuir and Langmuir-Blodgett Films of Mixed

Asphaltene and a Demulsifier. Langmuir 2003, 19, 9730-9741.

13 Saad, S.M.I.; Policova, Z.; Acosta, E.J.; Hair, M.L.; Neumann, A.W. Mixed DPPC/DPPG

Monolayers at Very High Compression. Langmuir 2009, 25, 10907-10912.

14 Östlund, J.-A.; Nydén, M.; Auflem, I.-H.; Sjöblom, J. Interactions between Asphaltenes and

Naphthenic Acids. Energy Fuels 2003, 17, 113-119.

15 Horváth-Szabó, G.; Czarnecki, J.; Masliyah, J.H. Sandwich Structures at Oil-Water Interfaces

under Alkaline Conditions. J. Colloid Interface Sci. 2002, 253, 427-434.

16 Walstra, P. Principles of Emulsion Formation. Chem. Eng. Sci. 1993, 48, 333-349.

17 Acosta, E.J.; Yuan, J.S.; Bhakta, A.S. The Characteristic Curvature of Ionic Surfactants. J.

Surfactants Deterg. 2008, 11, 145-158.

18 Aguilera, B.M.; Delgado, J.G.; Cárdenas, A.L. Water-in-Oil Emulsions Stabilized by

Asphaltenes Obtained from Venezuelan Crude Oils. J. Dispersion Sci. Technol. 2010, 31, 359-

363.

73

CHAPTER 4:

EVALUATING THE HYDROPHILIC-LIPOPHILIC NATURE

OF ASPHALTENIC OILS AND NAPHTHENIC AMPHIPHILES

USING MICROEMULSION MODELS

This chapter is derived from the following published manuscript:

Kiran, S.K.; Acosta, E.J; Moran, K. Evaluating the Hydrophilic-Lipophilic Nature of

Asphaltenic Oils and Naphthenic Amphiphiles Using Microemulsion Models. J. Colloid

Interface Sci. 2009, 336, 304-313.

74

4.1 ABSTRACT

In this chapter, microemulsion (µE) phase behavior models were applied to quantify the

hydrophilic-lipophilic nature of asphaltenic oils (bitumen, deasphalted bitumen, asphalt, and

naphthalene) and surface-active asphaltene aggregates and naphthenic amphiphiles. For the test

oils, the equivalent alkane carbon number (EACN) was determined by evaluating the salinity

shifts of µEs formulated with a reference surfactant (sodium dihexyl sulfosuccinate (SDHS))

and a reference oil (toluene) at different volume fractions. Similarly, the characteristic curvature

(Cc) of test surfactants was determined by evaluating their salinity shifts in mixtures with SDHS

at different molar fractions. As a part of the oil phase, asphaltenes and asphaltene-like species

were highly hydrophilic. This led to low EACN values despite their large molecular weight. As

a surfactant, asphaltenes were hydrophobic species that led to the formation of water-in-oil

emulsions. Naphthenic amphiphiles, particularly sodium naphthenates, were on the other hand

highly hydrophilic compounds that led to the formation of oil-in-water emulsions. These

hydrophilic-lipophilic characterization parameters, and the methods used to determine them,

could potentially be used in the future to understand the phase behavior of complex oil-water

systems (such as rag layers).

4.2 INTRODUCTION

A common methodology for determining the hydrophilic-lipophilic nature of different

oils and surfactants is to analyze the impact of these compounds on the observed microemulsion

(µE) phase behavior. The advantage of using the µE phase behavior as a “hydrophilic-lipophilic

scale” stems from the fact that these are systems in thermodynamic equilibrium. Therefore, any

parameter obtained from these studies reflects the balance of surfactant, oil, and water

75

interactions1. As was described in Chapter 1, the three basic µE types are: oil-in-water (o/w)

µEs in equilibrium with an excess oil phase (Type I), bicontinuous µEs in equilibrium with

excess oil and aqueous phases (Type III), and water-in-oil (w/o) µEs in equilibrium with an

excess aqueous phase (Type II). For µEs formulated with an ionic surfactant, an increase in the

electrolyte concentration leads to a Type IType IIIType II µE phase behavior transition. In

other words, increasing the salinity favors the partitioning of the surfactant into the oil phase.

Along this transition, the interfacial tension (γow) of the system changes significantly, reaching

an ultralow value at Type III conditions where the net curvature of the surfactant at the oil-water

interface is 0 (see Figure 4.1)2. This specific Type III condition is referred to as the “optimum”

point where one also typically finds equal volumes of solubilized oil and water3,4,5

. The

electrolyte concentration needed to obtain a net 0 curvature is the optimal salinity (S*).

From the above discussion, one can extrapolate that a larger salt concentration is

required to produce a Type IType IIIType II µE phase behavior transition in systems

formulated with a more hydrophobic oil as it is more difficult for the surfactant to partition into

such an oil. Similarly, a larger salt concentration is required to produce a Type IType

IIIType II µE phase behavior transition in systems formulated with a more hydrophilic

surfactant as it is highly soluble in water and resists leaving the aqueous phase. In this chapter,

the plan is to use these principles to evaluate the hydrophilic-lipophilic nature of asphaltenic oils

(bitumen, deasphalted bitumen, asphalt, and naphthalene) and surface-active asphaltene

aggregates, naphthenic acids (NAs), and sodium naphthenates (NaNs).

Unfortunately, these oils and surfactants do not form µEs on their own. As a result, µE

formation is facilitated by mixing each test oil and surfactant with a reference phase. The

reference oil and anionic surfactant used in this study is toluene and sodium dihexyl

76

sulfosuccinate (SDHS) respectively. By determining the shift in S* as a function of the fraction

of the test oil or surfactant in the mixture, the relative hydrophobicity of the test oils and

surfactants will be determined.

The hydrophilic-lipophilic deviation (HLD) model is used in this study to quantify the

relative impact of the hydrophilic-lipophilic nature of the above specified test oils and

surfactants on the µE phase behavior. The HLD equation developed by Salager et al. for µEs

containing an ionic surfactant, or mixture of ionic surfactants, is as follows1,2,6,7,8

:

cToc CTaAfKNSHLD ,ln (Eq. 1)

In this equation, the natural logarithm ln(S) describes the effect of the aqueous phase salt

concentration (S, in g/100 mL) on the suppression of the ionic surfactant’s electrical double

layer. The next term KNc,o reflects the ability of the surfactant to induce a dipole-dipole

interaction with the oil phase. An average K value of 0.17 is typically applied to most

surfactant-oil combinations. On the other hand, Nc,o is an explicit measure of the oil phase

hydrophobicity. For basic alkanes, Nc,o is simply equal to the number of oil molecule carbon

atoms and is thus commonly referred to as the alkane carbon number (ACN). For more complex

oils, Nc,o, which needs to be solved for experimentally, is instead referred to as the equivalent

alkane carbon number (EACN). The additional function f(A) varies with the co-surfactant type

and concentration. In the event where no co-surfactant is added, f(A) is set to 0. Furthermore,

the term ΔT is the difference in temperature from 298 K. An associated pre-factor value of

aT=0.01 K-1

is commonly reported in the literature. The final characteristic curvature term, Cc, is

indicative of the hydrophilic-lipophilic nature of the surfactant of surfactant mixture.

77

The effect of varying a specific parameter (e.g. salinity) while keeping all others

constant in Equation 1 is explored with respect to the phase behavior of SDHS-oil (20 wt%

naphthalene and 80 wt% toluene)-water µEs in Figure 4.1.

γo

w (

mN

/m)

NaCl (g/100 mL)

Figure 4.1: The phase behavior and corresponding γow of SDHS-oil (20 wt% naphthalene and

80 wt% toluene)-water µEs as a function of the salinity.

According to this figure, a HLD<0HLD=0HLD>0 shift corresponds to a Type

IType IIIType II µE phase behavior transition. The point of optimum formulation is at

HLD=0.

Limited data is currently available on the use of the HLD equation to characterize crude

oil components as most efforts have been focused on characterizing contaminants, oils used in

cosmetic and pharmaceutical formulations, and vegetable oils7,9,10,11

. Rondón et al. made the

first attempt to relate the HLD concept to the phase behavior of crude oil formulations by

adjusting the concentration of added demulsifier such that an ultralow γow is observed at

HLD=012

. They, however, fell short of obtaining both the EACN of the oil and the “apparent”

78

characteristic curvature ( a

cC ) of asphaltenes. In addition, these researchers did not consider the

role of naphthenic amphiphiles in their system. It should be clarified that the qualifier

“apparent” is used to reflect the fact that asphaltenes are better classified as a solubility class

rather than a pure component or well-defined mixture.

One of the most complete sets of data used to determine the EACN of a wide range of

oils was produced by Baran Jr. et al.13

. These researchers used the phase behavior of µEs

produced with SDHS at 25°C, without alcohols or other co-surfactants, to obtain the value of S*

for oils with different ACNs ranging from 6 to 10. The following linear correlation was

established to relate these parameters:

92.017.0ln * ACNS (Eq. 2)

From the measured value of S* for non-alkane test oils, these researchers were able to

employ this equation to solve for their EACN (e.g. the EACN of toluene is 1). By comparing

Equation 2 with Equation 1 for optimum formulations (HLD=0) under similar processing

conditions as those used by Baran Jr. et al. (f(A)=0 and ΔT=0), one may conclude that the Cc of

SDHS=-0.92. It should be noted that large EACN values represent hydrophobic oils whereas

negative Cc (or a

cC ) values represent hydrophilic surfactants.

Baran Jr. et al. also confirmed that the EACN of oil mixture (EACNmix) follows a linear

mixing rule13

:

iimix EACNxEACN (Eq. 3)

79

In this equation, xi and EACNi represent the fraction and EACN of component i

respectively. This equation will be made use of later to estimate the EACN of the test oils in

mixtures with toluene.

Recently, Acosta et al. determined the Cc of various surfactants, including that of NaNs

obtained from Eastman Kodak8. The solved for Cc is -2.4 (highly hydrophilic). This finding is

consistent with the fact that NaNs are water-soluble and therefore tend to produce o/w

emulsions as in Chapter 3.

In the same work of Acosta et al., the Cc of ionic surfactant mixtures (Cc,mix) was shown

to follow a linear mixing rule as well8:

icimixc CxC ,, (Eq. 4)

In this chapter, S* is determined for µEs formulated with mixtures of asphaltenes, NAs,

and NaNs with SDHS at 25°C (ΔT=0) and in the absence of other co-surfactants (f(A)=0). To

obtain Cc,mix, Equation 1 is used with the aforementioned conditions.

The significances of the EACN obtained for asphaltenic oils, the a

cC obtained for

asphaltenes, and the Cc obtained for naphthenic amphiphiles are discussed in light of the EACN

and Cc of other oils and surfactants. It is also later presented in Chapter 8 how these parameters

can be used to interpret the formation and stability of rag layers. The solved for EACN and Cc

values presented in this chapter will vary according to the source of the oils, except for

naphthalene, and surfactants analyzed due to their complexity and ill-defined makeup.

80

4.3 MATERIALS AND METHODS

4.3.1 Materials

The following chemicals were used as purchased from Sigma-Aldrich Corp. (Oakville,

ON, Canada): anhydrous toluene (product #244511, 99.8%), anhydrous hexadecane (product

#296317, ≥99%), anhydrous hexane (product #296090, 95%), naphthalene (product #147141,

99%), NAs (product #70340, technical-grade extract), SDHS (product #86146, ~80% in water),

and NaCl (product #S9625, ≥99.5%). The purchased NAs are a technical grade extract from

crude oil (no defined origin) with an acid number of 230 g/equivalent, which is used here as its

molecular weight14,15

. Recent studies have shown that naphthenic amphiphiles obtained from

Alberta’s oil sands have a similar composition to those found in this sample and are enriched in

C12-C16 fatty acid equivalents with 3-6 double bond equivalents16

. The calculated Cc values of

naphthenic amphiphiles presented later are reflective of the wide distribution of such surfactants

found naturally within Canadian crude oil reserves. NaOH (product #LC24270-5, 0.1 N) was

purchased from Caledon Laboratory Chemicals (Georgetown, ON, Canada), black asphalt

(Dominion Sure Seal Ltd.) was purchased from a local retailer (Toronto, ON, Canada), and

coker feed bitumen (16.3% saturates, 39.8% aromatics, 28.5% resins, and 14.7% asphaltenes

according to SARA analysis) was provided by Syncrude Canada Ltd. (Edmonton, AB,

Canada)17

. Tap water was deionized using an APS Ultra mixed bed resin to a conductivity <3

µs/cm.

4.3.2 Asphaltene Precipitation

The same procedure as that used to precipitate asphaltenes in Chapter 3 was also

employed here.

81

4.3.3 Formulation of Microemulsions with Test Oil and Toluene Mixtures

Each of the test oils was mixed with toluene using a wrist-action shaker until a

homogeneous oil phase was obtained. The concentration of all the test oils (other than

naphthalene in toluene) ranged from 10 to 50 wt%. The concentration of naphthalene was

limited to just 30 wt% due to its lower solubility18

. 5 mL aliquots of these oil mixtures were

mixed together with 5 mL aqueous phase solutions composed of 0.1 M SDHS and 1-10 g

NaCl/100 mL in 15 mL flat bottom glass centrifuge tubes. These mixtures were allowed to

equilibrate for 1 week.

4.3.4 Formulation of Microemulsions with Test Surfactant and SDHS Mixtures

A similar formulation procedure as that described in Section 4.3.3 was used here. The

major differences were that pure toluene was used as the oil phase and the surfactant was a 0.1

M mixture of the test surfactant and SDHS. In addition, the volumes of the added aqueous and

oil phases were 5 mL and 2 mL respectively. The purpose of only using 2 mL of toluene instead

of 5 mL here was to reduce its consumption. For such systems at a relatively low surfactant

concentration, it is not expected that a 2.5 to 1 water to oil volume ratio will affect the µE phase

behavior19

. This was indeed confirmed by repeating the µE phase behavior of only select

systems using a 1 to 1 water to oil volume ratio. Asphaltenes and NAs, being oil soluble, were

diluted with toluene prior to being added to the formulations. The final compositions of the test

surfactant analyzed in the surfactant mixtures were 10 mol%, 30 mol%, 50 mol%, and 70 mol%.

In the case of asphaltene-SDHS mixtures, the surfactant mixture compositions were

calculated by assuming an average molecular weight of 600 g/mol for asphaltenes. This

assumption is based on the observation of Pomerantz et al. who analyzed the molecular weight

82

distribution of asphaltenes in similar bitumen via two-step laser mass spectroscopy20

. To

produce NaNs, NAs were neutralized with the appropriate amount of NaOH. A salinity scan was

performed on each surfactant mixture to determine S*.

4.3.5 Interfacial Tension Measurements

The optimum formulation along each of the performed salinity scans for the various test

oil and test surfactant mixtures was pinpointed by detecting the minimum γow according to the

spinning drop tensiometer procedure outlined in Chapter 2. For Type III µEs, γow was measured

between the excess oil and aqueous phases.

4.3.6 Asphaltene Partitioning at the Oil-Water Interface

As has been described before (and will be reinforced later), asphaltenes can be treated as

both an oil phase component and surfactant. To determine the fraction of asphaltenes that adsorb

at the oil-water interface, the concentration of asphaltenes in the bulk oil phase before and after

producing µEs was assessed using the same UV-vis spectroscopic technique developed in

Chapter 2. It was also confirmed using this technique that the asphaltenes content of the

prepared deasphalted bitumen in Section 4.3.2 was <10% of the original asphaltenes content.

4.4 RESULTS

4.4.1 Microemulsion Phase Behavior Scans

A Type IType IIIType II µE phase behavior transition is observed in all of the

salinity scans. In addition to the example of these transitions already presented in Figure 4.1,

other examples for formulations composed of a bitumen and toluene mixture as the oil phase

83

and a NAs and SDHS mixture as the surfactant are presented in Figures 4.2 (a) and (b). In the

case of Figure 4.2 (a), a bicontinuous middle phase µE is obtained at 3 g NaCl/100 mL. Due to

the high SDHS concentration in that phase, however, and the fact that bitumen and SDHS have

relatively high densities, it sinks to the bottom of the vial. To accurately determine S* for these

systems and others, the γow is measured as a function of the salinity, using smaller salt

increments of 0.2 g NaCl/100 mL in the vicinity of the Type III µEs.

(b) (a)

Figure 4.2: µE phase behavior transitions as a function of the salinity for a (a) 30 wt% bitumen

and 70 wt% toluene oil phase mixture and (b) 20 mol% NAs and 80 mol% SDHS surfactant

mixture.

4.4.2 Interfacial Tension of Test Oil and Toluene Mixtures

The γow measurements of the test oil and toluene mixtures are presented as a function of

the salinity in Figures 4.3 (a)-(d).

84

(a) (b)

(c) (d)

NaCl (g/100 mL)

γo

w (

mN

/m)

Figure 4.3: The measured γow as a function of the salinity for (a) bitumen and toluene, (b)

asphalt and toluene, (c) naphthalene and toluene, and (d) deasphalted bitumen and toluene oil

phase mixtures.

According to this γow data, S* increases from ~3 g/100 mL to 3.5 g/100 mL as the

bitumen fraction in the oil phase is increased from 10 wt% to 50 wt%. However, for all the

fractions of asphalt and naphthalene dissolved in toluene, S* remains constant at 3 g/100 mL.

This is an indication that bitumen is more hydrophobic than both asphalt and naphthalene,

which have a similar polarity to that of toluene. For deasphalted bitumen, the increase in S*

reflects that the non-asphaltenic part of bitumen is more hydrophobic than whole bitumen and

its asphaltene-like components. This result is consistent with the fact that toluene is a suitable

solvent for asphaltenes whereas hexane and heptane are suitable solvents for maltenes. For

85

bitumen, asphalt, and deasphalted bitumen, the minimum γow increases with the test oil

composition. This observation is consistent with the fact that increasing the molar volume of the

solubilized oil reduces its solubilization capacity, likely due to the fact that it is more difficult to

disrupt the interaction amongst oil molecules21

. A reduction of the solubilization capacity is

directly linked to an increase in the minimum γow as expressed by the Chun Huh relationship

and the concept of interfacial rigidity19

.

4.4.3 EACN of Test Oils

To determine the average EACN of asphalt, bitumen, naphthalene, and deasphalted

bitumen, the EACNmix of each oil mixture is first calculated using the estimated S* from the γow

data presented in Figures 4.3 (a)-(d) as well as Equation 2. Knowing EACNmix and that the

EACN of toluene is 1, the EACN of each test oil can be calculated using Equation 3. However,

there is a question to be resolved before carrying out that calculation. That question is if the

composition xi in Equation 3 should be on the basis of moles, weight, or volume. While Baran

Jr. et al. assumed that the composition was based on the mole fraction, it is important to keep in

mind that the HLD equation is semi-empirical, and that there are therefore no fundamental bases

to argue one cause or the other. In fact, because the molecular weight of the different

compounds analyzed by Baran et al. is approximately the same, mole and weight fractions are

similar. The situation is different for mixtures of toluene with large asphaltenic crude oils,

making it imperative to determine if molar, weight, or volume fractions are more appropriate.

This issue is elucidated by determining, experimentally, the shifts in S* of hexadecane and

toluene mixtures and comparing these values to the salinity shifts calculated from the HLD

equation.

86

To obtain the equation for the salinity shift in the presence of electrolyte, the concept of

ideal mixing is applied to the HLD for a mixture of oils. Considering this mixture at optimal

conditions:

testtestrefrefmix HLDxHLDxHLD 0 (Eq. 5)

In this equation, the subscripts “ref” and “test” refer to the reference and test oils

respectively. Furthermore, considering Equation 1 at f(A)=0 and ΔT=0:

c

test

ocmixtestc

ref

ocmixref CKNSxCKNSx ,

*

,

* lnln0 (Eq. 6)

Knowing that *

refx + *

testx =1, one obtains:

c

test

octest

ref

ocrefmix CKNxKNxS ,,

*ln0 (Eq. 7)

For the reference oil alone, the following is true:

c

ref

ocref CKNS ,

*ln0 (Eq. 8)

Subtracting Equation 8 from Equation 7 yields the desired expression for the salinity

shift:

ref

oc

test

octest

ref

mix NNKxS

S,,*

*

ln

(Eq. 9)

To determine whether testx should be expressed on a molar, weight, or volume basis,

Equation 9 is applied to µEs produced in systems composed of hexadecane and toluene mixtures

as the oil phase and SDHS and sodium dioctyl sulfosuccinate (AOT) mixtures as the surfactant.

The results of this test are presented in Figure 4.4.

87

xhexadecane

Figure 4.4: Experimented and modeled shifts of S* for µEs composed of hexadecane and

toluene oil phase mixtures and 0.1 M of a 35 mol% SDHS and 65 mol% AOT surfactant

mixture.

The modeled shift in S* (solid line) for all fractions of hexadecane presented in Figure

4.4 are established using test

ocN , (hexadecane)=167. The data points presented as circles,

diamonds, and triangles in this figure are obtained by plotting the experimental salinity shift

versus the composition of hexadecane on a molar, weight, and volume basis respectively.

According to the data obtained, the volume fraction is the most accurate descriptor of the

composition terms in the EACN mixing rule. This finding is in agreement with that of Puerto et

al. who suggested that representing the oil phase composition on a molar basis reveals

nonlinearities in the phase behavior of surfactant-oil-water formulations22

. Until the

fundamental bases of the EACN mixing rule are elucidated, the significance of this finding may

not be fully appreciated. From a practical point of view, this finding provides a significant

advantage towards the elucidation of the EACN values of oils of unknown structure or complex

composition, such as bitumen and asphalt.

88

Equation 9 is subsequently applied (using testx as a volume fraction) to mixtures of the

test oils and toluene. Table 4.1 summarizes the parameters for these formulations and the

calculated EACN values for the oil mixtures and test oils.

Table 4.1: Calculated EACN of asphalt, bitumen, naphthalene, and deasphalted bitumen

Test Oil xtest ln( *

mixS ) EACNmix EACNtest

Asphalt

0.1 1.1 1.1

1.3 0.3 1.1 1.1

0.5 1.1 1.1

Bitumen

0.1 1.1 1.1

2.5 0.3 1.1 1.1

0.5 1.3 2.0

Naphthalene

0.1 1.1 1.1

1.3 0.2 1.1 1.1

0.3 1.1 1.1

Deasphalted

Bitumen

0.1 3.3 1.5

6.2 0.3 3.9 2.6

0.5 4.6 3.6

4.4.4 Interfacial Tension of Test Surfactant and SDHS Mixtures

The measured γow of µEs formulated for the test surfactant and SDHS mixtures

described in Section 4.3.4 are presented in Figures 4.5 (a)-(c) as a function of the salinity.

89

NaCl (g/100 mL)

(c) γ

ow (

mN

/m)

NaCl (g/100 mL) NaCl (g/100 mL)

γo

w (

mN

/m)

(a)

NaCl (g/100 mL)

(b)

γo

w (

mN

/m)

Figure 4.5: The measured γow as a function of the salinity for (a) NAs and SDHS, (b) NaNs and

SDHS, and (c) asphaltenes and SDHS surfactant mixtures at a total concentration of 0.1 M.

The data in Figure 4.5 suggests that NAs are more lipophilic than SDHS since an

increase in the surfactant mixture concentration of NAs produced a decrease in S*. On the other

hand, NaNs are more hydrophilic than SDHS since an increase in their concentration in the

surfactant mixture produced an increase in S*. By comparing Figures 4.5 (a) and (b), it is

evident that the displacement in S* for the case of NAs and SDHS mixtures is not substantial.

With regards to asphaltenes (Figure 4.5 (c)), a slight decrease in S* is observed as its

concentration within the surfactant mixture is increased from 10 mol% to 70 mol%. This

suggests that asphaltenes are lipophilic relative to both SDHS and NaNs. To compare the

hydrophilic-lipophilic nature of asphaltenes relative to NAs, however, the partitioning behavior

90

of asphaltenes at the oil-water interface must be accounted for. This data will be presented in

Section 4.4.5.

4.4.5 Cc of Test Surfactants

Acosta et al. previously developed Equation 10 below to determine the Cc of anionic

surfactants and fatty acids based on the shift of S* as a function of the surfactant composition

8:

test

c

ref

c

ref

mix CCS

S

test*

*

xln (Eq. 10)

The salinity shift of NAs and SDHS and NaNs and SDHS surfactant mixtures is

presented in Figures 4.6 (a) and (b) as a function of the test surfactant mole fraction within the

0.1 M surfactant mixture.

(a) (b)

xNA xNaN

Figure 4.6: Salinity shift of (a) NAs and SDHS and (b) NaNs and SDHS surfactant mixtures as

a function of the test surfactant mole fraction within the 0.1 M surfactant mixture.

91

The values of the salinity shifts in Figures 4.6 (a) and (b) are analyzed using the linear

regression tool within Microsoft Excel (2003). The linear regression analysis produced with this

tool yields the value of the slope and the 95% confidence interval for this value. These values

are presented in Table 4.2. The slope values can be introduced into Equation 10 to calculate the

Cc of NAs and NaNs. The results from these calculations are also presented in Table 4.2.

Table 4.2: Cc (* a

cC ) of NAs, NaNs, and asphaltene aggregates.

Test Surfactant Ln[ *

mixS / *

refS ]/X2 test

cC

NAs -0.31±0.05 -0.61±0.05

NaNs 1.7±0.1 -2.6±0.1

Asphaltene Aggregates* -2.0±0.2

* 1.1±0.2

*

Evaluating the a

cC of asphaltene aggregates requires an understanding of the partitioning

of asphaltenes between the bulk oil phase and the oil-water interface. Figure 4.7 (a) presents the

molar fraction of asphaltenes (in mixtures with SDHS) at the interface versus the equilibrium

volume fraction of asphaltenes (in mixture with toluene) in the excess bulk phase. According to

this figure, only when the volume fraction of asphaltenes in toluene is larger than 0.03 do the

asphaltenes adsorb significantly at the oil-water interface. A plateau fraction of asphaltenes at

the oil-water interface is reached at ~0.12. At asphaltene volume fractions in the oil phase larger

than 0.05, these asphaltenes tend to precipitate out of solution. While asphaltenes in toluene

alone are completely soluble, the observed precipitation is consistent with earlier observations

by Yarranton et al. that the presence of oil-water interfaces induces the precipitation of

asphaltenes23

.

92

(a) (b)

Figure 4.7: (a) Asphaltene partitioning and (b) the shift in S* for optimum µEs formulated with

an asphaltenes and SDHS surfactant mixture.

In previous work involving the characterization of the Cc of polar organic molecules,

like lipophilic linkers, partitioning has been identified as a source of error8. In order to account

for these effects, the shift in S* is plotted in Figure 4.7 (b) as a function of the calculated molar

fraction of asphaltenes in mixtures with SDHS at the interface. Using the same mathematical

approach as used for the characterization of the Cc of NAs and NaNs, the a

cC of asphaltene

aggregates may be determined. Unlike the case for NAs and NaNs, a single regression line

cannot be used to fit the data presented here. To overcome this issue, the data is separated into a

non surface-active asphaltenes cluster and a surface-active asphaltenes cluster at the transition

point indicated in the above discussion (~0.1). In doing so, separate regression lines can be fit to

each of the clusters. Considering that the concept of Cc applies to surface-active species, the

slope of the regression line for surface-active asphaltenes and Equation 10 were used to

calculate the a

cC of surface-active asphaltene aggregates as presented in Table 4.2.

93

4.5 DISCUSSIONS

4.5.1 Analysis of the Hydrophilic-Lipophilic Nature of Asphaltenic Crude Oils

The calculated EACN of the asphaltenic oils listed in Table 4.1 confirm that, indeed,

these oils have a similar hydrophilic-lipophilic nature as that of toluene. It is counter intuitive to

think that such heavy crude oil fractions are less hydrophobic than light crude oils, which are

typically simulated (in µE phase behavior studies) with oils such as octane (EACN=8)24

. Such

counterintuitive behavior has been reported before, but this is the first time that it has been

quantified through the EACN value25

.

Another interesting observation is that the EACN of bitumen is slightly larger than that

of asphalt. This could be explained by the fact that bitumen contains less asphaltenes (~15%)

than asphalt (~45%)26

. The impact of asphaltenes on the polarity of these crude oils may be

elucidated from the EACN of the deasphalted (maltene-like) fraction of bitumen. The EACN of

this deasphalted bitumen fraction is between that of hexane and heptane. Considering that

asphaltenes only represent ~15% of bitumen but reduces the EACN of bitumen to a value of

approximately 2 suggests that the polar nature of asphaltenes dominates the hydrophilic-

lipophilic nature of bitumen. Furthermore, the relatively low EACN of asphaltenic oils also

suggests that they should have amphiphilic properties. For example, it has been shown that

benzene is a polar molecule (EACN=0) that is also surface-active and tends to segregate near

the oil-water interface27

.

The simple method of mixing a test oil and reference oil to determine the EACN of the

test oil is successful in principle. The fact that volume fractions, instead of mole fractions, are

more in line with the HLD model allows the use of this method for oils of unknown structure

94

and composition. This simple test and reference oil phase mixture method is simpler than other

available methods to determine the EACN of a wide range of oils7.

4.5.2 Analysis of the Hydrophilic-Lipophilic Nature of Naphthenic Amphiphiles

and Asphaltene Aggregates

To put the Cc of NAs and NaNs presented in Table 4.2 into context, it is important to

recall that Acosta et al. determined that the Cc value of NaNs obtained from Eastman Kodak is

-2.4±0.28. While the NaNs source studied here is different (obtained by neutralizing NAs

purchased from Sigma-Aldrich Canada, Inc.), the obtained Cc is similar. In the case of NAs, the

calculated Cc is in between that of oleic acid (Cc=0) and SDHS (Cc=-0.92). In other words, NAs

are more hydrophilic than oleic acid. Another interesting point of comparison is that sodium

oleate, the salt of oleic acid, has a Cc of -1.78. In this case, the neutralization of the fatty acid

produced a decrease in Cc of 1.7. In the case of NAs, the neutralization to NaNs produced a

similar decrease in Cc of 2.

With regards to the a

cC of asphaltene aggregates, it should be re-emphasized that the

molecular weight of asphaltenes is a probability distribution with a peak intensity at 600 g/mol.

Assuming that the molecular weight distribution of asphaltenes varies from 500 g/mol to 1000

g/mol according to the analysis performed by Pomerantz et al., the a

cC limits of the asphaltene

aggregates may range from 0.8 to 2.320

. In comparison to both NAs and NaNs, it may be

concluded that asphaltene aggregates are much more lipophilic. In general terms, the upper a

cC

limit of asphaltene aggregates is comparable to that of AOT (Cc=2.55), which is one of the most

lipophilic surfactants identified thus far8.

95

4.6 CONCLUSIONS

Using the HLD model for ionic surfactants as well as several basic mixing rules, it was

shown that bitumen had an equivalent alkane carbon number (EACN) of 2.5, a value slightly

larger than that of asphalt (EACN=1.3) and naphthalene (EACN=1.3). These low EACN values

reflect the polar nature of polyunsaturated aromatic compounds present in these oils. The polar

nature of asphaltenes was additionally justified by the fact that the EACN of deasphalted

bitumen was ~6.2. It was furthermore demonstrated that asphaltenes were surface-active under

certain conditions and that they were significantly more lipophilic than both naphthenic acids

(NAs, Cc=-0.61) and sodium naphthenates (NaNs, Cc=-2.6).

4.7 REFERENCES

1 Salager, J.-L.; Antón, R.E.; Sabatini, D.A.; Harwell, J.H.; Acosta, E.J.; Tolosa, L.I. Enhancing

Solubilization in Microemulsions-State of the Art and Current Trends. J. Surfactants Deterg.

2005, 8, 3-21.

2 Acosta, E.J. The HLD-NAC Equation of State for Microemulsions Formulated with Nonionic

Alcohol Ethoxylate and Alkylphenol Ethoxylate Surfactants. Colloids Surf., A 2008, 320, 193-

204.

3 Healy, R.N.; Reed, R.L. Physicochemical Aspects of Microemulsion Flooding. Soc. Pet. Eng.

AIME J. 1974, 14, 491-501.

96

4 Healy, R.N.; Reed, R.L.; Stenmark, D.G. Multiphase Microemulsion Systems. Soc. Pet. Eng.

AIME J. 1976, 16, 147-160.

5 Antón, R.E.; Salager, J.-L. Effect of the Electrolyte Anion on the Salinity Contribution to

Optimum Formulation of Anionic Surfactant Microemulsions. J. Colloid Interface Sci. 1990,

140, 75-81.

6 Salager, J.-L.; Morgan, R.; Schechter, R.S.; Wade, W.H.; Vasquez, E. Optimum Formulation

of Surfactant/Water/Oil Systems for Minimum Interfacial Tension or Phase Behavior. Soc. Pet.

Eng. AIME J. 1979, 19, 107-115.

7 Nardello, V.; Chailloux, N.; Poprawski, J.; Salager, J.-L.; Aubry, J.-M. HLD Concept as a

Tool for the Characterization of Cosmetic Hydrocarbon Oils. Polym. Int. 2003, 52, 602-609.

8 Acosta, E.J.; Yuan, J.S.; Bhakta, A.S. The Characteristic Curvature of Ionic Surfactants. J.

Surfactants Deterg. 2008, 11, 145-158.

9 Van Hecke, E.; Catté, M.; Poprawski, J.; Aubry, J.-M.; Salager, J.-L. A Novel Criterion for

Studying the Phase Equilibria of Nonionic Surfactant-Triglyceride Oil-Water Systems. Polym.

Int. 2003, 52, 559-562.

10 Tongcumpou, C.; Acosta, E.J.; Scamehorn, J.F.; Sabatini, D.A.; Yanumet, N.; Chavadej, S.

Enhanced Triolein Removal using Microemulsions Formulated with Mixed Surfactants. J.

Surfactants Deterg. 2006, 9, 181-189.

97

11 Thakur, R.K.; Villette, C.; Aubry, J.-M.; Delaplace, G. Formulation-Composition Map of a

Lecithin-Based Emulsion. Colloids Surf., A 2007, 310, 55-61.

12 Rondón, M.; Bouriat, P.; Lachaise, J.; Salager, J.-L. Breaking of Water-in-Crude Oil

Emulsions. 1. Physicochemical Phenomenology of Demulsifier Action. Energy Fuels 2006, 20,

1600-1604.

13 Baran Jr., J.R.; Pope, G.A.; Wade, W.H.; Weerasooriya, V.; Yapa, A. Microemulsion

Formation with Mixed Chlorinated Hydrocarbon Liquids. J. Colloid Interface Sci. 1994, 168,

67-72.

14 Smith, D.F.; Schaub, T.M.; Kim, S.; Rodgers, R.P.; Rahimi, P.; Teclemariam, A.; Marshall,

A.G. Characterization of Acidic Species in Athabasca Bitumen and Bitumen Heavy Vacuum

Gas Oil by Negative-Ion ESI FT-ICR MS with and without Acid-Ion Exchange Resin

Prefractionation. Energy Fuels 2008, 22, 2372-2378.

15 Frank, R.A.; Kavanagh, R.; Kent Burnison, B.; Arsenault, G.; Headley, J.V.; Peru, K.M.; Van

Der Kraak, G.; Solomon, K.R. Toxicity Assessment of Collected Fractions from an Extracted

Naphthenic Acid Mixture. Chemosphere 2008, 72, 1309-1314.

16 Headley, J.V.; Peru, K.M.; Armstrong, S.A.; Han, X.; Martin, J.W.; Mapolelo, M.M.; Smith,

D.F.; Rogers, R.P.; Marshall, A.G. Aquatic Plant-Derived Changes in Oil Sands Naphthenic

Acid Signatures Determined by Low-, High-, and Ultrahigh-Resolution Mass Spectrometry.

Rapid Commun. Mass Spectrom. 2009, 23, 515-522.

98

17 Akbarzadeh, K.; Alboudwarej, H.; Svrcek, W.Y.; Yarranton, H.W. A Generalized Regular

Solution Model for Asphaltene Precipitation from n-Alkane Diluted Heavy Oils and Bitumens.

Fluid Phase Equilib. 2005, 232, 159-170.

18 Heric, E.L.; Posey, C.D. Interaction in Nonelectrolyte Solutions. II. Solubility of Naphthalene

at 25°C in some Mixed Solvents Containing Toluene, Ethylbenzene. J. Chem. Eng. Data 1964,

9, 161-165.

19 Acosta, E.; Szekeres, E.; Sabatini, D.A.; Harwell, J.H. Net-Average Curvature Model for

Solubilization and Supersolubilization in Surfactant Microemulsions. Langmuir 2003, 19, 186-

195.

20 Pomerantz, A.E.; Hammond, M.R.; Morrow, A.L.; Mullins, O.C.; Zare, R.N. Two-Step Laser

Mass Spectrometry of Asphaltenes. J. Am. Chem. Soc. 2008, 130, 7216-7217.

21 Acosta, E.J.; Nguyen, T.; Witthayapanyanon, A.; Harwell, J.H.; Sabatini, D.A. Linker-Based

Bio-Compatible Microemulsions. Environ. Sci. Technol. 2005, 39, 1275-1282.

22 Puerto, M.C.; Reed, R.L. Three-Parameter Representation of Surfactant/Oil/Brine Interaction.

Soc. Pet. Eng. J. 1983, 23, 669-682.

23 Yarranton, H.W.; Hussein, H.; Masliyah, J.H. Water-in-Hydrocarbon Emulsions Stabilized by

Asphaltenes at Low Concentrations. J. Colloid Interface Sci. 2000, 228, 52-63.

99

24 Bourrel, M. and Schechter, R., 1988. Microemulsions and Related Systems: Formulation,

Solvency, and Physical Properties. New York: Marcel Dekker, Inc.

25 Acevedo, S.; Borges, B.; Quintero, F.; Piscitelly, V.; Gutierrez, L.B. Asphaltenes and Other

Natural Surfactants from Cerro Negro Crude Oil. Stepwise Adsorption at the Water/Toluene

Interface: Film Formation and Hydrophobic Effects. Energy Fuels 2005, 19, 1948-1953.

26 Shakirullah, M.; Ahmad, I.; Arsala Khan, M.; Ali Shah, A.; Ishaq, M.; Habib-ur-Rehman.

Conversion of Asphalt into Distillate Products. Energy Convers. Manage. 2008, 49, 107-112.

27 Szekeres, E.; Acosta, E.; Sabatini, D.A.; Harwell, J.H. A Two-State Model for Selective

Solubilization of Benzene-Limonene Mixtures in Sodium Dihexyl Sulfosuccinate

Microemulsions. Langmuir 2004, 20, 6560-6569.

100

CHAPTER 5:

EXPERIMENTAL EVALUATION OF EMULSION STABILITY

VIA SURFACTANT-OIL-WATER PHASE BEHAVIOR SCANS

This chapter is derived from the following submitted manuscript:

Kiran, S.K.; Acosta, E.J. Experimental Evaluation of Emulsion Stability Via Surfactant-Oil-

Water Phase Behavior Scans. J. Surfactants Deterg. (ID: JSD-13-0018).

101

5.1 ABSTRACT

The stability of anionic surfactant (sodium dihexyl sulfosuccinate (SDHS))-oil (toluene)-

water emulsions was characterized in this chapter in terms of aggregation (ap), drainage (dp), and

coalescence (cp) time periods. These time periods, determined by tracking the normalized

displacement of a set of phase separation fronts, were evaluated as a function of the quantified

proximity to the phase inversion point using the hydrophilic-lipophilic deviation (HLD)

framework. The secondary effects of surfactant concentration and temperature on emulsion

stability were also evaluated. It was determined that at a low surfactant concentration (0.01 M

SDHS), the cp time period was most dominant for oil droplets in water-continuous formulations

(HLD<0) and water droplets in oil-continuous formulations (HLD>0). At larger surfactant

concentrations, the ap time period was most dominant at HLD<0 whereas all of the time periods

were comparable at HLD>0. In addition, the apparent drainage and coalescence activation

energies were shown to be strongly influenced by the interfacial tension, whose minimum

coincides with that of the overall emulsion stability at the phase inversion point (HLD=0).

5.2 INTRODUCTION

It was previously described in Chapters 1 and 4 that microemulsion (µE) phase behavior

scans can be used to evaluate formulation conditions that produce an inversion in the surfactant

partitioning between the aqueous and oil phases. Under net formulation conditions that favor an

overall hydrophilic (hydrophobic) environment, the surfactant preferentially partitions into the

aqueous (oil) phase as oil (water)-swollen micelles (or Type I (Type II) µEs) in equilibrium with

an excess oil (aqueous) phase. At balanced formulation conditions, where phase inversion takes

place, oil and water co-solubilize into a bicontinuous (or Type III) µE phase that co-exists with

102

excess oil and aqueous phases. A Type IType IIIType II phase behavior scan for sodium

dihexyl sulfosuccinate (SDHS)-toluene-water µEs is illustrated in Figure 5.1. The ability to

evaluate various equilibrium properties including solubilization capacities, interfacial tensions

(γow), densities (ρµE), and, as will be described in the next chapter, viscosities (ηµE) from such a

simple phase behavior scan is useful in applications ranging from surfactant enhanced oil

recovery to the design of cosmetic and pharmaceutical drug delivery systems1,2

.

NaCl

(g/100 mL): 10752.41.50.8 432.3

0HLD: -1.4 -0.7 -0.3 -0.2 0.3 0.5 0.9 1.2

Type I Type III Type II

Figure 5.1: A Type IType IIIType II phase behavior scan for 0.1 M SDHS-toluene-water

µEs at 25°C.

Phase behavior scans also produce information that is relevant to the dynamics of

emulsification3,4,5,6,7,8,9,10,11

. Under the same formulation conditions as for Type I (Type II) µEs,

emulsions take the form of the excess oil (aqueous) phase dispersed throughout the oil-in-water

(o/w) (water-in-oil (w/o)) µE phase itself. Under Type III conditions, both the excess oil and

aqueous phases are dispersed throughout the bicontinuous µE phase. The corresponding

connection to the stability of these emulsions is less well understood. The aim of this work is to

103

provide a new insight into this connection by evaluating the kinetics of the different separation

processes that take place as a function of the surfactant concentration and temperature.

Salager et al. and Binks et al. proposed different separation processes that influence the

kinetics of demulsification around the phase inversion point7,9,10

. By identifying and tracking the

normalized displacement of the µE and excess phase separation fronts versus time, these authors

argued that the aggregation of emulsion droplets into flocs leads to a faster rate of settling. The

subsequent onset of film thinning into the resulting Plateau borders of a highly concentrated

emulsion layer further accelerates the onset of coalescence. A schematic overview of each of

these processes is illustrated in Figure 5.2.

Aggregation

Film Thinning

Coalescence

Settling

Figure 5.2: Emulsion aggregation, settling, film thinning, and coalescence processes.

104

While the studies of Salager et al. and Binks et al. showed that the µE and excess phases

separate the fastest at the phase inversion point, they failed to provide a detailed description of

its kinetics. The first objective of this chapter is to assess how aggregation, settling, film

thinning, and coalescence contribute to the total separation time. This will be conducted as a

function of the proximity to the phase inversion point. Figure 5.3 (a) illustrates how the

observable phase separation stages evolve over time in a test tube for emulsified water droplets

(black) dispersed throughout a continuous Type II µE phase (gray). Figure 5.3 (b) presents the

normalized displacement of the phase separation fronts versus time and the extrapolated time

periods.

μE Phase Aqueous Phase

Δhw(t)=Δhw,tot

ΔhμE(t)=ΔhμE,totΔhμE(t)ΔhμE(t)

Δhw(t)Δhw(t)

ΔhμE(t)

Time

ap dp

No

rmalized

Dis

pla

cem

en

t

Timecp

(a)

(b)

Figure 5.3: (a) Evolution of the observable phase separation stages versus time in a test tube for

emulsions produced in Type II µEs. (b) Normalized displacement of the µE and excess aqueous

phase separation fronts versus time and the extrapolated ap, dp, and cp time periods.

105

No observable changes to the emulsion can be visually detected over the initial induction

(or aggregation) time period (ap) until a µE phase separation front develops. The period of time

between the onset of the µE phase separation front and the appearance of an aqueous phase

separation front is referred to as the drainage time period (dp). Over this time period, the

drainage of the continuous µE phase essentially leads to the production of a more concentrated

emulsion layer. The final period of time between the onset of the aqueous phase separation front

and its complete separation is referred to as the coalescence time period (cp). The onset of the

final cp time period is realized once a critical packing is reached and represents the rejection of

the excess phase from the emulsion. It is important to clarify that the intent of selecting the

names for these phase separation time periods as such is to reflect the rate-limiting process that

is expected to dominate demulsification. In reality, however, a multitude of these processes may

be acting together at any point in time.

The second objective of this chapter is to develop an understanding of how the µE phase

behavior properties influence the duration of the emulsion phase separation time periods. The

initial ap time period likely depends on the original emulsion droplet size. The original emulsion

droplet size is in turn a function of the mixing energy input and γow12

. Lower γow should

theoretically lead to the formation of smaller emulsion droplets that settle slowly but that are

much more mobile as a result of Brownian motion13,14

. This increase in mobility should lead to

an increase in the collision frequency of emulsion droplets and a therefore greater probability of

their aggregation. It is very possible that the emulsion droplets that make up this aggregated

state undergo an internal coalescence process that does not result in any separation of the excess

phase. The next dp time period is more complex in nature as it likely involves the settling of

aggregated flocs of growing emulsion droplets as well as film thinning. The final cp time period

106

is likely associated with coalescence processes resulting from the breakage of oil-water-oil

and/or water-oil-water interfaces. The rupture of these interfaces may be explained by the hole

nucleation theory. This theory predicts that the activation energy of coalescence is directly

proportional to γow15,16,17

.

Another objective of this chapter is to explore the influence of the surfactant

concentration on the emulsion phase separation time periods. According to the literature, an

increase in the surfactant concentration may lead to several possible outcomes. For one, the

initial ap time period may be lengthened due to an increase in the electrostatic and steric

repulsions caused by the additional µE droplets in the continuous phase medium that separates

the approaching emulsion droplets18

. On the other hand, it is also possible for the ap time period

to decrease via attractive depletion interactions at low volume fractions of µE droplets17,18,19,20

.

The ap time period should also decrease in the event of µE-mediated Ostwald ripening10,21,22

.

With regards to the subsequent dp time period, an increase in its timescale can be expected

because of the hindered settling of emulsion droplets throughout a continuous µE phase of

increased ηµE6,10

. This effect may become even more pronounced if the number of metastable

µE layers that need to be expelled during film thinning increases23,24

. For the final cp time

period, the added amount of adsorbed surfactant should lead to an increase in its timescale due

to a better ability to withstand any sort of film rupturing11

. Additionally, depletion-repulsion

mechanics (i.e. the osmotic pressure cost of excluding µE droplets from the inter-droplet

continuous phase medium) may even further prolong the cp time period25,26

.

The last objective of this work is to study the effect of temperature (T) on the emulsion

phase separation time periods. By treating aggregation, drainage, and coalescence as a series of

chemical reactions that conform to Arrhenius’s law, an increase in T should foreseeably lead to

107

a decrease in their time periods across the µE phase behavior spectrum. The apparent activation

energy (W*) for a given one of aggregation, drainage, and coalescence can be calculated using

the following expression of Kabalnov et al.16,27

:

Tk

Wft

B

p

*exp (Eq. 1)

Here, tp is the emulsion phase separation time period of interest, f is a proportionality

constant that is assumed to be independent of T, and kB is the Boltzmann constant. Kabalnov et

al. applied Equation 1 to determine W* for only the coalescence of Type III emulsions. Other

efforts by Acosta et al. and Witthayapanyanon et al., which are equally as constrained as those

of Kabalnov et al., looked to instead fit W* according to changes in the emulsion coalescence

rate as a function of the interfacial rigidity of the system28,29

.

In this chapter, the ap, dp, and cp time periods will be studied for SDHS-toluene-water

emulsions. As will be shown in greater detail later, the total stability of these emulsions falls in

the range of just seconds to a few hours. These are much more reasonable timeframes for

performing high throughput analysis compared to days as has been published for sodium dioctyl

sulfosuccinate (AOT) stabilized emulsions9,10

. The proximity to the phase inversion point will

be quantified for SDHS-toluene-water emulsions using the following hydrophilic-lipophilic

deviation (HLD) expression for ionic surfactants that was described in detail in Chapter 4:

coc CTAfNSHLD 01.017.0ln , (Eq. 2)

In short, S is the electrolyte concentration (in g/100 mL), Nc,o is the equivalent alkane

carbon number (EACN) of the oil phase (EACN=1 for toluene), f(A) is a function of the co-

surfactant type and concentration, ΔT is the difference in temperature from 25°C, and Cc is the

108

characteristic curvature of the surfactant (Cc=-0.92 for SDHS). A HLD<0HLD=0HLD>0

scan reflects a Type IType IIIType II µE phase behavior transition.

5.3 MATERIALS AND METHODS

5.3.1 Materials

NaCl (product #S9625, ≥99.5%) and SDHS (product #86146, ~80% in water) were

purchased from Sigma-Aldrich Corp. (Oakville, ON, Canada). Reagent grade acetone (product

#1200-1-10) and toluene (product #9200-1-10) were purchased from Caledon Laboratory

Chemicals (Georgetown, ON, Canada). All of these materials were stored at 25°C and used as

received. Tap water was deionized using an APS Ultra mixed bed resin to a conductivity <3

µs/cm.

A multipoint turbidimeter was constructed in-house using laser modules (product

#VLM-650-03-LPA), linear photodiodes (product #OPT101P), and digital potentiometers

(product #AD5206BN10) acquired from Digi-Key (Thief River Falls, MN, USA). Almost all

other electrical components (solderless breadboards, insulated wires, hex inverters (product

#74LS04), and an Arduino Diecimilia microcontroller) were obtained from Creatron Inc.

(Toronto, ON, Canada). To expand the number of input/output ports on the Arduino Diecimilia

microcontroller, a third party multiplexer (mux) shield was purchased from Mayhew

Laboratories (Greenville, SC, USA).

5.3.2 Microemulsion Phase Behavior Scans and Emulsification

The baseline µE phase behavior scan, shown in Figure 5.1, was conducted in 15 mL flat-

bottom glass test tubes at 25°C by adding 5.5 mL of toluene to 5.5 mL of an aqueous phase

109

solution composed of 0.8-10 g/100 mL NaCl and 0.1 M SDHS. After having mixed these phases

together for 1 minute at 3200 rpm using a 150 W VWR analog vortex mixer, they were allowed

to equilibrate. The resulting sphere-equivalent µE droplet sizes, as calculated using the net-

average curvature (NAC) model in the next chapters, were on the order of ~1-30 nm.

Emulsification was induced by re-agitating these pre-equilibrated µEs. It is safe to assume from

the similar viscosities of toluene and water that emulsion droplet breakup occurred

instantaneously under the applied mixing conditions12

. It can further be concluded from the

interplay of hydrodynamic and thermodynamic effects that the resulting emulsion droplets

immediately took on the expected morphology as predicted by the HLD model30

.

Additional µE phase behavior scans were conducted at 0.01 M and 0.3 M SDHS. To

account for SDHS’s contribution to the total electrolyte concentration (S in Equation 2), 15% of

its sodium counterion was assumed to dissociate. This assumption is consistent with previous

observations for SDHS phase behavior scans31,32

. Only select baseline (0.1 M SDHS) µEs at

HLD=0, ±0.4, ±0.7, and ±1.1 were formulated at T=7°C, 15°C, 35°C, and 44°C. At 7°C and

15°C, all emulsification and phase separation experiments were carried out in a refrigerated

incubator (Fisher Scientific International Inc.). At 35°C and 44°C, all emulsification and phase

separation experiments were carried out in an oven (Cenco Instruments Corp.).

5.3.3 Interfacial Tension of Baseline Microemulsions

The γow between the heavy and light phases of equilibrium baseline (0.1 M SDHS) µEs

at 25°C was measured using the spinning drop tensiometer technique described in Chapters 2.

110

5.3.4 Average Diameter of Baseline Emulsion Droplets

Optical micrographs of a 10 µL emulsified baseline (0.1 M SDHS) sample at 25°C were

taken using an Olympus wide zoom digital camera at 4× magnification mounted on top of an

Olympus BX-51 microscope set at 100× magnification within 1 minute after the test tube was

shaken. From the number frequency (Fi) of observed emulsion droplets of measured diameter

E

id in a population size of 30, the average emulsion droplet diameter ( E

dd ) was calculated as per

Chapter 2:

2

3

E

ii

E

iiE

d

dF

dFd (Eq. 3)

5.3.5 Emulsion Phase Separation Profiles

Emulsion phase separation profiles were assessed by evaluating the normalized

displacement of the µE (ΔsµE(t)), excess oil (Δso(t)), and/or excess aqueous (Δsw(t)) phase

separation fronts versus time as follows:

totE

E

Eh

thts

,

(Eq. 4)

toto

o

oh

thts

,

(Eq. 5)

totw

ww

h

thts

,

(Eq. 6)

In these equations, ΔhµE(t), Δho(t), and Δhw(t) are the tracked net displacement of the

moving µE, excess oil, and excess aqueous phase separation fronts whereas ΔhµE,tot, Δho,tot, and

111

Δhw,tot are their respective total required displacement for equilibrium to be reached. The net and

total required displacements of the relevant phase separation fronts for a Type II emulsion are

described in Figure 5.3 (a). The primary method of tracking the net displacement of moving

phase separation fronts was by taking time-lapse images every 15 seconds. The gradient in the

gray level for each series of images pertaining to a given emulsion was analyzed using

MATLAB’s image processing toolbox in order to help identify the location of the moving phase

separation fronts.

The net displacement of moving phase separation fronts was additionally tracked using

high frequency turbidity (τ(t)) measurements with the array of laser-photodiode pairs (switched

on/off for 0.5 seconds/cycle in a preset order) shown in Figure 5.4.

0.8 cm

2.2 cm

3.7 cm

5.1 cm

6.5 cm

1.5 cm

3 cm

4.5 cm

6 cm

Figure 5.4: In-house multipoint turbidimeter used to track the net displacement of µE and

excess phase separation fronts at a high time resolution.

The τ(t) was calculated using this multipoint turbidimeter at the different heights as

follows28,29

:

112

tV

V

Lt

p

maxln1

(Eq. 7)

In this equation, Lp is the laser path length, V(t) is the voltage of the photodiode obtained

at a given time during the measurement, and Vmax is the maximum attainable voltage of the

photodiode in a transparent sample. In this case, the passage of a moving phase separation front

through a given set of laser-photodiode pairs was identified by a sharp decay in τ(t) to 0. The

data taken from the time-lapse images and multipoint τ(t) measurements were both used (for

validation purposes) to plot the normalized displacement of the moving µE and excess phase

separation fronts versus time as in Figure 5.3 (b). These graphs were then used to determine the

resulting ap, dp, and cp time periods.

5.4 RESULTS

5.4.1 Interfacial Tension and Average Emulsion Droplet Diameter

The baseline (0.1 M SDHS) γow profile at 25°C is presented in Figure 5.5 (a) as a

function of the HLD. The measured data, which follows the same trend as that observed for the

other µEs in Chapter 4, shows that an ultralow minimum is attained at HLD=0.

113

0.01

0.1

1

-1.5 -1 -0.5 0 0.5 1 1.5

γo

w(m

N/m

)

(a)

Type I Type III Type II

(μm)

HLD

(b)

0.1

1

10

-1.5 -1 -0.5 0 0.5 1 1.5

Type I Type III Type II

Figure 5.5: Measured (a) interfacial tension (γow) and (b) emulsion droplet diameter ( E

dd )

profiles for baseline (0.1 M SDHS) emulsions at 25°C.

The baseline (0.1 M SDHS) E

dd profile at 25°C in Figure 5.5 (b) is also presented as a

function of the HLD. It can be extrapolated from the data that E

dd tends to a minimum at

HLD=0. This trend is directly related to the changes in γow as a function of the HLD. The

relationship between E

dd and γow can be expressed based on the balance of hydrodynamic and

interfacial forces as discussed in Chapter 3 as follows:

ow

E

dd (Eq. 8)

It should be noted here that the reported E

dd are of the same order of magnitude as the

calculated number average droplet diameter. Furthermore, E

dd are not reported for emulsions in

the vicinity of the Type III region as their E

id could not accurately be measured.

dd

E (

µm

)

114

5.4.2 Emulsion Phase Separation Profiles and Time Periods

In Figure 5.6, ΔsµE(t) and Δsw(t) profiles versus time for a Type II baseline (0.1 M

SDHS) emulsion at HLD=0.9 and 25°C are presented along with sample time-lapse images that

show the real progression of the phase separation fronts.

0

0.2

0.4

0.6

0.8

1

0 1750 3500 5250 7000

y1 = 0.0008x1 - 1.46R² = 0.95

0

0.2

0.4

0.6

0.8

1

0 1750 3500 5250 7000

y2 = 0.0008x2 - 2.13R² = 0.96

0

0.2

0.4

0.6

0.8

1

0 1750 3500 5250 7000

0

0.2

0.4

0.6

0.8

1

ep

0

0.2

0.4

0.6

0.

8

1

0 1750 3500 5250 7000

cpdpap

Time (Seconds)Time

Images - ΔsμE(t)

Images - Δsw(t)

τ(t) - ΔsμE(t)

τ(t) - Δsw(t)

Linear Approximation - ΔsμE(t)

Linear Approximation - Δsw(t)

ΔsμE(t) (or Δs

w(t

))

Figure 5.6: Example of sample time-lapse images and ΔsµE(t) and Δsw(t) profiles versus time

for a Type II baseline (0.1 M SDHS) emulsion at HLD=0.9 and 25°C.

The ap time period is represented by the initial lag time before which the µE phase

separation front begins to advance. To calculate the duration of the ap time period, the advance

of the µE phase separation front is fitted to a linear equation that assumes a constant advance

rate. The intercept of this linear equation at ΔsµE(t)=0 corresponds to the ap time period. A

similar method is used to calculate the dp time period. What is different in this case, however, is

that the ap time period is subtracted from the intercept of this second linear equation at Δsw(t)=0.

The final cp time period is calculated by subtracting the ap and dp time periods from the time at

115

which the second linear equation intercepts Δsw(t)=1. An interesting feature of these phase

separation profiles, which is clearly illustrated in the accompanying time-lapse images, is that

the advance of the µE phase separation front does not leave behind a “clean” µE phase. Instead,

a small volume fraction of isolated emulsion droplets settles very slowly before eventually

coalescing. The impact of this phenomenon on the total time required to reach equilibrium (ep)

is ignored. The exact same approach applies to Type I emulsions (but for Δso(t) instead of

Δsw(t)). For Type III emulsions, where oil and water droplets co-exist, the dp time period is

allowed to span the time interval between the onset of ΔsµE(t) and the more stable emulsion

type. The cp time period is thereafter allowed to proceed until equilibrium is reached. At the

phase inversion point (i.e. HLD=0), no ap and dp time periods are observed as the coalescence of

oil and water emulsion droplets occurs instantaneously. The maximum calculated error between

the times estimated using the time-lapse images and multipoint turbidimeter is, on average,

15%.

Figure 5.7 presents the net rate of advance of the µE and excess phase separation fronts

as a function of the HLD for baseline (0.1 M SDHS) emulsions at 25°C. Overall, the net rate of

advance of both fronts tend to be similar and increase upon approaching the phase inversion

point at HLD=0. It should be noted that the net rate of advance of the µE and excess phase

separation fronts were not originally expected to be so similar. The fact that they are however so

similar may suggest some strong commonalities with respect to their rate-controlling separation

processes. Near the phase inversion point, the µE phase separation front advances faster than the

excess phase separation front. The net rate of advance of the phase separation fronts for Type III

emulsions are not included because these emulsions separate almost instantaneously.

116

0

0.1

0.2

0.3

-1.5 -1 -0.5 0 0.5 1 1.5

HLD

Net

Rate

of

Ad

van

ce (

mm

/s)

μE Phase Separation Front Excess Phase Separation Front

Type I Type III Type II

Figure 5.7: Net rate of advance of the µE and excess phase separation fronts as a function of the

HLD for baseline (0.1 M SDHS) emulsions at 25°C.

In Figures 5.8 (a)-(c), a compilation of the ap, dp, cp, and ep (ap+dp+cp) time periods is

presented as a function of the HLD at 0.01 M, 0.1 M, and 0.3 M SDHS and 25°C. In general, all

of these time periods tend to decrease upon approaching HLD=0. For emulsions at 0.01 M

SDHS, the overall stability is dominated by the cp time period whereas the ap time period

represents the smallest contribution. For emulsions at 0.1 M and 0.3 M SDHS, the ep time

periods, which are substantially larger than those obtained at 0.01 M SDHS, appear close.

However, the ap time period for Type I emulsions at 0.1 M SDHS, which is the major

contributor to their overall stability, is noticeably larger than those obtained at 0.3 M SDHS.

117

epcpdpap

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Tim

e (

Seco

nd

s)

Type I Type III Type II(a)

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Tim

e (

Seco

nd

s)

(b)

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Tim

e (

Seco

nd

s)

(c)

HLD

Figure 5.8: Compiled ap, dp, cp, and ep (ap+dp+cp) time periods for emulsions at (a) 0.01 M, (b)

0.1 M, and (c) 0.3 M SDHS and 25°C.

5.4.3 Effect of Temperature on Emulsion Phase Separation Time Periods

The effect of T on the ap, dp, cp, and ep time periods is illustrated in Figures 5.9 (a)-(d) as

a function of the HLD for baseline (0.1 M SDHS) emulsions. All of these phase separation time

118

periods tend to decrease with an increase in T, but this decrease is more noticeable for

emulsions that are far from HLD=0. One point that deserves closer inspection is to compare the

effect of T on the ap time periods at HLD=-0.4 and HLD=0.4. At HLD=-0.4, the increase in T

produces a substantial decrease in the ap time period. At HLD=0.4, however, the ap time period

does not change very much with T and remains quite low. This finding suggests that, at least for

the ap time period, the morphology of the emulsion plays a role (particularly close to HLD=0).

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

(a) (b)

(c) (d)

Tim

e (

Se

co

nd

s)

Tim

e (

Se

co

nd

s)

Type I Type III Type II

epcpdpap

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

HLD HLD

Type I Type III Type II

Figure 5.9: Compiled ap, dp, cp, and ep (ap+dp+cp) time periods for baseline (0.1 M SDHS)

emulsions at (a) T=7°C, (b) T=15°C, (c) T=35°C, and (d) T=44°C.

119

In Figure 5.10 (a), an example of the Arrhenius-like relationship proposed by Kabalnov

et al. (Equation 1) to describe the emulsion phase separation time periods is shown versus 1/T

for baseline (0.1 M SDHS) emulsions at HLD=0.4.

DrainageAggregation Coalescence

(a) y = 1E-10e9141x

R² = 0.98

y = 3E-06e5648x

R² = 0.55

y = 2E-05e5391x

R² = 0.94

1

10

100

1000

10000

100000

0.003 0.0032 0.0034 0.0036 0.0038 0.004

Tim

e (

Se

co

nd

s)

1/T (1/K)

(b)

0

10

20

30

40

50

-1.5 -1 -0.5 0 0.5 1 1.5

HLD

W*(k

BT

)

Type I Type III Type II

Figure 5.10: (a) Sample W* fittings for the aggregation, drainage, and coalescence of Type I

baseline (0.1 M SDHS) emulsion at HLD=-1 and (b) their values as a function of the HLD at

298 K.

120

The exponential factors taken from these regressions are in turn used to calculate W* for

aggregation, drainage, and coalescence as presented in Figure 5.10 (b) in equivalent kBT units at

298 K and as a function of the HLD. From this figure, it appears that W* for all these time

periods tends to decrease, and reach a minimum, towards HLD=0. Although these W* are

consistent with previous observations regarding the stability of Type III emulsions close to

HLD=0, they are not in complete agreement with the modeled predictions of Kabalnov et al. for

Type I and Type II emulsions16,27

. These researchers suggested that W* is constant for Type I

and Type II emulsions. Another interesting feature in Figure 5.10 (b) is that W* for the

aggregation of Type I emulsions is considerably greater than that for their drainage and

coalescence. This finding is consistent with the previous observation in Figure 5.8 (b) that the ap

time period generally seems to control the separation of Type I emulsions. It is important to

keep in mind that although these values produce reasonable estimates for W* around the phase

inversion point, they do not reflect any changes in the pre-exponential factor f of Equation 1.

5.5 DISCUSSIONS

It was previously mentioned that the ap time period likely involves the aggregation of

emulsion droplets. To evaluate this possibility, one can assume that the Brownian motion of

emulsion droplets is what leads to the collisions responsible for this overall process. An estimate

for the experimental ap time period can thus be obtained from the characteristic diffusion time

(tdiff). To calculate tdiff, the characteristic diffusion length ( diffDtl 2 ) is taken together with

the Einstein-Stokes relation for the diffusivity of an emulsion droplet after mixing

(E

dE

B

d

TkD

3 ) and is in turn set proportional to E

dd as follows:

121

Tk

dt

B

E

dE

diff4

33

(Eq. 9)

The assumption of l being proportional to E

dd is believed to be valid as E

dE

d

dl 3

3

4

2

1

in

both dilute and concentrated emulsions (where E

d is the emulsion droplet volume fraction). In

Figure 5.11, the ap time period and tdiff are presented as a function of the experimental E

dd for

baseline (0.1 M SDHS) emulsions at 25°C (data taken from Figure 5.5 (b)).

y = 72.3x2.5

1

10

100

1000

10000

100000

0 2 4 6 8

d32 (μm)

ap

or

t dif

f(S

ec

on

ds

)

Figure 5.11: Dependence of the ap time period for baseline (0.1 M SDHS) emulsions at 25°C on

E

dd . The solid line represents the fit of tdiff from Equation 9 (R2=0.84).

Despite having neglected hydrodynamic interactions, the power correlation in Figure

5.11 presents a fitted exponent of 2.5. This fitted exponent is fairly close to the predicted value

ddE (µm)

122

of 3 in Equation 9. The dispersion in the data is pronounced at larger values of E

dd . For Type I

emulsions, the power correlation underpredicts the experimental ap time period by 2 orders of

magnitude. This observation suggests that the aggregation of Type I emulsions is perhaps

hindered at low electrolyte concentrations by repulsive electrostatic forces. Furthermore, it is

likely that these emulsion droplets must collide numerous times before they can form large

enough aggregates for settling and film thinning to continue the separation process. For Type II

emulsions, the power correlation only slightly overpredicts the experimental ap time period.

The additional likelihood of such growth in the emulsion droplet aggregate size over the

ap time period can be evaluated by calculating the apparent average emulsion droplet diameter

( E

appdd , ) at the onset of settling. This can be done with the aid of Stokes’ law as follows10,33

:

E

E

appdE

E

d

s

dgv

18

2

, (Eq. 10)

In this equation, vs is the settling rate, E

d is the density of the emulsified phase, and g is

the gravitational acceleration constant (9.81 m/s2). By taking the vs of baseline (0.1 M SDHS)

emulsion droplets at 25°C as the net rate of advance of the µE phase separation front (Figure

5.7), ηµE as 1 cP, and Δρ ( E

d -ρµE) as 65 kg/m3, E

appdd , is calculated as presented in Figure 5.12.

The Reynolds number associated with E

appdd , is on the order of 10-3

, which is consistent with the

use of Stokes’ law. One limitation of Stokes’ law is that it does not consider hindered settling. If

one were to consider this phenomenon by introducing the appropriate empirical correction factor

(e.g. (1- E

d )4.5

), even larger E

appdd , than those presented in Figure 5.12 would need to be

produced34

. The data of Figure 5.12 proves that the emulsion droplets formed upon mixing do

123

not settle on their own and that the aggregation of pre-coalesced emulsion droplets that takes

place over the ap time period is in fact a necessary prerequisite. These conclusions are consistent

with the observations of others35,36

.

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

HLD

Emulsion Droplet Diameter (μm)

Initial (Measured) Settling (Equation 10) Film Thinning (Equation 11)

Type I Type III Type II

Figure 5.12: Comparison of the originally measured E

dd for baseline (0.1 M SDHS) emulsion

droplets at 25°C after mixing and the estimated E

appdd , required for settling (Equation 10) and

film thinning (Equation 11) as a function of the HLD.

Any further growth in E

appdd , due to gravitationally-induced collisions occurring over the

dp time period can be calculated by considering the process of film thinning as follows37,38

:

223

3

13

4

ratioratioEc

HowH

ffR

t

dt

dt

(Eq. 11)

124

In this expression, which was derived by taking Reynolds’ equation together with the

Laplace pressure at the Plateau border (neglecting DLVO forces), -dtH/dt is the rate of film

thinning, tH is the film thickness, Rc (estimated as E

appdd , /2) is the radius of curvature at the

Plateau border, and fratio is the ratio between the radius of the circumference of contact between

emulsion droplets (or the radius of the film) and Rc. By considering a dodecahedron foam-like

structure, fratio=0.6. To estimate tH, a simple ratio of the volume of the continuous µE phase to

the total area of the originally formed emulsion droplets is calculated. An initial estimated range

for tH of 0.2-0.9 µm is calculated as a function of the HLD for baseline (0.1 M SDHS)

emulsions at 25°C. At these separation distances, the magnitude of the DLVO forces acting to

squeeze toluene and water emulsion droplets together is only 0.1%-1% of the total Laplace

pressure38

. An estimate for -dtH/dt is further calculated by taking a ratio between the flow rate of

the liberated µE phase (determined from Figure 5.7 as the net rate of advance of the µE phase

separation front) and the area of the originally formed emulsion droplets. Using Equation 11 and

all of the values listed above, Rc can be calculated as a function of the HLD. These values are

presented in Figure 5.12 as E

appdd , =2Rc. From this, the calculated E

appdd , appear substantially

larger than the values calculated based on vs in Equation 10 and are close to 2 orders of

magnitude larger than the originally measured E

dd in Figure 5.5 (b). According to Figure 5.12, a

foam-like structure should be formed at HLD=0.9 with a characteristic cell size of ~1 mm. As

can be observed from the series of pictures in Figure 5.6, such structures do become apparent in

the latter stages of film thinning.

To evaluate the hole nucleation theory of coalescence, W* for baseline (0.1 M SDHS)

emulsions at 25°C (or 298 K) in Figure 5.10 (b) are plotted as a function of γow in Figure 5.13.

125

0

10

20

30

40

50

0 0.2 0.4 0.6 0.8 1

W*(k

BT

)

γow (mN/m)

DrainageAggregation Coalescence

HLD=-0.4

HLD=-0.7

Figure 5.13: General correlation between all of the fitted W* and the measured γow for baseline

(0.1 M SDHS) emulsions at 25°C.

The data in Figure 5.13 suggests that the apparent activation energy, W*, for drainage

and coalescence increases linearly with increasing γow as predicted by the hole nucleation

theory. This suggests that coalescence plays an important role in controlling these phase

separation time periods. If coalescence is the rate limiting process for the dp and cp time periods,

this would be consistent with the fact that the rate of advance of the µE and excess phase

separation fronts is almost the same. On the other hand, W* for aggregation deviates from the

linear trend predicted by the hole nucleation theory for Type I emulsions at intermediate

HLDs<0. This deviation is not surprising as the ap time period seems to be more or less

controlled by the previously described Brownian collisions. The larger W* for aggregation at

HLD<0 suggests that electrostatic forces strongly oppose the formation of flocs.

126

The emulsion phase separation studies carried out at 0.01 M SDHS and 25°C provide

some insight into the effect of the surfactant concentration. One of those effects is that reducing

the surfactant concentration from 0.1 M to 0.01 M SDHS reduces the ap time period by more

than 1 order of magnitude at the HLD extremes. It is unlikely that the presence of surfactant

would modify the frequency of emulsion droplet collisions as ηµE is almost the same. Also, γow

at these surfactant concentrations should be the same since they both lie above the critical µE

concentration (cµc) of 8 mM for SDHS-toluene-water µEs32

. Therefore, according to Equation

8, E

dd at both surfactant concentrations should be the same. The difference in the final emulsion

stability results is likely associated with the probability of these Brownian collisions leading to

the formation of flocs. At a higher surfactant concentration, more µE droplets are present in the

film that separates emulsion droplets. In squeezing out this inter-droplet film to allow for

coalescence to take place, depletion-repulsion interactions result from the increase in the

osmotic pressure associated with the overlap of structured µE droplet zones25

. The magnitude of

these interactions is presented in Chapter 7.

In comparing the emulsion phase separation time periods at 0.1 M and 0.3 M SDHS and

25°C (Figure 5.8), the ap and ep time periods appear to slightly decrease with an increase in the

surfactant concentration for Type I emulsions. This suggests that Ostwald ripening may play a

role in helping for these emulsion droplets to grow and more quickly assemble into an aggregate

size at which subsequent drainage can occur. It is expected that the prolonged dp time period

observed at an increased surfactant concentration for these same emulsions is a result of the

increased number of metastable µE layers that need to be expelled during film thinning. The

contributions of an increased ηµE leading to a decreased vs should be minimal.

127

5.6 CONCLUSIONS

The objective of this chapter was to study and describe how the stability of emulsions

prepared with an anionic surfactant (sodium dihexyl sulfosuccinate, SDHS), oil (toluene), and

water change around the phase inversion point of corresponding microemulsions (µEs). To do

so, emulsion phase separation profiles were described in terms of aggregation (ap), drainage

(settling + film thinning) (dp), and coalescence (cp) time periods. These phase separation time

periods were assessed as a function of the hydrophilic-lipophilic deviation (HLD). Here, the

phase inversion point lies at HLD=0. At a low surfactant concentration (0.01 M SDHS), the cp

time period dominated for oil (HLD<0) and water (HLD>0) emulsion droplets. At larger

surfactant concentrations, the ap time period dominated at HLD<0 whereas all of the time

periods were more comparable at HLD>0. The ap time period seemed to be controlled by the

frequency of Brownian collisions that effectively led to the formation of aggregated flocs of

larger emulsion droplets. The frequency of these underlying Brownian collisions was hindered

by the presence of µE droplets in the emulsion’s continuous phase. The dp and cp time periods

seemed to be controlled to a greater extent by the coalescence of neighboring emulsion droplets.

The apparent activation energies (W*) associated with these latter phase separation time periods

was proportional to the oil-water interfacial tension, suggesting that they can be explained by

the hole nucleation theory for emulsion coalescence.

This work illustrates the amount of information that can be obtained from relatively

simple phase separation studies, as well as the risks of oversimplifying the interpretation of the

data by just using a single equation (e.g. settling velocity). Having differentiated the stability

analysis of emulsions according to their different phase separation time periods and associated

W* further opens the possibility for a more detailed modeling of these processes.

128

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133

CHAPTER 6:

PREDICTING THE MORPHOLOGY AND VISCOSITY OF

MICROEMULSIONS USING THE NAC MODEL

This chapter is derived from the following published manuscript:

Kiran, S.K.; Acosta, E.J. Predicting the Morphology and Viscosity of Microemulsions using the

HLD-NAC Model. Ind. Eng. Chem. Res. 2010, 49, 3424-3432.

134

6.1 ABSTRACT

This chapter focused on extending the net-average curvature (NAC) model to help

predict the shape and viscosity (ηµE) of anionic surfactant (sodium dihexyl sulfosuccinate

(SDHS))-oil (toluene)-water microemulsions (µEs) from readily available formulation

parameters. To do so, a new shape-based NAC model was introduced which related the net and

average curvatures to the length and radius of µE droplets possessing a hypothesized cylindrical

core with hemispherical end caps. Knowing the shape of these µE droplets, theoretical scattering

profiles and maximum hydrodynamic radii were predicted. Furthermore, considering the

predicted volume fraction of the dispersed µE droplets alongside their shape allowed for the

accurate prediction of ηµE. It was found that treating the µE phase as a dilute suspension of rigid

rods yielded predicted ηµE close to experimental values near the bicontinuous phase transition

limits. These correlations were further extended to published experimental ηµE data of non-ionic

surfactant systems. The predicted µE morphology and ηµE may be useful in the design of

formulations for nanoparticle synthesis, enhanced oil recovery, and various environmental

remediation technologies.

6.2 INTRODUCTION

As has been highlighted in earlier chapters, there are three basic microemulsion (µE)

types: oil-in-water (o/w) µEs in equilibrium with an excess oil phase (Type I), bicontinuous µEs

in equilibrium with excess oil and aqueous phases (Type III), and water-in-oil (w/o) µEs in

equilibrium with an excess aqueous phase (Type II). Various researchers have experimentally

demonstrated that the length to radius ratio of µE droplets varies significantly according to

formulation conditions1,2,3,4,5

. The resulting changes in µE droplet shape and size, which arise

135

from underlying changes to the surfactant packing at the interface, have a profound effect on the

manufacturing of nanoparticles, latexes, and other nanostructured fluids6,7

. Another important

property of µEs is their viscosity (ηµE). As was previously noted by Trickett et al., ηµE ranges

from a fluid-like state to a solid-like gel8. Examples of instances where ηµE is an important

parameter to consider, and control, include in soil remediation, oil recovery, alternative fuels,

and hydraulic fluids9,10

.

The size of µE droplets has been estimated by Eastoe and others from the ratio of the

dispersed phase to the mass of the surfactant11,12,13

. Furthermore, Acosta et al. have predicted the

size of µE droplets using the inverse of their sphere-equivalent radius of curvature as calculated

from the net-average curvature (NAC) model14,15

. While there are a number of theoretical

equations dedicated to explaining the shape and changes in shape of µE droplets, thus far there

are no theoretical or empirical equations that can predict the shape of µE droplets from

formulation conditions (e.g. type of surfactant and oil, electrolyte concentration, and

temperature) 16,17,18,19,20,21,22,23,24,25,26

. The first objective of this chapter is to extend the NAC

model to predict the shape of Type I and Type II µE droplets. The predicted shape will be used

to generate theoretical small angle neutron scattering (SANS) profiles that will be compared to

experimental scattering profiles. Furthermore, the predicted shape will also be used to estimate

the maximum hydrodynamic radius of µE droplets. The second objective of this chapter is to

use the predicted volume fraction ( E

d

) and shape of µE droplets to estimate ηµE.

Numerous correlations have been established by several researchers to help predict ηµE.

Einstein first developed the following model for the dilute dispersion of hard spheres within a

liquid medium9:

136

E

d

E

cE

5.21 (Eq. 1)

In this equation, E

c

is the viscosity of the µE’s continuous phase. An extension of this

model was later proposed by Taylor et al. for the dilute suspension of liquid spheres within a

reference continuous phase9,27

:

E

d

E

c

E

d

E

cE

d

E

cE

5.21 (Eq. 2)

In Equation 2, E

d

represents the viscosity of the µE’s dispersed phase. For the hard-

sphere scenario where E

d

>> E

c

, Taylor’s model simplifies to that of Einstein. To describe ηµE

under more concentrated regimes ( E

d

>0.02), Saito et al. derived the following expression by

considering the hydrodynamic interactions between non-correlated spheres9,28

:

E

d

E

dE

cE

1

5.11 (Eq. 3)

A shortcoming of Equation 3 is that it does a poor job of replicating experimental ηµE

peaks at the maximum packing condition of dispersed droplets. This was corrected for in the

Dougherty-Krieger relationship as follows9,29,30

:

Ed

E

d

E

dE

cE

max,5.2

max,

1

(Eq. 4)

In Equation 4, E

d

max, reflects the maximum volume fraction of rigid dispersed spheres,

which ranges from 0.6 to 0.7 according to the shear rate31

. In addition to Equations 1-4, Thomas

137

et al. established an empirical expression for predicting the ηµE of concentrated µEs by

expanding Einstein’s model to higher orders of E

d

9,32:

E

d

E

d

E

d

E

cE

6.16exp00273.005.105.21

2 (Eq. 5)

A key assumption in the derivation of Equations 1-5 is that the shape of the dispersed µE

droplets is spherical. As was previously described, however, cylindrical and elliptical µE

droplets are also common. Doi and Edwards treated the case for a dilute and semi-dilute

concentration of rigid rods by taking into account their rotary diffusion in Equations 6 and 7

respectively9,33

:

E

E

rods

E

rodsgE

c

E

rods

E

cE

d

lclN

2

2

3 411 , where

31

E

rodslN

(Eq. 6)

E

rods

E

rodsE

rodsE

E

rodsgE

cE

d

ld

lc

ln5

96

632

63

, where 23

11

E

rods

E

rods

E

rods ldN

l (Eq. 7)

In these equations, N is the number density of rigid µE rods, E

rodsl and E

rodsd are their

respective length and diameter, cg is their concentration in mass per unit volume, and E is the

density of the µE phase. Similar correlations have yet to be developed for liquid rods.

Additional models for µEs consisting of prolate and oblate ellipsoids as the dispersed phase are

as in Equations 8 and 9 respectively9,34

:

138

15

14

5.12

ln155.02

ln5

1

22

E

ell

E

ell

E

ell

E

ell

E

ell

E

ell

E

ell

E

ell

E

d

E

cE

b

a

b

a

b

a

b

a

(Eq. 8)

E

ell

E

ell

E

ell

E

ell

E

d

E

cE

b

a

b

a

arctan

15

16

1 (Eq. 9)

Equations 8 and 9 are applicable when the ratio of the ellipsoid’s major axis ( E

ella ) to

minor axis ( E

ellb ) is greater than 10.

The E

d

and shape of the dispersed µE droplets predicted by the NAC model will be

incorporated into Equations 1-9 to predict the ηµE of Type I and Type II anionic surfactant

(sodium dihexyl sulfosuccinate (SDHS))-oil (toluene)-water µEs. The predicted ηµE will be

compared to experimental results obtained for similar µEs.

The NAC model allows for the prediction of µE properties (e.g. solubilization capacity,

phase volumes, phase transitions, and interfacial tensions) as a function of thermodynamic

formulation parameters (e.g. electrolyte concentration, temperature, and oil and surfactant

hydrophobicity)14,15

. A key element of this model is the calculation of the hydrophilic-lipophilic

deviation (HLD). As was previously described in earlier chapters, the HLD for ionic surfactants

can be expressed as follows:

coc CTAfNSHLD 01.017.0ln , (Eq. 10)

139

In this equation, S is the electrolyte concentration, Nc,o is the equivalent alkane carbon

number, f(A) is a function of the co-surfactant type and concentration, ΔT is the temperature

deviation from 25°C, and Cc is the characteristic curvature of the ionic surfactant. The HLD

equation for non-ionic surfactants differs ever so slightly15,35

:

noc CTANSbHLD 06.017.0 , (Eq. 11)

In this equation, b(S) accounts for the “salting” out of the non-ionic surfactant. Published

values of b include 0.1 for CaCl2 and 0.13 for NaCl. Also, ϕ(A) represents the effect of alcohol

(co-surfactant) and Cn is the characteristic curvature of the non-ionic surfactant. For both

Equations 10 and 11, a HLD<0HLD=0HLD>0 scan reflects a Type IType IIIType II

µE phase behavior transition.

To obtain the desired physical properties of µEs, the HLD is included within the NAC

model as a scaling parameter for the net curvature (Hn) of the surfactant film at the oil-water

interface14,15

:

L

HLD

rrH

E

w

E

o

n

11 (Eq. 12)

In Equation 12, E

or and E

wr are the sphere-equivalent radii of hypothetically co-existing

oil and water µE droplets whereas L is a length scaling parameter that is proportional to the

extended length of the surfactant tail (~1.2 times). Hn is 0 for bicontinuous (Type III) µEs

containing similar amounts of oil and water ( E

or = E

wr ), >0 for Type I µEs ( E

or << E

wr ), and <0

for Type II µEs ( E

or >> E

wr ). The average curvature term (Ha) in the NAC model is used to

describe the size of the µE’s oil and water domains14,15

:

140

111

2

1

E

w

E

o

arr

H (Eq. 13)

The inequality on the right-hand size of Equation 13 suggests that the size of µE droplets

is limited by their characteristic length (), a concept that was previously introduced by De

Gennes et al.36

. These authors demonstrated that this parameter, which represents the distance at

which an oil or water molecule can be separated from a surfactant membrane and still interact

with it, is calculated as follows14,15,36

:

s

E

E

w

E

o

A

V

6

(Eq. 14)

Here, E

o

and E

w

are the respective oil and water volume fractions within the µE phase

of volume EV . At HLD=0, is a maximum and serves as a benchmark for the transition to a

bicontinuous µE.

6.3 DEVELOPMENT OF EXPRESSIONS FOR THE SHAPE-BASED

NAC MODEL AND MAXIMUM HYDRODYNAMIC RADIUS

To predict the shape of Type I and Type II µE droplets, it is assumed that these droplets

have a cylindrical neck region of length E

dl with hemispherical end caps of radius E

dr . In this

way, a smooth transition from spheres ( E

dl =0) to rods ( E

dl >> E

dr ) can be obtained. The total

integrated curvature of these droplets at the oil-water interface is equivalent to the surface area-

averaged curvature (CSA,avg) whereby the curvature of the cylindrical neck region is 1/2 E

dr and

the curvature of the hemispherical end caps is 1/ E

dr :

141

AreaSurfaceTotal

HemisphereofAreaSurfacer

CylinderofAreaSurfacer

C

E

d

E

d

avgSA

12

2

1

, (Eq. 15)

As demonstrated in Appendix 1, a new revised net curvature term ( '

nH ) is a reasonable

approximation of the sphere-equivalent curvature of these droplets. Incorporating this parameter

within Equation 15 yields the following expression in terms of E

dr and E

dl :

E

d

E

d

E

d

E

d

E

d

E

d

n

nrlr

l

rl

HH

42

1

2

2

2

'

(Eq. 16)

Using neutron scattering data, it has been shown that Ha represents the sphere-equivalent

surface area to volume ratio of µE droplets37

. Ha can therefore be related to E

dr and E

dl as

follows:

243

42

E

d

E

d

E

d

E

d

E

da

rrl

rlH

(Eq. 17)

Obtaining Hn (and thus '

nH ) and Ha from the original NAC model, the shape of the µE

droplets can be determined by simultaneously solving Equations 16 and 17 for E

dr and E

dl .

Furthermore, these solved values of E

dr and E

dl , as well as the surfactant tail length (δtail), can

be used to estimate the maximum hydrodynamic radius of rigid µE droplets ( E

hr

max, ) as follows:

tail

E

dE

d

E

h

lrr

2max, (Eq. 18)

The actual hydrodynamic radius ( E

hr ) is only equal to the calculated

E

hr

max, for rigid and

fully extended µE droplets. Droplets demonstrating an experimental E

hr smaller than

E

hr

max, (as

142

determined via light scattering) are flexible. Such types of light scattering experiments have

been performed in the past to assess the flexibility of sodium dodecyl sulfate (SDS) micelles38

.

6.4 MATERIALS AND METHODS

6.4.1 Materials

The following materials were used as purchased from Sigma-Aldrich Corp. (Oakville,

ON, Canada): anhydrous toluene (product #244511, 99.8%), SDHS (product #86146, ~80% in

water), and NaCl (product #S9625, ≥99.5%). Acetonitrile (product #1401-7-10, HPLC grade)

was obtained from Caledon Laboratory Chemicals (Georgetown, ON, Canada) and deionized

water (conductivity <3 µs/cm) was prepared in the laboratory using an anion exchange resin.

6.4.2 Microemulsion Phase Behavior Scans

A series of SDHS-toluene-water µE phase behavior scans at a SDHS concentration of

0.1 M and 1-10 g/100 mL NaCl in the aqueous phase was prepared according to the procedure

outlined in Chapter 5.

6.4.3 Oil-Water Solubilization

Solubilization of toluene in water was measured experimentally using a Dionex 3000

ionic chromatographic system equipped with a C18 column. The mobile phase, which consisted

of water and acetonitrile in a 55 to 45 volume ratio, was pumped through the column at a flow

rate of 1 mL/min. An AD25 absorbance detector was used to detect toluene at a wavelength of

220 nm. Water solubilization in the oil phase was measured using a Kam Control Inc. Karl

Fischer moisture analyzer.

143

6.4.4 Viscosity Measurements

A Gilmont Instruments falling ball viscometer (model #GV-2200) was used in order to

target ηµE measurements <10 cp. The vertical glass tube apparatus was first filled with ~4.5 mL

of the µE sample of interest. Once filled, a glass ball of diameter 0.6 cm was released within the

tube and its time of descent (td, in minutes) between two sets of red reference markings was

measured using a stopwatch. A glass ball was used here as opposed to a heavier stainless steel

ball in order to reduce the shear rate ( ) and minimize the possible risk of shear-thinning

behavior reported by Bennett et al.39

. The maximum applied was limited to 0.1 seconds-1

. For

each sample, triplicate measurements were performed. The ηµE was subsequently calculated in

units of centipoise (cP) by taking into consideration the density of the glass ball (2.53 g/mL), the

density of the µE phase (ρµE, in g/mL), and the viscometer constant (3.3):

EdE t 53.23.3 (Eq. 19)

6.4.5 Dynamic Light Scattering

Dynamic light scattering (DLS) measurements were performed at 25°C using a

Brookhaven Instruments Corp. 90Plus Particle Size Analyzer. This instrument is equipped with

a 15 mW solid state laser (wavelength of 635 nm) and a detector positioned at a scattering angle

of 90° with respect to the incident light beam. To operate, standard glass cuvettes were filled

with the µE samples 15 minutes prior to being inserted within the sample chamber for

analysis40

. Droplet sizes were measured on the basis of decay times of fluctuating intensity

readings. Reference viscosities and refractive indices are provided in Table 6.1.

144

Table 6.1: Reference viscosities () and refractive indices (n).

Parameter Value

water (25°C) 0.89 cP41

toluene (25°C) 0.56 cP42

nwater (25°C) 1.3343

ntoluene (25°C) 1.4944

6.4.6 Prediction of Oil and Water Solubilization with the NAC Model

To calculate Hn for Type I (Type II) µEs, a fictitious sphere-equivalent radius of the

continuous aqueous (oil) phase ( E

IIIowr

, ) was first calculated as follows14,15

:

s

E

IIIowE

IIIowA

Vr

,

,

3 (Eq. 20)

si

E

wsis aVcA 231002.6 (Eq. 21)

In Equations 20 and 21, E

IIIowV , represents the continuous volume of water (oil) within

the Type I (Type II) µE phase, As is the total surfactant interfacial area, and csi and ai represent

the concentration and surface area per molecule of surfactant or co-surfactant species i pre-

dissolved within the aqueous phase. Here, the critical micelle concentration (cmc) of non-

adsorbed surfactant is neglected since the total surfactant concentration is more than 1 order of

magnitude larger14,15

. By establishing E

IIIowV , and csi as experimental constants and using

literature values for ai, E

IIIowr

, was calculated as per Equations 20 and 21. Knowing this

parameter as well as L, and having calculated the HLD using either Equation 10 or 11, the

sphere-equivalent radius of the dispersed oil (E

Ior

, ) and water (E

IIwr

, ) droplets for Type I and

Type II µEs were respectively estimated through the use of Equation 12. The volumes of oil

(E

IoV , ) and water (

E

IIwV , ) solubilized within the µE phase were respectively solved for by subbing

145

E

Ior, and E

IIwr

, back into Equation 20. For Type III µEs, Ha=1/ and Equations 12 and 13 were

solved simultaneously for E

IIIor, and E

IIIwr, . Oil and water solubilization for SDHS-toluene-water

µEs across a Type IType IIIType II phase behavior scan was determined from the list of

HLD and NAC modeling parameters in Table 6.2.

Table 6.2: HLD and NAC modeling parameters for SDHS-toluene-water µEs.

Parameter Value

S (g/100 mL) 1-10 g/100 mL

Nc,o 1

f(A) 0

∆T 0

Cc -0.92

L (Å) 10 Å14,37

E

IIIowV , (mL) 5 mL

csi 0.1 mol/L

ai 100 Å2/molecule

14

(HLD≈0) 68 Å37

This same approach was used to calculate E

Ior

, and E

IoV , for Type I µEs of the non-ionic

surfactant (C12E4)-oil (hexadecane)-water and non-ionic surfactant (C12E5)-oil (cyclohexane and

hexadecane in a 1 to 1 weight ratio)-water systems studied by Leaver and Olsson28

. The list of

HLD and NAC modeling parameters for these µEs is summarized in Table 6.3.

Table 6.3: HLD and NAC modeling parameters for the non-ionic µEs of Leaver and Olsson28

.

Parameter C12E4-Hexadecane-Water C12E5-Cyclohexane and Hexadecane-Water

S (g/100 mL) 0 0

Nc,o 1615,45

9

ϕ(A) 0 0

∆T (-8) to (+2) (-5) to (+3)

Cn 1.815

0.815,46

L (Å) 2515

2515

E

IwV , (mL) 73 83

csi (M) 0.4 0.2

ai (Å2/molecule) 54

15 60

15

146

From the above calculations for the different µEs of interest, E

dr and E

dl were

calculated. These values were then used to predict ηµE.

6.4.7 Prediction of Small Angle Neutron Scattering Profiles

The small angle neutron scattering (SANS) profiles of cylindrical droplets with

hemispherical end caps were predicted by incorporating the NAC model within the theoretical

profile developed by Cusack et al.47,48

:

2/

0

2 sin

dqAqI E

d (Eq. 22)

In this equation, I(q) is the scattering intensity, q is the scattering vector, θ is the

scattering angle, and A(q) represents the scattering amplitude defined by:

1

0

3

3

312)(sin4

)cos()sin(cos4

sin

sin2sindttfXr

qr

qrqrqrXr

qr

qrJ

X

Xlr

b

qA E

dE

d

E

d

E

d

E

dE

dE

d

E

dE

d

E

d

(Eq. 23)

In Equation 23, X=sin(q( E

dl /2)cosθ), f(t)=(1-t

2)sin(q E

dr tcosθ)[J1(q

E

dr sinθ(1-

t2)0.5

)]/(q E

dr sinθ(1-t

2)0.5

), and Δb is the difference in the scattering length density between the

µE’s dispersed and continuous phases.

For both the Type I and Type II SDHS-toluene-water µEs, scattering profiles were

generated by predicting E

dr and E

dl using Equations 16 and 17. The predicted profiles were

then compared to experimental scattering profiles obtained from Acosta et al. for similar

systems37

. For Type I scattering profiles, which were obtained using deuterated toluene

solubilized in the micelles, Δb=4.8×10-6

Ǻ-2

. For Type II µEs, deuterated water was solubilized

in the reverse micelles and Δb=5.4×10-6

Ǻ-2

.

147

6.5 RESULTS AND DISCUSSIONS

6.5.1 Comparison of Spherical Viscosity Models and Experimental

Measurements

As was previously described, predicting ηµE for µEs composed of spherical droplets

requires the knowledge of E

d

. Figures 6.1 (a) and (b) present the experimental and predicted

E

IoV , and E

IIwV , as well as E

d

as a function of the HLD. This data suggests that the experimental

and predicted data are in good agreement. Similar agreements have been reported in the past for

other ionic and non-ionic surfactant µEs2,14,15

.

0

0.5

1

1.5

-1.5 -1 -0.5 0 0.5 1 1.5

HLD

(b)(a)

0

0.2

0.4

0.6

0.8

1

-1.5 -1 -0.5 0 0.5 1 1.5

0

0.5

1

1.5

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3(mL

)

HLD

(b)(a)

0

0.2

0.4

0.6

0.8

1

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3NAC Model

NAC Model

E

d

E

d

(Oil)

(Water)

0

0.5

1

1.5

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3(mL

)

HLD

(b)(a)

0

0.2

0.4

0.6

0.8

1

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3NAC Model

NAC Model

E

d

E

d

(Oil)

(Water)

0

0.5

1

1.5

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3(mL

)

HLD

(b)(a)

0

0.2

0.4

0.6

0.8

1

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3NAC Model

NAC Model

E

d

E

d

(Oil)

(Water)

0

0.5

1

1.5

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3(mL

)

HLD

(b)(a)

0

0.2

0.4

0.6

0.8

1

-1.5 -1 -0.5 0 0.5 1 1.5

Y-Values

Column1

Column3NAC Model

NAC Model

E

d

E

d

(Oil)

(Water)

Figure 6.1: Experimental and predicted (a) E

IoV , and E

IIwV , as well as (b) E

d

as a function of the

HLD for SDHS-toluene-water µEs.

The ηµE can be calculated by substituting the predicted E

d

, in addition to the volume

fraction of surfactant, as well as E

c

(from Table 6.1) into Equations 1-5. Figure 6.2 presents

148

the measured E

c

along with the values predicted using Equations 2 (for dilute liquid spheres)

and 3 (for concentrated hard spheres). The values predicted using Equations 1, 4, and 5 all fall in

between the calculated ηµE using Equations 2 and 3. When compared to the experimental data, it

is apparent that away from the Type I-Type III and Type II-Type III transitions (HLD<-0.5 and

HLD>0.5 respectively) is where the predicted ηµE assuming spherical droplets is most

reasonable. The predicted ηµE are however substantially lower than the measured values in the

transition regions. The inaccuracy of these spherical droplet models near the transition regions

has been reported before9. It has been proposed that the ηµE models for spherical droplets fail as

they severely underestimate the interaction amongst rod-like µE droplets.

1

2

ηµ

E (

cP)

5

0.5

-1.5 -1 -0.5 0 0.5 1 1.5

Experimental

Dilute liquid

spheres (Eq. 2)

Concentrated hard

spheres (Eq. 3)

Figure 6.2: Experimental and NAC modeled (for dilute liquid spheres (Equation 2) and

concentrated hard spheres (Equation 3)) ηµE of SDHS-toluene-water µEs as a function of the

HLD.

The new shape-based NAC model can potentially resolve this issue since the predicted

E

dr and E

dl of the µE droplets can be used to calculate the ηµE of non-spherical droplets. The

149

E

dr and E

dl as well as the total aspect ratio ( E

dl /2 E

dr ) of SDHS-toluene-water µE droplets as

predicted by the NAC model are displayed in Figures 6.3 (a) and (b) as a function of the HLD.

Consistent with the trends predicted by conceptual models, and previous experimental

observations, E

dl of the µE droplets increases upon transitioning into the bicontinuous

region19,26

. Such elongated structures demonstrate enhanced interactions via overlapping which

results in quasi-stable clusters of droplets49,50

. Predictions of ηµE for bicontinuous µEs will be

omitted here as their complex interconnected structural network is not yet well understood and

falls outside the scope of this chapter.

(Å)

Asp

ect

Ra

tio

HLD

(b)(a)

Rd

Ld

(NAC Model)

(NAC Model)

0

3

6

9

12

15

-1.5 -1 -0.5 0 0.5 1 1.5

Figure 6.3: Predicted (a) E

dr and E

dl as well as (b) aspect ratio ( E

dl / E

dr ) as a function of the

HLD for SDHS-toluene-water µEs using the NAC model.

150

6.5.2 Comparison of Predicted and Experimental SANS Profiles for Type I and

Type II Microemulsions

The predicted values of E

dr and E

dl in Figure 6.3 are used to generate theoretical SANS

profiles for Type I and Type II µE droplets assuming a polydispersity value of 0.3 for E

dr ,

consistent with the average polydispersity obtained in DLS experiments. At a low electrolyte

concentration (1.2 g/100 mL NaCl where HLD=-0.9), Figure 6.4 (a) shows that the predicted

scattering curves have a transitional feature (related to E

dr ) at high q values (~0.15Å

-1), but that

transitional feature is observed at lower q values (~0.07Å-1

) in the experimental profile. From

this observation, one infers that the µE droplets are more spherical than cylindrical at low

electrolyte concentrations and that the NAC model underestimates their size. This

underestimation of the size of µE droplets at such a low electrolyte has been observed before37

.

Such deviation has been explained by the fact that the simple NAC model discussed here only

accounts for solubilization in the core of the micelles and not in their palisade layer. Aromatic

oils like toluene also solubilize in the palisade layer of the micelles, a phenomenon that has been

reproduced using a modified form of the NAC model51

.

151

0.01

0.1

1

10

100

0.01 0.1 1

0.01

0.1

1

10

100

0.01 0.1 1

0.01

0.1

1

10

100

0.01 0.1 1

0.01

0.1

1

10

100

0.01 0.1 1

0.01

0.1

1

10

100

0.01 0.1 1

0.01

0.1

1

10

100

0.01 0.1 1

q, Å-1

I(q

), c

m-1

(c)

q, Å-1

I(q

), c

m-1

I(q

), c

m-1

(b)

(a) (d)

(e)

(f)

1.2 g NaCl/100 ml

1.7 g NaCl/100 ml

2.0 g NaCl/100 ml

3.8 g NaCl/100 ml

4.3 g NaCl/100 ml

5.5 g NaCl/100 ml

Figure 6.4: Predicted SANS profiles of SDHS-toluene-water µEs formulated at (a) HLD=-0.9,

(b) HLD=-0.6, (c) HLD=-0.4, (d) HLD=0.3, (e) HLD=0.4, and (f) HLD=0.6 using the NAC

model.

As the electrolyte concentration increases, solubilization in the micelle core increases

and the scattering profiles predicted by the NAC model match more closely the experimental

scattering profiles, as shown in Figures 6.4 (b)-(c). For Type II µEs (Figures 6.4 (d)-(f)), the

predicted scattering profiles also produce a reasonable match with the experimental trends. The

good agreement between the predicted and experimental profiles supports the assumption and

approximations made in the new shape-based NAC model.

152

6.5.3 Comparison of Maximum Predicted and Experimental Hydrodynamic Radii

Values of E

hr

max, predicted via Equation 18 as well as experimentally measured E

hr using

DLS are plotted in Figure 6.5 as a function of the HLD. Although the predicted and

experimental trends are similar for Type I µEs, E

hr

max, consistently appears slightly larger than

the measured values. This indicates that the dispersed droplets are flexible as they do not fully

reach the maximum extended length. For Type II µEs, however, the predicted E

hr

max, more

closely traces the experimental data, suggesting that these cylinder-like droplets are able to

better maintain their shape as they randomly move throughout the continuous oil phase.

1

-1.5 -1 -0.5 0 0.5 1 1.5

(n

m)

Type I (Experimental)

Type II (Experimental)

NAC Model

Figure 6.5: Experimental E

hr and predicted

E

hr

max, for Type I and Type II SDHS-toluene-water

µEs as a function of the HLD.

153

6.5.4 Comparison of Non-Spherical Viscosity Models and Experimental

Measurements

The ηµE of elongated µE droplets can be calculated by incorporating the predicted E

d

in

Figure 6.1, the volume fraction of the surfactant, and the predicted E

dr and E

dl in Figure 6.3

into Equations 6-9. Figure 6.6 presents the same experimentally measured ηµE previously

presented in Figure 6.2 along with the predicted ηµE using Equations 6 (for dilute rigid rods) and

8 (for prolate ellipsoids). The values of ηµE predicted using Equation 7 (for semi-dilute rigid

rods) are substantially lower than the measured ηµE values for µEs away from the transition

regions and higher than the measured ηµE values near the transition regions. Equation 9 (for

oblate ellipsoids) produces ηµE values that are similar to those predicted for prolate ellipsoids. In

comparing Figures 6.2 and 6.6, it is shown that modeling the ηµE of SDHS-toluene-water µEs as

non-spherical droplets yields a closer approximation to the experimental data near the Type I-

Type III and Type II-Type III transition regions. Away from these transition regions, the

spherical models work best. Of the non-spherical models, that for dilute rigid rods is the most

accurate. This observation is believed to be a result of the fact that the condition N<1/ 3E

dl is

satisfied across all HLD values.

154

1

2

1.5 1 0.5 0 -0.5 -1 -1.5

5

ηµ

E (

cP

)

Experimental

Dilute rigid rods (Eq. 6)

Prolate ellipsoids (Eq. 8)

Figure 6.6: Experimental and NAC modeled (for dilute rigid rods (Equation 6) and prolate

ellipsoids (Equation 8)) ηµE of SDHS-toluene-water µEs as a function of the HLD.

The NAC model is also applied to the non-ionic µEs of Leaver and Olsson28

. For the

C12E4-hexadecane-water µEs, the NAC model predicts that the oil-swollen micelles take on the

shape of rods at 16°C and that these rods continuously lengthen with temperature. Figure 6.7 (a)

presents the resulting experimental viscosities for these µEs (plotted as a ratio of E / E

c

)

along with the predicted values obtained by applying the NAC model to dilute hard spheres

(Equation 1) and dilute rigid rods (Equation 6). Similar to the case for Type I SDHS-toluene-

water µEs, the dilute hard-sphere model greatly underestimates the experimental E / E

c

ratio

of these non-ionic µEs over 17-21°C where the maximum increase in E

d

is only 0.05. The

dilute rigid rods model produces a closer match with the experimental values. Using the semi-

dilute rigid rods model produces an overestimation of the E / E

c

ratio. An interesting feature

of Figure 6.7 (a) is the presence of an experimental E / E

c

maximum at 23°C. This maximum

coincides with the transition predicted by the NAC model where the rod structure is no longer

155

stable, yielding alternative morphologies with lower curvatures at higher temperatures. The

same trends are observed for the C12E5-cyclohexane and hexadecane-water µEs in Figure 6.7

(b). However, the predicted trend in the E / E

c

ratio for dilute rigid rods is shifted by about

3°C from the experimental data. This shift could be related to the value of Cn used in Table 6.3.

Using a value of Cn of 0.9 instead of 0.8 would produce a predicted ηµE curve for dilute rigid

rods that overlaps with the experimental data. This is a reasonable deviation considering that the

standard deviation in the estimated values of Cc and Cn range between ±0.1 and ±0.515,46

.

0

10

20

30

40

50

60

17 18 19 20 21 22 23 24 25 26 27

r

Temperature ( C)

Data of Leaver and Olsson

Dilute rigid spheres, Eq. 1

Dilute rigid rods, Eq. 6

0

5

10

15

20

25

30

20 21 22 23 24 25 26 27 28

r

Temperature ( C)

(a)

(b)

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

18

19

20

21

22

23

24

25

26

27

28

29

30

31

32

33

34

35

36

37

38

39

40

41

T(°C)

T(°C)

Figure 6.7: Experimental and predicted E / E

c

ratio for the (a) C12E4-hexadecane-water and

(b) C12E5-cyclohexane and hexadecane-water µEs of Leaver and Olsson28

.

156

While there are a number of gaps to still be filled to accurately predict ηµE (e.g.

bicontinuous and concentrated µEs), incorporating shape considerations can be used to estimate

the significant changes in ηµE that occur over a relatively narrow range of formulation

conditions near the phase transition regions.

6.6 CONCLUSIONS

The new shape-based NAC model effectively estimated the shape of microemulsion

(µE) droplets assuming that they possess a cylindrical core with hemispherical end caps. This

proposed structure was confirmed via neutron scattering profiles as well as DLS measurements

of hydrodynamic radii. Introducing the predicted µE droplet shape and volume fraction into pre-

existing viscosity (ηµE) models allowed for the ηµE of µEs to be estimated. For toluene-water

µEs produced with 0.1 M SDHS, it was determined that the model of Doi and Edwards for a

dilute concentration of rigid rods produced a reasonable prediction of ηµE for oil-in-water (or

Type I) and water-in-oil (or Type II) µEs, especially near the bicontinuous (or Type III)

transition limits. Similar trends were observed in an attempt to model the experimental ηµE data

of non-ionic µEs taken from the literature. The capability of predicting µE morphologies and

ηµE from formulation conditions may potentially be useful in applications including nanoparticle

synthesis, transdermal drug delivery, environmental remediation, and enhanced oil recovery.

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Tool for the Characterization of Cosmetic Hydrocarbon Oils. Polym. Int. 2003, 52, 602-609.

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Méthode du HLD. Tech. Ing. 2001, 157, 1-19.

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Aggregation Behavior of Viscoelastic Detergents. Adv. Colloid Interface Sci. 1982, 17, 275-298.

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164

CHAPTER 7:

MODELING THE SIZE AND STABILITY OF EMULSIONS

AROUND THE PHASE INVERSION POINT

165

7.1 ABSTRACT

In this chapter, the size and stability of emulsions around the phase inversion point was

estimated with the aid of the hydrophilic-lipophilic deviation (HLD) and net-average curvature

(NAC) models. The HLD and NAC models were used to predict the interfacial tensions,

densities, and viscosities of microemulsions (µEs) prepared with an anionic surfactant (sodium

dihexyl sulfosuccinate (SDHS)), toluene, and aqueous phase solutions of sodium chloride.

These predicted properties were introduced into known models that estimated the average

emulsion droplet size as a function of the mixing conditions. The predicted oil droplet sizes

compared well with those determined via optical microscopy. A simplified model of emulsion

stability, which was previously proposed in the literature, was used here to estimate the stability

of the emulsions formed after mixing. This emulsification model combined the assumption of

emulsion droplet diffusion-collision with an activation energy term for coalescence. This

activation energy term was calculated using the hole nucleation theory and by assuming that the

critical spacing between 2 approaching emulsion droplets was equal to the sphere-equivalent

diameter of their corresponding µE droplets. With the aid of the HLD and NAC models, the

stability of SDHS-toluene-water emulsions was reasonably predicted at surfactant

concentrations of 0.01 M, 0.1 M, and 0.3 M SDHS and temperatures ranging from 7°C to 44°C.

The largest deviations were observed at temperature and HLD extremes where the activation

energies might be dominated by factors that were not accounted for in the hole nucleation

theory. The HLD and NAC models were also used to predict literature data on the size and

stability of emulsions produced with other surfactants and oils. Overall, the applied models were

capable of producing reasonable estimates except for in situations that involved charge-

166

stabilized emulsion droplets and/or emulsion droplets stabilized by highly rigid surfactant

phases (e.g. liquid crystals).

7.2 INTRODUCTION

Large changes to the size and stability of emulsions dispersed throughout a

microemulsion (µE) continuum have been observed in Chapter 5 to occur around the

bicontinuous (or Type III) phase inversion point where oil-in-water (o/w, or Type I) µEs invert

into water-in-oil (w/o, or Type II) µEs. At this point, the interfacial tension (γow) reaches an

ultralow value, facilitating the production of submicron emulsions with a minimum energy

input. Furthermore, the stability of emulsions also reaches a minimum at this point. To make use

of these trends in the design of emulsification and phase separation processes, it is important to

be able to quantify the proximity to the phase inversion point. The most effective way of doing

so has been shown to be through the use of the following hydrophilic-lipophilic deviation

(HLD) models for ionic (Equation 1) and non-ionic (Equation 2) surfactants1,2,3,4,5,6

:

coc CTAfNSHLD 01.017.0ln , (Eq. 1)

noc CTANSHLD 06.017.013.0 , (Eq. 2)

In these equations, S is the NaCl concentration (in g/100 mL), Nc,o, for non-alkane oils,

is the equivalent alkane carbon number (EACN), f(A) and ϕ(A) are functions of the co-

surfactant type and concentration, ΔT is the difference in temperature from 25°C, and Cc and Cn

are the respective characteristic curvatures of ionic and non-ionic surfactants. Scanning

Equations 1 and 2 from HLD<0HLD=0HLD>0 reflects a Type IType IIIType II µE

phase behavior transition.

167

While the HLD model can be used to quantify the general approach of emulsions to the

phase inversion point (HLD=0), it fails to reveal how specific physical properties related to their

size and stability change. These properties, and more, can be estimated through the use of the

net-average curvature (NAC) model that was introduced for µEs in Chapter 6. The key

equations of the NAC model are as follows7,8

:

E

w

E

o

nrrL

HLDH

11 (Eq. 3)

E

w

E

o

arr

H

11

2

1 (Eq. 4)

In these equations, Hn and Ha are the respective net and average curvatures of the

surfactant at the oil-water interface, L is a scaling constant this is proportional to the extended

length of the surfactant tail (~1.2 times), and E

or and E

wr are the respective sphere-equivalent

radii of hypothetically co-existing oil and water µE droplets. Based on convention, a shift in

Equation 3 from Hn>0Hn=0Hn<0 corresponds to a HLD<0HLD=0HLD>0 transition.

By solving Equations 3 and 4, some of the continuous µE phase properties that can in turn be

predicted include solubilization capacities, phase transition and volumes, densities (ρµE), γow,

actual droplet shapes, and viscosities (ηµE).

The average emulsion droplet radius ( E

dr ) depends on the physical properties of the

system being considered and its mixing conditions. For turbulent mixing conditions in stirred

tanks (Re>10,000), E

dr can either be estimated via the energy dissipation per unit volume (εmix)

in Equation 5 or the impeller speed (Nimp) and diameter (Dimp) in Equation 69,10

:

168

5/15/2

5/3

1

Emix

owE

d

Cr

(Eq. 5)

5/35/45/6

5/3

2

Eimpimp

owE

dDN

Cr

(Eq. 6)

In these equations, C1 is a fitting constant that is close to 0.5 and C2 is 0.25 according to

the Hinze-Kolmogorov model. Given the power consumption (Pcons) and the total volume of the

emulsion (VE) in the mixer, ε can be estimated as Pcons/VE. Alternatively, Nimp and Dimp can be

used together instead of εmix.

The NAC model may also be used to predict the stability of emulsions around the µE

phase inversion point. That being said, the process of demulsification is quite complex and

system-dependent, involving a number of different stages, each one with its own governing

equation. For the anionic surfactant (sodium dihexyl sulfosuccinate (SDHS))-toluene-water

emulsions in Chapter 5, each one of these demulsification stages was experimentally

characterized by a time period. In the first aggregation time period (ap), it was proposed that

small emulsion droplets collide with neighboring emulsion droplets via Brownian motion to

form larger flocs. The ap time period is followed by a drainage time period (dp) where all flocs

settle into a more concentrated emulsion layer and undergo film thinning. The individual

emulsion droplets continue to grow via an internal coalescence process over this time period. In

the final coalescence time period (cp), all of the concentrated emulsion droplets gradually begin

to phase separate into a free excess phase. Based on similar phase separation profiles at different

SDHS concentrations, it was also determined that Ostwald ripening only pertains to emulsions

formulated at high surfactant concentrations, away from the phase separation point (HLD<-1

and HLD>1).

169

The simplest of the available demulsification models is that of Davies and Rideal for the

rate of emulsion droplet coalescence (Uc)11,12,13

:

Tk

EaTnkU

B

c

E

EBc exp

3

4 2

(Eq. 7)

The pre-exponential part of this model, which accounts for the average number of

Brownian collisions, is based on von Smoluchowski’s theory of colloid flocculation (truncated

to the formation of doublets) and Einstein’s definition of the diffusion of spheres. The undefined

terms in the pre-exponential part of this model include the number concentration of

monodisperse emulsion droplets (nE), the Boltzmann constant (kB, 1.38×10-23

J/K), and the

absolute temperature (T). The additional exponential activation energy term (Eac) further

describes the probability of an emulsion droplet collision leading to coalescence.

The schematic in Figure 7.1 shows a simplified overview of the collision-coalescence-

separation (CCS) mechanism used to implement Equation 7. According to this mechanism, as 2

emulsion droplets collide and coalesce, they are quickly released into the excess free phase and

the continuous µE phase then drains relatively quickly to help re-establish the initial number

concentration of emulsion droplets (nEo). In this simplified mechanism, both nE (=nEo) and the

average emulsion droplet size (as described by E

dr ) are considered as steady state constants.

This simplification is questionable as experimental observations for SDHS-toluene-water

emulsions in Chapter 5 suggest that although E

dr does not seem to change substantially over the

ap time period, it does increase in size over the latter dp and cp time periods. The CCS

mechanism is nevertheless still able to reproduce the fact that Uc is experimentally observed to

remain constant and that the ap time period is approximately proportional to ( E

dr )3.

170

Collision

Aggregation

CoalescenceSeparated phase

Continuous phase

Figure 7.1: Overview of the simplified CCS demulsification mechanism used to model

emulsion stability.

Finally, the total emulsion stability time period (ep) can be estimated using the following

expression:

c

Eo

pU

ne (Eq. 8)

To model Uc, the Eac term in Equation 7 will be estimated using the hole nucleation

theory for emulsion droplets14,15,16,17

. According to this theory, neighboring emulsion droplets

adjoin, as illustrated in Figure 7.2, by forming a hole across a film of thickness tH and diameter

(radius) dH (rH). Whether this formed hole evolves into a coalescence event depends on the size

of the hole. Small holes tend to contract and close, but large holes continue to grow and lead to

coalescence. Neglecting electrostatic or solvation effects, the hole nucleation theory accounts

for the interfacial rigidity (Er) of the surfactant self-assembly and γow.

171

dH=2rH

tH

rdE

Figure 7.2: Cross-section of coalescing emulsion droplets showing the formation of the

nucleating hole and its characteristic dimensions.

Figure 7.3 presents a summary of the strategies outlined above that will be used in this

chapter to predict the size and stability of Type I and Type II emulsions around the phase

inversion point. This modeling approach will primarily be applied to the SDHS-toluene-water

emulsions characterized at surfactant concentrations of 0.01 M, 0.1 M, and 0.3 M SDHS and

T=280 K, 288 K, 298 K, 308 K, and 317 K in Chapter 5. This same modeling approach will also

be used to predict the size and stability of the emulsions studied by Kabalnov et al., Salager et

al., Binks et al., and Lin et al.16,18,19,20

. The discussion section concentrates on highlighting the

advantages and limitations of the proposed modeling approach.

172

Formulation conditions(surfactant, oil, temperature,

electrolyte concentration, volume fraction)

Mixing conditions

Calculation of HLD

Calculation of SOW properties (HLD-NAC)

Calculation of emulsion initial drop size

Calculation of emulsion stability

Figure 7.3: General algorithm used to predict the size and stability of emulsions.

7.3 DEVELOPMENT OF EMULSION STABILITY MODEL SOLUTION

To model emulsion stability according to Equations 7 and 8, nEo is first solved for by

dividing the volume fraction of the emulsified phase ( E

d ) by the volume of an emulsion droplet

of radius E

dr (calculated using either Equation 5 or 6) as follows:

3

3

4 E

d

E

d

Eo

r

n

(Eq. 9)

The oil to water volume ratio for all of the emulsions modeled in this work is 1 (except

for those of Lin et al.). For this reason, E

d is set equal to 0.5. The sensitivity of Equations 7 and

173

8 to this assumption is tested by considering a range of E

d values up to, and including, the hard

sphere packing limit (=0.74). Increasing E

d any further would result in the transformation of

emulsion droplets into foam-like polyhedrons that no longer collide in a Brownian fashion17

.

For the SDHS-toluene-water emulsions in Chapter 5, εmix is estimated as 14 W/mL based

on a Pcons of 150 W and a VE of 11 mL. A similar value of εmix is assumed for the emulsions of

Kabalnov et al., Salager et al., and Binks et al. as these authors did not provide enough

information to solve for it16,18,19

.

To solve for Eac, the necessary energies of first forming the hole described in Figure 7.2

(E1) and then fully opening it (E2) are considered. Acosta et al. produced an empirical estimate

for E1 of Type III emulsions at HLD=0 (where γow~0)12

:

rc EfE 1 (Eq. 10)

In this equation, fc is a fitting constant. The fitted value of fc ~4 used by Acosta et al. is

employed here. The additional E2 contribution was modeled by de Vries as follows14

:

owHtE 2

2 73.0 (Eq. 11)

Several researchers have proposed different expressions for predicting the critical value

of tH (tH,crit). One such set of expressions derived by de Vries looked to calculate tH,crit according

to the geometrical packing constraints of bubbles21

. Another set of expressions derived by

Klaseboer et al. and Yaminsky et al. looked to calculate tH,crit by balancing the hydrodynamic

pressure at the center of a thinning film with the capillary pressure of fast approaching

bubbles22,23

. However, both sets of expressions do not take into account any of the interaction

energies at play. A more detailed analysis of how tH,crit is influenced by the inclusion of DLVO

174

interactions was issued by Vrij et al. and Chesters for thinning films at low surfactant

concentrations24,25

. Both of these research groups represented tH,crit as the film thickness below

which the attractive component of the DLVO interactions drives the immediate flocculation and

coalescence of emulsion droplets. The emulsions considered in this chapter are however at a

relatively high surfactant concentration where coalescence, and not film thinning, is more rate

limiting. For these emulsions, the thin film remains stable even after tH,crit is reached due to a

disjoining pressure effect. Ivanov et al. and others have demonstrated that the osmotic pressure

contribution to this disjoining pressure effect is most dominant and is closely associated with

expelling µE droplet layers during the film thinning process26,27

. The free energy penalty (ut) of

expelling the last layer of µE droplets, which can be modeled as per the equations of Basheva et

al. in Appendix 2, is the greatest28

. It is therefore proposed that tH,crit can be estimated as

follows:

Lrt E

dcritH 2, (Eq. 12)

In thereby summing E1 and E2 together, the following expression for Eac is obtained:

ow

E

drcc LrEfEa 292.2 (Eq. 13)

An alternative expression for Eac was developed by Kabalnov et al. using the hole

nucleation theory and Helfrich’s model that considers surfactant bending energies15.16,17

. Despite

the several similarities with Kabalnov et al., the simplified Equation 13 uses variables that are

easily predicted by the NAC model and its parameter database.

175

7.4 MATERIALS AND METHODOLOGIES

7.4.1 Materials

The following chemicals were purchased from Sigma-Aldrich Corp. (Oakville, ON,

Canada): NaCl (product #S9625, ≥99.5%) and SDHS (product #86146, ~80% in water).

Furthermore, reagent grade toluene (product #9200-1-30) was obtained from Caledon

Laboratory Chemicals (Georgetown, ON, Canada). Tap water was deionized to a conductivity

<3 µS/cm in the laboratory using an APS Ultra mixed bed resin.

7.4.2 NAC Modeling Methodology

The first step in applying the NAC model was to solve for Hn as a function of the change

in the HLD of ionic (Equation 14) and nonionic (Equation 15) surfactants at salinities S and S*:

L

S

S

L

HLDHn

*

ln

(Eq. 14)

L

SSb

L

HLDHn

* (Eq. 15)

In these equations, S is treated as a scanning parameter and S* is the optimal (or

reference) salinity at HLD=0 where emulsion stability is a minimum. It should be noted that this

alternative approach to solving for Hn (rather than that described in Chapter 6) was used here as

published values of S* for all of the emulsions under consideration were more easily obtained

from the literature instead of all the individual HLD parameters listed in Equations 1 and 2.

Aside from the above calculated Hn, the sphere-equivalent radius of the µE’s Type I (Type II)

continuous ( E

IIIowr

, ) and dispersed ( E

IIIwor

, ) phases, their respective volumes E

IIIowV , and

176

E

IIIwoV , , their respective droplet volume fractions

E

IIIow

, and E

IIIwo

, , and Ha were calculated

using the exact same procedure outlined in Chapter 6. With the aid of E

IIIwor

, as a basis of

calculation, Er was fitted to the measured isotherm γow as follows7,8

:

2,4 E

IIIwo

row

r

E

(Eq. 16)

By also taking E

IIIowV , and

E

IIIwoV , (or

E

IIIow

, and E

IIIwo

, ) together, ρµE was solved

for via the following expression:

E

IIIwo

E

IIIow

E

IIIwo

E

IIIwo

E

IIIow

E

IIIow

EVV

VV

,,

,,,,

(Eq. 17)

In this equation, E

IIIow

, and E

IIIwo

, are the respective densities of the µE’s Type I

(Type II) continuous and dispersed phases. The real length ( E

dl ) and radius ( E

dr ) of the µE

droplets was calculated as per Chapter 6:

E

reald

E

reald

E

reald

E

reald

E

reald

E

reald

nrlr

l

rlH

,,,

,

,, 42

12

2

4 (Eq. 18)

2,,,

,,

43

42

E

reald

E

reald

E

reald

E

reald

E

reald

a

rrl

rlH

(Eq. 19)

The impact of a change in the µE droplet shape from spheres (where E

dl =0) to rods

(where E

dl >> E

dr ) on ηµE was also estimated by assuming that the model for dilute rigid rods is

valid as per Chapter 6:

177

2

2

1E

dE

E

dgE

cE

r

lc

(Eq. 20)

In this equation, E

c

is the viscosity of the µE’s continuous phase and cg is the mass

concentration of µE droplet rods. The required HLD and NAC modeling parameters for

predicting the properties of the different emulsions of interest are listed in Table 7.1.

Table 7.1: Required HLD and NAC parameters for predicting the properties of the emulsions of

Kiran et al. (Chapter 5), Kabalnov et al., Salager et al., Binks et al., and Lin et al.16,18,19,20

.

Parameter Kiran et al. Kabalnov et al. Salager et al. Binks et al. Lin et al.

Surfactant SDHS

Alcohol

Ethoxylate

(C12E5)

Sodium

Dodecyl

Sulfate (SDS)

Aerosol OT

(AOT)

Nonylphenol

Co-surfactant N/A N/A Pentanol N/A N/A

Oil Toluene Octane Kerosene 0.65 cS

PDMS Mineral oil

S (g/100 mL) 1-10 1-20 1-10 0.1-6 0

S* (g/100 mL) 3 9 3.3 0.7 --

T (K) 298 293 298* 298 294

L (Ǻ) 106 25

8 20

7 11

6 18

8

ρoil (g/mL) 0.8729

0.730

0.8731

0.7632

0.84

ρwater (g/mL) f(S)33

ηoil (cP) 0.5629,34,35

0.5130,35

1.3531

0.49 2536

ηwater (cP) 0.8937

Voil (mL) 5.5 1 5.5* 5 20 g

Vwater (mL) 5.5 1 5.5* 5 75 g

csi (M) 0.01, 0.1,

and 0.3 0.08

SDS = 0.009

Pentanol =

0.6

0.04

5 g

ai

(Ǻ2/molecule)

1006 60

8

SDS = 607

Pentanol =

307

100

--

In this table, all of the fields marked by an asterisk “*” are unknown experimental

constants whose values are not available in the literature and that were therefore assumed to be

178

equal to those specified in Chapter 5. Furthermore, ρwater was extrapolated as a function of the

salinity from the referenced website.

7.4.3 Density and Viscosity Measurements

The validity of Equations 17 and 20 were tested for Type I and Type II SDHS-toluene-

water µEs. These µEs were prepared by allowing the mixed components in Table 1 to

equilibrate for 1 week. An estimate for ρµE was obtained by weighing out a 1 mL µE sample

using a Denver Instruments analytical scale (model #TP-214). For ηµE, a Gilmont Instruments

falling ball viscometer (model #GV-2200) was used. A detailed operating procedure for this

technique is provided in Chapter 6.

7.5 RESULTS

7.5.1 Experimental and Predicted SDHS-Toluene-Water Microemulsion

Properties

The experimental and NAC modeled µE properties required for the prediction of the size

and stability of SDHS-toluene-water emulsions are presented in Figures 7.4 (a)-(c). The γow at

0.01 M, 0.1 M, and 0.3 M SDHS can be represented by the single experimental isotherm taken

from Chapter 5 and replotted in Figure 7.4 (a). A fitted Er=1.5 kBT yields an accurate prediction

of γow at all of the tested surfactant concentrations.

The experimental and predicted ρµE in Figure 7.4 (b) are also in good agreement. Here,

Type I and Type II ρµE at a given SDHS concentration are shown to remain approximately

constant. A decrease (increase) in the Type I (Type II) ρµE is however observed with an increase

in the SDHS concentration.

179

In the final plot of ηµE in Figure 7.4 (c), the experimental and predicted values at both

0.01 M and 0.1 M SDHS are very close. At 0.01 M SDHS, ηµE across the Type I and Type II

regions is essentially equal to E

c

. At 0.1 M SDHS, ηµE peaks appear near the Type I-Type III

and Type II-Type III phase behavior limits. This system is an exact reproduction of the previous

modeling efforts in Chapter 6. Unlike at 0.01 M and 0.1 M SDHS, Equation 20 greatly over-

predicts ηµE at 0.3 M SDHS near the Type I-Type III and Type II-Type III phase behavior

limits. The dilute rigid rods model is likely inaccurate at this surfactant concentration and some

level of interaction amongst the rods must lead to the lower ηµE values. One thought by the likes

of Quemada et al. is that the interaction between µE droplets may lead to the early onset of

bicontinuous structures before an actual transition into the Type III region is physically

observed38

. The results in Figure 7.4 (c) suggest that the predicted ηµE should be limited to

values less than 5 cP, consistent with the experimental observations in Chapter 6.

180

0.8

0.9

1

1.1

1.2

-1.5 -1 -0.5 0 0.5 1 1.5

0.01

0.1

1

-1.5 -1 -0.5 0 0.5 1 1.5

0.3

3

30

-1.5 -1 -0.5 0 0.5 1 1.5

HLD

γo

w(m

N/m

)ρμE

(g/m

L)

ημE

(cP

)

(a)

(b)

(c)

Type I Type III Type II

Experimental (Kiran et al.) NAC (0.01 M SDHS) NAC (0.1 M SDHS) NAC (0.3 M SDHS)

Experimental (0.01 M SDHS) Experimental (0.1 M SDHS) Experimental (0.3 M SDHS)

NAC (0.01 M SDHS) NAC (0.1 M SDHS) NAC (0.3 M SDHS)

Figure 7.4: Experimental and NAC modeled (a) γow, (b) ρµE, and (c) ηµE of SDHS-toluene-water

µEs.

7.5.2 Predicted Droplet Size of 0.1 M SDHS-Toluene-Water Emulsions

Figure 7.5 presents the predicted E

dr , calculated using Equation 5, as a function of HLD.

The fitted value of C1 is 0.4 for Type I emulsions (HLD<0) and 0.6 for Type II emulsions

181

(HLD>0). The need for two different fitting constants might be due to the different fluid

properties of the continuous phase (water and toluene respectively). Overall, the experimental

values reported in Chapter 5 are closely reproduced, except for the E

dr of Type I emulsions

close to the Type III transition that are slightly underpredicted. These fitted C values are

employed in all the other emulsion modeling efforts, except for the systems of Lin et al.20

. r d

E(μm)

0.1

1

10

-1.5 -1 -0.5 0 0.5 1 1.5

Type I Type III Type II

Experimental (Kiran et al.)

NAC

HLD

Figure 7.5: Predicted E

dr of 0.1 M SDHS-toluene-water emulsions at 298 K using the NAC

model. Experimental data are obtained from the work presented in Chapter 5.

7.5.3 Predicted Stability of 0.1 M SDHS-Toluene-Water Emulsions

A theoretical demonstration of the prediction of ut from the equations listed in Appendix

2 is presented in Figure 7.6 (a) as a function of the number of µE droplet layers (tH/2 E

dr ) at

HLD=±0.5. These predictions were made by varying the µE droplet volume fraction ( E

d

) from

0.1 to 0.2 and setting the second term in the expression for π1 as P~

. A maximum in ut is clearly

observed in this figure upon expelling the last layer of droplets. The increase in ut at increased

182

E

d

is also consistent with the previous results of Basheva et al.28

. To estimate tH,crit, the NAC

modeled E

dr is incorporated into Equation 12. Figure 7.6 (b) presents the estimated values of

tH,crit as a function of the HLD. As illustrated in this figure, the value of tH,crit increases from 3 to

11 nm upon approaching the Type I-Type III and Type II-Type III phase boundaries. The

reason for this increase in tH,crit is that E

dr also increases as the formulation approaches the

phase boundaries. While close, the predicted tH,crit values are smaller than the 10-20 nm range

measured by Nikolov et al. for the anionic surfactant SDS using a reflected light interferometric

technique39

. Nevertheless, they are still greater than the minimum range for black films (1.3-2.5

nm) reported by Sonneville-Aubrun et al.40

.

-60

-40

-20

0

20

40

60

0 1 2 3 4 5

0

3

6

9

12

15

-1.5 -1 -0.5 0 0.5 1 1.5

t H,c

rit(n

m)

Type I Type III Type II

HLD

ut(k

BT

)

(a) (b)

# µE Droplet Layers

1.0,5.0 E

dHLD

2.0,5.0 E

dHLD

15.0,5.0 E

dHLD

1.0,5.0 E

dHLD

2.0,5.0 E

dHLD

15.0,5.0 E

dHLD

Figure 7.6: Predicted (a) ut and (b) tH,crit of 0.1 M SDHS-toluene-water emulsions at 298 K

using the NAC model.

The predicted Eac are presented in Figure 7.7 as a function of the HLD, along with the

apparent activation energies determined experimentally in Chapter 5. The predicted Eac values

are close to the apparent values at -0.5<HLD<0.5. At the HLD extremes, however, where γow

183

effects are more prominent, the predicted Eac are significantly smaller than the apparent values.

It is possible that the apparent activation energies may contain systematic errors because they

were obtained by assuming that the pre-exponential factor does not vary with changes in

temperature.

1

10

100

-1.5 -1 -0.5 0 0.5 1 1.5

Ea

c(k

BT

)

Type I Type III Type II

Experimental (Kiran et al.)

NAC

HLD

Figure 7.7: Predicted Eac of 0.1 M SDHS-toluene-water emulsions at 298 K using the NAC

model. Experimental data are obtained from the work presented in Chapter 5.

Figure 7.8 presents the estimated stability for 0.1 M SDHS-toluene-water emulsions at

room temperature and as a function of the HLD. Within the range of E

d from 0.5 to 0.74, the

predicted emulsion stabilities compare well with the measured ep time periods in Chapter 5. The

largest deviations occur near the Type I-Type III phase boundary where the stability is under-

predicted by one order of magnitude. However, the predicted values at the Type I-III phase

boundary are consistent with the experimental ap time periods. A smaller over-prediction of the

ep time period occurs at HLD>1.

184

epcpdpap

NAC ( =0.5) NAC ( =0.74)

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Type I Type III Type II

Tim

e (

Se

co

nd

s)

HLD

E

dE

d

Figure 7.8: Predicted stability (solid lines) of 0.1 M SDHS-toluene-water emulsions at room

temperature with E

d =0.5 (initial volume fraction) and 0.74 (closely packed hard spheres). All

experimental data is taken from Chapter 5.

7.5.4 Effect of Surfactant Concentration on the Predicted Stability of SDHS-

Toluene-Water Emulsions

Figure 7.9 presents the experimental and NAC modeled (solid lines) stability of

emulsions prepared with (a) 0.01 M and (b) 0.3 M SDHS. At 0.01 M SDHS, the model over-

predicts, by at least 1 order of magnitude, the stability of emulsions at HLD<-1 and HLD>1.

Conversely, the stability is still underpredicted near the Type I-Type III and Type II-Type III

phase boundaries. At 0.3 M SDHS, relatively accurate stability predictions are obtained for all

the emulsions. It is important to reiterate that the predicted ηµE at 0.3M SDHS is limited to no

more than 5 cP to be consistent with experimental results. Larger predicted ηµE would have

produced local stability peaks near the μE phase boundaries. Although these stability peaks are

185

not experimentally observed for these particular emulsions, they have been observed in the past

for others41

.

epcpdpap NAC

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Tim

e (

Se

co

nd

s)

Type I Type III Type II

(a)

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Tim

e (

Se

co

nd

s)

(b)

HLD

Figure 7.9: Predicted stability (solid lines) of (a) 0.01 M and (b) 0.3 M SDHS-toluene-water

emulsions at 298 K and E

d =0.5. Dashed line in (a) corresponds to predicted values using

tH,crit=0. All experimental data is taken from Chapter 5.

7.5.5 Effect of Temperature on the Predicted Stability of SDHS-Toluene-Water

Emulsions

Figure 7.10 presents the stability of 0.1 M SDHS emulsions at (a) 280 K, (b) 288 K, (c)

308 K, and (d) 317 K. It should be noted here that the effect of temperature on the HLD was

taken into account at the time of calculating the properties of the μEs. Overall, within the range

186

of temperatures considered in Figure 7.10, the stability of these emulsions is reasonably

predicted except at the phase boundaries where it is consistently underpredicted. One also

observes that the stability at HLD=-1.1 and 317 K is overpredicted by a factor of 2.

These observations suggest that the predicted activation energies of Figure 7.7 are, for

the most part, adequate and that the deviations from the experimental values observed at

extreme HLDs might be explained by the method used to estimate those experimental values.

epcpdpap NAC

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

(a) (b)

(c) (d)

Tim

e (

Seco

nd

s)

Tim

e (

Seco

nd

s)

Type I Type III Type II

HLD HLD

Type I Type III Type II

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

1

10

100

1000

10000

100000

-1.5 -1 -0.5 0 0.5 1 1.5

Figure 7.10: Predicted stability (solid lines) of 0.1 M SDHS-toluene-water emulsions at

E

d =0.5 and (a) 280 K, (b) 288 K, (c) 308 K, and (d) 317 K. All experimental data was taken

from Chapter 5.

187

7.6 DISCUSSIONS

7.6.1 Prediction of the Droplet Size and Stability of SDHS-Toluene-Water

Emulsions

The match between the predicted E

dr using the NAC model and the experimental data is

not surprising considering that the NAC modeled µE properties in Figures 7.4 (a)-(c) fitted the

experimental values reasonably well. However, it is important to note that the droplet sizes near

the Type I-Type III boundary are underpredicted. This seems to be connected with the under-

prediction of γow in that region (Figure 7.4 (a)). This deviation in γow may be due to a small

underestimation of the value of S*. The deviation in E

dr near the Type I-Type III phase

boundary might also explain the underprediction of emulsion stabilities at the Type I-Type III

phase boundary for nearly all the SDHS systems. According to the CCS mechanism (Figure 7.1

and Equations 7 and 8), the stability of the emulsion is proportional to the inverse of the initial

drop concentration (epnEo-1

), which in turn means that emulsion stability is proportional to the

cube of the initial emulsion droplet size (ep ( E

dr )3). Therefore an underprediction in E

dr would

produce an even more severe underprediction of emulsion stability.

The surprising element in this work is that the implementation of the simple kinetic

model of Davies together with the simple, and physically inaccurate, CCS mechanism is able to

predict reasonably well (within 1 order of magnitude) the stability of SDHS-toluene-water

emulsions. This might be explained by the fact that the selection of the model and the CCS

mechanism is guided by the features of the separation profiles: constant and nearly identical

coalescence and drainage rates, a relatively long ap time period where almost no changes in the

emulsion are observed, and that the aggregation time scaled to ( E

dr )2.5

. It is unlikely that any

188

other single equation can represent the exact physics of the complex multistep demulsification

process. However, the reason that the kinetic model is successful for the SDHS-toluene-water

emulsions is because it captures the main features of the stages of aggregation and coalescence

that seem to be the rate limiting steps. It is unlikely that this kinetic model and CCS

approximation would work for emulsions where the drainage rate controls the overall

separation.

All the droplets considered in this work are relatively small for creaming or

sedimentation to contribute to droplet collisions. However, for emulsions produced with more

positive or negative HLDs, the larger γow would produce larger droplets that should lead to an

increase in creaming and sedimentation and an acceleration of the overall separation of the

emulsion. Another important clarification is that for the SDHS-toluene-water emulsions

evaluated in this work, Ostwald ripening is not an important factor. Ostwald ripening can

accelerate the overall separation process, particularly for systems with more positive or negative

HLDs that experience larger γow.

One of the consequences of the kinetic model and CCS mechanism is that most of the

changes in emulsion stability around the phase inversion point might be explained by the

changes in E

dr rather than the changes in activation energies as initially proposed by Kabalnov

et al.15,16

. As indicated in Figure 7.7, the changes in Eac around the phase inversion point are

relatively small due to the fact that tH,crit grows as the HLD approaches 0 while γow reduces.

These activation energies produce reasonable predictions of emulsion stability at different

temperatures. The experimentally-determined apparent activation energies used the assumption

of Kabalnov et al. that the pre-exponential factor is constant (i.e. does not change with

temperature) which might have produced an over-estimation of some of the values of activation

189

energies, particularly for large (negative or positive) values of HLD. Another aspect that

deserves some consideration is the estimation of tH,crit as 2( E

dr +L). Overall, this assumption

seems to be consistent as the predicted activation energies help reproduce the experimental

emulsion stabilities at different temperatures. However, it is important to remember that this

estimation of tH,crit is justified by the osmotic pressure associated with removing this last layer of

μE drops. Therefore, surfactant concentration (i.e. the concentration of μE drops) must play a

role on the value of this thickness. In particular, for the case of the 0.01 M SDHS system, the

surfactant concentration is just above the critical μE concentration for this system, therefore the

osmotic pressure contribution is very low for this system and the emulsion drops can approach

each other with negligible resistance. To obtain a better prediction of emulsion stability at 0.01

M SDHS, one could assume that the E2 term can be neglected by setting tH,crit=0. The dashed

line in Figure 7.9 (a) shows the prediction of the emulsion stabilities with no E2 contribution,

which clearly produces a better approximation to the experimental values. Two elements that

were not considered in the estimation of the activation energy were solvation effects and

electrostatic repulsion. Hydration can be important for systems of ethoxylated nonionic

surfactants, particularly at low temperatures, and electrostatic repulsion is likely to be important

for ionic surfactant systems prepared with little or no added electrolyte.

To the best of our knowledge, all previous efforts in studying the changes in emulsion

stability around the phase inversion point have been limited to the qualitative description of

these changes from proposed mechanisms. No quantitative predictions have been produced. The

NAC model used in combination with the kinetic model of Davies and the CCS mechanism has

been able to quantitatively predict the order of magnitude of these stabilities for the first time.

Despite this advancement, there are several modeling limitations that still need to be addressed.

190

The most notable of these limitations is the case where settling becomes more rate-limiting.

Such a scenario is likely to occur when either the gravitational field is << 1 g or ∆ρ approaches

0. Under these types of conditions, the validity of the CCS mechanism breaks down as one can

no longer assume that growing emulsion droplets immediately phase separate. This furthermore

introduces polydispersity effects which cannot be ignored (especially when dealing with

Ostwald ripening). The proposed model conversely breaks down at the other extreme where

either the gravitational field is >> 1 g or ∆ρ is very large. Under these alternative conditions, the

separation of dispersed droplets no longer depends on their Brownian collisions, but instead on

the rate of collisions induced by creaming/sedimentation. Another limitation of the proposed

emulsion stability model is that it ignores solvation and electrostatic repulsion effects. Under

conditions where these effects are prominent, an increase in tH,crit and hence emulsion stability is

expected. One last limitation is that the stability model does not account for rigid structures (i.e.

liquid crystals) at the interface. These types of structures enhance emulsion stability via an

increase in Er.

The idea of using the proposed modeling framework to evaluate the formation and

stability of other emulsions already in the literature is considered in the next sections.

7.6.2 Prediction of the Emulsion Droplet Size of Lin et al.

Lin et al. reported the average emulsion drop diameters produced with 65 wt% water, 5

wt% nonylphenol ethoxylates with different degrees of ethoxylation, and 30 wt% mineral oils.

The emulsions were prepared by mixing 100g of the emulsion at 150 rpm and an impeller with 6

cm diameter. The HLD of nonylphenol ethoxylate - mineral oil systems was evaluated using

Equation 2, with an EACN of 13 for mineral oil, and characteristic curvatures (Cn) for

191

nonylphenol ethoxylates estimated as 7-#ethylene oxide groups, in the absence of salt (S=0) and

at 21°C (ΔT=-4°C)8,42

. The γow of these nonylphenol-mineral oil systems were predicted using

Er=4kBT8. Figure 7.11 presents the drop diameter calculated using Equation 6 and the

properties of the nonylphenol-mineral oil system predicted with the NAC model. One of the

points that Lin et al. made in their contribution was that the optimal emulsification conditions

could be predicted by the determining, experimentally, the maximum solubilization capacity of

their system. Not knowingly, these authors were getting at the point that the optimal

emulsification (lowest drop size) is obtained at HLD=0. The data in Figure 7.11 accurately

predicts (without experiments) that the optimal emulsification occurs in nonylphenols where the

ethylene oxide groups represent 50% of the mass of the surfactant. The minimum drop size of

the emulsion was also predicted. The predicted range of optimal emulsification is relatively

narrow compared to the range of optimal emulsification proposed by Lin et al.

0

25

50

75

100

30 40 50 60 70

Dro

p d

iam

ete

r, u

m

wt% PEO in the surfactant

NAC predicted drop size

Trend proposed by

Lin et al.

Figure 7.11: Predicted emulsion droplet size (solid line) for the nonylphenol ethoxylate-mineral

oil-water systems at 294 K and as a function of the wt% of polyethylene oxide (PEO) in the

192

surfactant. See Table 7.1 for additional simulation conditions. The experimental data is taken

from Lin et al.

7.6.3 NAC Prediction of Initial Emulsion Droplet Size and Stability for the

Systems of Binks et al. with AOT and 0.65 cSt Silicone Oil

The prediction of the HLD and solubilization for the AOT-silicone system has been

reported by Castellino et al.43

. Figure 7.12 (a) presents the predicted γow using Er=1 kBT and its

comparison to the experimental values of Binks et al. The predicted γow match the experimental

values reasonably well. It is important to mention that Binks et al. do not describe the power

input for the mixer. They do, however, present some average drop sizes. The drop sizes

predicted by using the NAC model together with Equation 5 and εmix=14 W/mL range from 0.5

to 5 μm across a HLD range of -0.3 to -2. Binks et al. measured drop size ranges from 1 to 4 μm

across the same range of HLD values. The predicted emulsion stabilities (ep) are presented as

the solid lines in Figure 7.12 (b). In accordance with the ep definitions of Binks et al., the

predicted Type I results are multiplied by a factor of 0.05 and the predicted Type II results are

multiplied by a factor of 0.5. It is observed that the predicted ep is smaller than those obtained

experimentally, particularly at the Type I and Type II extremes. Solvation effects, electrostatic

repulsion, formation of rigid interfaces (e.g. liquid crystals), and/or film drainage issues could

be responsible for these inconsistencies. One crude method to test for solvation/electrostatic

effects is to simply add an additional term to the estimated tH,crit. The estimated tH,crit as per

Equation 12 ranges from 3 to 15 nm as the HLD is reduced from 2 to 0.2. Adding an additional

thickness of 15 nm to all the estimated values of tH,crit produced emulsion stabilities represented

by the dashed lines in Figures 7.12 (b), which represent more closely the experimental trends.

This additional thickness is consistent with measurements of disjoining pressures determined in

193

a liquid surface force apparatus using the system of AOT-water-dodecane (similar EACN to

0.65 cSt silicone oil), in the presence of very low surfactant concentrations and electrolyte

concentrations similar to those used in the Type I μE systems44

. In that work, the authors

determined that the emulsions films could only be compressed until about 10 to 20 nm of

separation. In the presence of higher surfactant concentrations, this separation could grow as

large as 60 nm.

0.001

0.01

0.1

1

10

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2

NACExperimental (Binks et al.)

γo

w(m

N/m

)

Type I Type III Type II

HLD

Tim

e (

Se

co

nd

s)

0.001

0.1

10

1000

100000

10000000

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2

(b)

(a)

Figure 7.12: Predicted (a) interfacial tension and (b) emulsion stability (solid lines) for the

AOT-0.65 cSt silicone oil system. See Table 7.1 for additional simulation conditions. The

dashed line in (b) represents the predicted stability considering an additional contribution to

tH,crit of 15 nm. The experimental data is taken from Binks et al.

194

7.6.4 NAC Prediction of Emulsion Stability for the Systems of Kabalnov et al.

with C12E5 and Octane

The NAC parameters and prediction of the interfacial tension for the C12E5 system has

been presented in a previous work. Unfortunately, in the work of Kabalnov et al., the method of

mixing is simply described as vigorous hand shaking (~5 times) and that other methods of

mixing did not introduce substantial changes. To this end, the “vigorous” hand mixing is simply

simulated as εmix=14W/mL. The work of Kabalnov does not present data on the initial drop size

that could be used to compare/validate the predicted values. Figure 7.13 (a) presents the NAC

predicted and experimental stabilities for C12E5-octane systems. It is important to clarify that

Kabalnov et al. did not follow the data for more than one month and that upward pointing

arrows (as presented by the authors) indicate that the stability of those systems is larger than the

time reported. It is also worth noting that these predicted stabilities are multiplied by a factor of

0.5, consistent with the definition of stability used by Kabalnov et al. Despite all the

extrapolations required to predict the stability of the emulsions of Kabalnov, the stabilities near

the phase transitions are of the same order of magnitude as the experimental values. The rest of

the data, considering that it was not tracked for more than one month, is also consistent with the

predictions of the model.

195

Tim

e (

Se

co

nd

s)

Tim

e (

Se

co

nd

s)

1.E+00

1.E+02

1.E+04

1.E+06

1.E+08

1.E+10

-1.5 -1 -0.5 0 0.5 1 1.5

(a)

Type I

Experimental (Kabalnov et al.)

NAC

HLD

(b)

1

10

100

1000

10000

100000

1000000

-1.5 -1 -0.5 0 0.5 1 1.5

Experimental (Salager et al.)

NAC

Type III Type II

Type I Type III Type II

Er = 2.5 kBT

Er = 1.8 kBT

Er = 1.0 kBT

Figure 7.13: Predicted emulsion stability (solid lines) for the system (a) C12E5–octane and (b)

SDS-pentanol-kerosene. See Table 7.1 for additional simulation conditions. The experimental

data for the systems in (a) and (b) is obtained from Kabalnov et al. and Salager et al.

respectively.

7.6.5 NAC Prediction of Emulsion Stability for the Systems of Salager et al. with

SDS, Pentanol, and Kerosene

For the SDS-pentanol-kerosene system there is no information to validate the prediction

of γow and E

dr . For this reason, Er values of 1 kBT, 1.8 kBT, and 2.5 kBT were used in predicting

the emulsion stabilities as shown in Figure 7.13 (b). The predicted emulsion stabilities (solid

196

lines) for o/w systems (negative HLDs), using Er=2.5 kBT, produces reasonable predictions

whereas the stability of w/o systems is reproduced using Er=1.8 kBT. Although the γow of most

SOW systems can be predicted using a single value of Er, one study reported a case where the

interfacial rigidities of o/w and w/o systems were different45

. It is possible that the partition of

pentanol into the oil-continuous and water-continuous environments is responsible for the

different Er obtained in both scenarios. One feature of the proposed modeling framework that is

illustrated in Figure 7.13 (b) is that even relatively small changes in Er can produce substantial

changes in emulsion stability. This also explains why liquid crystal systems (Er~10 kBT)

produce highly stable emulsions.

7.7 CONCLUSIONS

In this work, the formation and stability of SDHS-toluene-water emulsions around the

phase inversion point was reasonably reproduced using HLD and NAC predicted properties of

the surfactant-oil-water (SOW) system and mixing and demulsification modeling parameters. A

simple collision-coalescence-separation (CCS) mechanism was used to implement the

demulsification model of Davies. This CCS mechanism incorporates the main features of the

drop aggregation and coalescence that were identified as the rate-controlling steps in Chapter 5.

The activation energy of the kinetic model of demulsification was evaluated using the hole

nucleation theory and an assumption that the critical film thickness was defined by the cross

sectional dimension of the μE drops that need to be removed from the film. Overall, the model

was able to predict the size of the emulsion drops and the order of magnitude of the stability of

the SDHS-toluene-water systems around the phase inversion point (HLD from -1.5 to +1.5) for

systems prepared at temperatures ranging from 7 to 44°C and SDHS concentration of 0.01 M to

0.3 M. When this methodology was extrapolated to explain other literature data, the predictions

197

were consistent with the data. The model can be used to produce benchmark estimates that need

to be further discussed in light of the relative importance of other phenomena including

creaming/sedimentation, Ostwald ripening, film drainage, surfactant solvation and electrostatic

interactions, and the potential formation of rigid interfaces, including liquid crystals.

7.8 REFERENCES

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Emulsions. 1. Physicochemical Phenomenology of Demulsified Action. Energy Fuels 2006, 20,

1600-1604.

2 Baran Jr., J.R.; Pope, G.A.; Wade, W.H.; Weerasooriya, V.; Yapa, A. Microemulsion

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3 Wu, B.; Sabatini, D.A. Using Partitioning Alcohol Tracers to Estimate Hydrophobicity of High

Molecular Weight LNAPLs. Environ. Sci. Technol. 2000, 34, 4701-4707.

4 Nardello, V.; Chailloux, N.; Poprawski, J.; Salager, J.-L.; Aubry, J.-M. HLD Concept as a

Tool for the Characterization of Cosmetic Hydrocarbon Oils. Polym. Int. 2003, 52, 602-609.

5 Queste, S.; Salager, J.L.; Strey, R.; Aubry, J.M. The EACN Scale for Oil Classification

Revisited Thanks to Fish Diagrams. J. Colloid Interface Sci. 2007, 312, 98-107.

6 Acosta, E.J.; Yuan, J.S.; Bhakta, A.S. The Characteristic Curvature of Ionic Surfactants. J.

Surfactants Deterg. 2008, 11, 145-158.

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7 Acosta, E.; Szekeres, E.; Sabatini, D.A.; Harwell, J.H. Net-Average Curvature Model for

Solubilization and Supersolubilization in Surfactant Microemulsions. Langmuir 2003, 19, 186-

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8 Acosta, E.J. The HLD-NAC Equation of State for Microemulsions Formulated with Nonionic

Alcohol Ethoxylate and Alkylphenol Ethoxylate Surfactants. Colloids Surf., A 2008, 320, 193-

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9 Walstra, P. Principles of Emulsion Formation. Chem. Eng. Sci. 1993, 48, 333-349.

10 Angle, C.W.; Hamza, H.A. Predicting the Sizes of Toluene-Diluted Heavy Oil Emulsions in

Turbulent Flow. Part 2: Hinze-Kolmogorov Based Model Adapted for Increased Oil Fractions

and Energy Dissipation in a Stirred Tank. Chem. Eng. Sci. 2006, 61, 7325-7335.

11 Davies, J.T. and Rideal, E.K., 1963. Interfacial Phenomena. 2

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12 Acosta, E.J.; Le, M.A.; Harwell, J.H.; Sabatini, D.A. Coalescence and Solubilization Kinetics

in Linker-Modified Microemulsions and Related Systems. Langmuir 2003, 19, 566-574.

13 Rosen, M.J. 2004. Surfactants and Interfacial Phenomena. 3

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16 Kabalnov, A.; Weers, J. Macroemulsion Stability within the Winsor III Region: Theory

versus Experiment. Langmuir 1996, 12, 1931-1935.

17 Babak, V.G.; Stébé, M.-J. Highly Concentrated Emulsions: Physicochemical Principles of

Formulation. J. Dispersion Sci. Technol. 2002, 23, 1-22.

18 Salager, J.L.; Loaiza-Maldonado, I.; Minana-Perez, M.; Silva, F. Surfactant-Oil-Water

Systems near the Affinity Inversion. Part I. Relationship between Equilibrium Phase Behavior

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Silicone Oil+Water+AOT Mixtures. Phys. Chem. Chem. Phys. 1999, 1, 2335-2344.

20 Lin, T.J. Low Surfactant Emulsification. J. Soc. Cosmet. Chem. 1979, 30, 167-180.

21 de Vries, A.J. Morphology, Coalescence, and Size Distributions of Foam Bubbles. In

Adsorptive Bubble Separation Techniques; Lemlich, R., Ed.; Academic Press, Inc.: New York,

1972; pp. 7-32.

22 Klaseboer, E.; Chevaillier, J.P.; Gourdon, C.; Masbernat, O. Film Drainage between Colliding

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23 Yaminsky, V.V.; Ohnishi, S.; Vogler, E.A.; Horn, R.G. Stability of Aqueous Films between

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Thickness. J. Am. Chem. Soc. 1968, 90, 3074-3078.

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Conditions. Colloids Surf., A 1997, 128, 155-175.

27 Danov, K.D. Effect of Surfactants on Drop Stability and Thin Film Drainage. In Fluid

Mechanics of Surfactant and Polymer Solutions; Starov, V.M., Ivanov, I.B., Eds.; Springer:

New York, 2004; pp. 1-38.

28 Basheva, E.S.; Kralchevsky, P.A.; Danov, K.D.; Ananthapadmanabhan, P.; Lips, A. The

Colloid Structural Forces as a Tool for Particle Characterization and Control of Dispersion

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Refractive Index of the Binary Mixtures of Cyclohexane with Hexane, Heptane, Octane,

Nonane, and Decane at (298.15, 303.15, and 308.15) K. J. Chem. Eng. Data 1996, 41, 521-525.

31 Sato, M.; Saito, M.; Hatori, T. Emulsification and Size Control of Insulating and/or Viscous

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32 Polydimethylsiloxane; MSDS No. DMS-T00 [Online]; Gelest, Inc.: Morrisville, PA, March

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Temperature Ranges 25-75°C. Int. J. Thermophys. 1987, 8, 641-647.

35 Harris, K.R.; Malhotra, R.; Woolf, L.A. Temperature and Density Dependence of the

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37 Thompson, J.W.; Kaiser, T.J.; Jorgenson, J.W. Viscosity Measurements of Methanol-Water

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38 Quemada, D.; Langevin, D. A Viscosity Model of Winsor Microemulsions. In Surfactants in

Solution; Mittal, K.L., Ed.; Springer US: New York, 1989, pp. 123-138.

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Surfactant Solutions. I. Experimental. J. Colloid Interface Sci. 1989, 133, 1-12.

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Phase Behavior and Interfacial Tension of Surfactant Systems. J. Colloid Interface Sci. 1979,

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Emulsion Films. Langmuir 1997, 13, 7180-7185.

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Microemulsions. J. Surfactants Deterg. 2012, 15, 495-504.

204

CHAPTER 8:

REVIEW OF THE FORMATION AND STABILITY OF

BITUMEN EMULSIONS

205

The overall goal of this thesis was to study the formation and stability of emulsions both

experimentally and with the aid of the hydrophilic-lipophilic deviation (HLD) and net-average

curvature (NAC) models. The formation and stability of bitumen emulsions (i.e. rag layers) was

of particular interest due to its relevance in the processing of oil sands. In Chapters 2 and 3, the

ability of asphaltenes and naphthenic amphiphiles to stabilize rag layers was experimentally

assessed as a function of solvent-bitumen-water ratios, solvent aromaticity, and temperature. In

Chapter 4, the hydrophilic-lipophilic natures of bitumen as well as that of endogenous surface-

active asphaltenes and naphthenic amphiphiles were quantified. Using toluene as a model oil

phase for bitumen in Chapters 5-7, a new approach for modeling emulsion stability as a function

of the equilibrium phase behavior of related microemulsions (µEs) was developed. The aim of

this review chapter is to revisit the formation and stability of bitumen emulsions and provide

new insight into how they themselves can be predicted.

In considering the formation of bitumen emulsions, the following HLD model is used to

predict the preferential oil-water partitioning behavior of the surfactant (or surfactant mixture):

coc CTAfNSHLD 01.017.0ln , (Eq. 1)

The salinity (S) of the aqueous phase used to formulate these emulsions can be expressed

in terms of an equivalent NaCl concentration of 0.3 g/100 mL. Additionally, the hydrophobicity

of bitumen (Nc,o), expressed in terms of its equivalent alkane carbon number (EACN), is 2.5 and

that of the solvents used to dilute bitumen, heptol 80/20 (80 vol% heptane and 20 vol% toluene)

and heptol 50/50 (50 vol% heptane and 50 vol% toluene), are 5.8 and 4 respectively. For the

case when asphaltenes, with a maximum hydrophobicity (or characteristic curvature (Cc)) of 2.3,

are used as the sole surfactant in the absence of any co-surfactant (f(A)=0) at 25°C (ΔT=0), the

HLD can be expressed as in Figures 8.1 as function of the heptol to bitumen dilution ratio. Here,

206

the range of heptol 80/20 (heptol 50/50) to bitumen dilution ratios from 1/1.5 to 10/1 represents

a corresponding range in the EACN of the mixed oils from 3.6 to 5.4 (3 to 3.8). The fact that a

calculated HLD>0 is observed across all these heptol to bitumen dilution ratios is consistent

with the experimental observance of water-in-oil (w/o) bitumen emulsions in Chapter 2. It

should be noted that the maximum Cc of 2.3 used here is for high molecular weight asphaltenes

(~1000 g/mol). By using a minimum Cc of 0.8 for low molecular weight asphaltenes (~500

g/mol), a range of HLD<0 values is instead observed.

0

0.2

0.4

0.6

0.8

1

0 2 4 6 8 10

Heptol 80/20

Heptol 50/50

Heptol to Bitumen Dilution Ratio (w/w)

HL

D

Figure 8.1: The HLD as a function of the heptol to bitumen dilution ratio for asphaltene-

stabilized rag layers prepared with heptol 80/20 and heptol 50/50 at 25°C.

The effect of 3 wt% sodium naphthenates (NaNs) on the HLD of the above bitumen

emulsions is plotted in Figure 8.2 at an oil to water ratio of 1 to 1 and as a function of the heptol

to bitumen dilution ratio.

207

-5

-4

-3

-2

-1

0

0 2 4 6 8 10

Heptol 80/20

Heptol 50/50

Heptol to Bitumen Dilution Ratio (w/w)

HL

D

Figure 8.2: The effect of 3 wt% sodium naphthenates (NaNs) on the HLD of bitumen emulsions

at an oil to water ratio of 1 to 1 and as a function of the heptol to bitumen dilution ratio.

It is assumed in Figure 8.2 that bitumen is composed of 15% asphaltenes and that all

these asphaltenes are high molecular weight aggregates that are surface-active. The observed

HLD<0 values across the heptol 80/20 and heptol 50/50 to bitumen dilution ratios are in

agreement with the experimentally observed o/w emulsions reported in Chapter 3. While an

increase in the oil to water ratio up until 10 to 1 is observed to drive a more positive HLD as a

result of the increase in the ratio of asphaltenes to NaNs, HLD<0 values are still observed. A 10

wt% NaN concentration would produce ever more negative HLDs.

A similar procedure can be used to evaluate the effect of introducing 3 wt% naphthenic

acids (NAs) on the HLD of bitumen emulsions. The estimated HLD values are presented in

Figure 8.3 as a function of the heptol to bitumen dilution ratio.

208

-5

-4

-3

-2

-1

0

0 2 4 6 8 10

Heptol 80/20

Heptol 50/50

Heptol to Bitumen Dilution Ratio (w/w)

HL

D

Figure 8.3: The effect of 3 wt% naphthenic acids (NAs) on the HLD of bitumen emulsions as a

function of the heptol to bitumen dilution ratio.

The results depicted in Figure 8.3 show that HLD<0 values are predicted at the heptol

80/20 and heptol 50/50 to bitumen dilution ratios considered. These HLD<0 values are to a

certain degree logical in that some signs of o/w emulsions are observed within the bitumen

emulsions evaluated in Chapter 3. For the most part, however, dominant w/o morphologies are

observed. A potential reason for which the formation of o/w bitumen emulsions is predicted

instead of w/o emulsions is because no special attention is given to the effect of NA partitioning

on the actual HLD calculation itself. Also, according to Figure 3.5 (surface pressure isotherm

for mixtures of asphaltenes and NAs) in Chapter 3, even a mixture containing 3 parts (per

weight) of asphaltenes to 1 part of NAs still shows the same surface pressure isotherm of pure

asphaltenes. Considering that bitumen contains 15% asphaltenes, it would take more than 5 wt%

NAs to start replacing some of the asphaltenes from the interface. This would suggest that the 3

209

wt% NAs added to the system has a limited influence on changing the Cc of the surfactant

mixture.

In addition to being able to predict the formation of the above rag layers, the use of the

HLD framework can also be employed to qualitatively interpret their relative stabilities. For

example, asphaltenes alone produce w/o rag layers that are only ~0 to 0.4 units above the

optimum point of minimum emulsion stability (HLD=0). On the other hand, mixtures of

asphaltenes and NaNs produce o/w rag layers that are ~1.5 to 4 units below HLD=0. What this

should signify, at least according to Chapters 5 and 7, is that o/w rag layers stabilized by

mixtures of asphaltenes and NaNs are much more stable than w/o rag layers stabilized by

asphaltenes alone. This hypothesis can indeed be confirmed from the experimental rag layer

stability results presented in Chapters 2 and 3. It can further be concluded from this theory that

mixtures of asphaltenes and NAs should consistently produce rag layers with an HLD~0

(instead of ~-1 to -2 as in Figure 8.3) as they are highly unstable.

The last aspect to be considered in this chapter is how the NAC model can be introduced

to quantitatively predict the stability of o/w bitumen emulsions stabilized by asphaltene and

NaN mixtures. This modeling approach is considered valid for these emulsions as their

measured interfacial tension (γow) values in Chapter 3 are comparable to those for the different

SDHS-based µEs in Chapters 4 and 5. The experimental and predicted γow of these o/w

emulsions as a function of their HLD are presented in Figure 8.4.

210

Figure 8.4: Experimental and predicted γow as a function of the HLD for o/w bitumen emulsions

stabilized by a mixture of asphaltenes and NaNs.

To obtain this figure, it is assumed from the competitive adsorption results presented in

Chapter 3 that NaNs populate the primary adsorbed surfactant layer at the oil-water interface.

Furthermore, the interfacial rigidity (Er) of NaNs is 1 kBT (typical for ionic surfactant systems)

by fitting the extended tail length (L) as 20 Ǻ (assuming that the molecule has the equivalent of

12 carbon groups in length) and taking an area per molecule (asi) of 50 Ǻ2. The predicted

estimates of γow presented in Figure 8.4 are quite reasonable when one considers the complexity

of the system being modeled.

From the predicted γow values in Figure 8.4, predictions of the average o/w bitumen

emulsion droplet diameter ( E

dd ) are obtained as a function of the HLD in Figure 8.5. To

calculate E

dd , Equation 5 of Chapter 7 is used with a known energy dissipation rate ε=14W/mL

and a fitted proportionality constant C1=2. The fact that the value of C1 required to reproduce

211

this data is 4 times larger than that used for toluene emulsions may be due in part to the larger

viscosity ratio between the dispersed and continuous phases. This would also explain why the

emulsion droplet sizes are underpredicted at HLDs closer to 0 (lower heptol to bitumen dilution

ratios) even when using a value of C1=2.

0

3

6

9

12

15

-5 -4 -3 -2 -1 0

Experimental

Predicted

HLD

dd

E(µ

m)

Figure 8.5: Experimental and predicted E

dd as a function of the HLD for o/w bitumen emulsion

droplets stabilized by a mixture of asphaltenes and NaNs.

By taking γow and E

dd together, as well as the calculated activation energies of

coalescence (Eac) from modeled film thicknesses of 7 nm to 9 nm, predictions of o/w bitumen

emulsion stability as a function of the HLD are obtained as in Figure 8.6.

212

1.E+00

1.E+02

1.E+04

1.E+06

1.E+08

1.E+10

-5 -4 -3 -2 -1 0

HLD

Tim

e (S

econ

ds)

Figure 8.6: Predicted stability of o/w bitumen emulsions stabilized by a mixture of asphaltenes

and NaNs as a function of the HLD.

The predicted stability trend depicted in Figure 8.6 closely follows those of the other

emulsions studied in Chapter 7. Unfortunately, however, these predictions cannot be compared

against the experimental rag layer stabilities presented in Chapter 3 as they were subjected to

centrifugal separation. As such, their separation cannot be simply described as per Brownian

collisions in this work. They instead require a careful consideration of hydrodynamic

interactions. It should also be noted here that the predicted emulsion stability times are

minimum values that do not consider the added rigidity imparted by the adsorption of asphaltene

skins.

213

CHAPTER 9:

CONCLUSIONS

214

This thesis explored the stability of bitumen emulsions from an experimental and

modeling perspective. In Chapter 2, it was shown via microscopic and spectroscopic techniques

that the phase separation of water-in-oil (w/o) bitumen emulsion droplets was impacted by the

extent to which asphaltenes adsorbed at the interface. By increasing the water fraction, the

amount of interfacial area produced that asphaltenes could potentially adsorb to also increased.

The likelihood of asphaltenes actually adsorbing depended on their aggregation state. For

instance, monomeric asphaltenes, which were formed by diluting bitumen with an aromatic

solvent, were not very surface-active and therefore tended to remain solubilized within the oil

phase. On the other hand, asphaltene aggregates, which were formed by diluting bitumen with a

more aliphatic solvent, were much more surface-active and therefore tended to readily adsorb at

the interface. While an increase in the formulation temperature did not seem to have an effect on

asphaltene adsorption, phase separation was nevertheless assisted by the enhanced drainage of

the continuous oil phase.

In Chapter 3, the added effect of naphthenic amphiphiles (naphthenic acids (NAs) and

sodium naphthenates (NaNs)) on the stability of the above emulsions was assessed. It was

revealed that under acidic conditions at pH 4, the inclusion of NAs (pre-solubilized in the oil

phase) significantly weakened the stability of governing w/o emulsions. Potential mechanisms

responsible for this behavior included the formation of mixed monolayers and mixed

aggregates. Under more basic conditions (pH 7.5 and 10), NaNs were observed to reverse the

emulsion morphology to oil-in-water (o/w) as a result of displacing asphaltenes from direct

contact with the interface. By also significantly reducing the interfacial tension (γow) of these

systems, the potential for asphaltenes to adsorb as a secondary layer at the interface was

enhanced. This proved to further worsen emulsion stability.

215

Chapter 4 assessed the hydrophilic-lipophilic nature of asphaltenic oils (bitumen,

deasphalted bitumen, asphalt, and naphthalene) as well as surface-active asphaltenes and

naphthenic amphiphiles. The obtained results showed that bitumen and toluene share a similar

hydrophobicity, which explains why they are soluble in one another. It was also demonstrated

that the asphaltene fraction of bitumen was primarily responsible for this similar hydrophobicity

as deasphalted bitumen had a hydrophobicity closer to that of hexane. With regards to the

studied surfactants, asphaltenes and NaNs were shown to be characteristically hydrophobic and

hydrophilic respectively. This supported the formed w/o and o/w emulsion droplet morphologies

of these surfactants in Chapters 2 and 3. Additionally, NAs were shown to have an intermediate

hydrophobicity. This could explain why o/w and w/o emulsion droplets were formed for

asphaltene and NA mixtures.

The stability of anionic surfactant (SDHS)-toluene (used to model the hydrophobicity of

bitumen)-water emulsions, expressed in terms of their aggregation, drainage, and coalescence

time periods, was assessed in Chapter 5 as a function of their proximity to the phase inversion

point. This proximity to the phase inversion point was quantified according to the hydrophilic-

lipophilic deviation (HLD) of related microemulsions (µEs). A notable result from this work

was that emulsion stability tended towards a minimum upon approaching the phase inversion

point (where HLD=0). Furthermore, coalescence appeared to be the rate-limiting

demulsification process for o/w (HLD<0) and w/o (HLD>0) emulsion droplets at a low SDHS

concentration (0.01 M). At larger SDHS concentrations (>0.1 M), aggregation became most

dominant. By also measuring emulsion stability at temperatures ranging from 7°C-44°C, it was

found that the “apparent” activation energies associated with drainage and coalescence were

correlated with γow and tended towards a minimum at HLD=0.

216

In Chapter 6, efforts were made to model the shape of µE droplets and their resulting

viscosity (ηµE) as a function of the HLD using a revised form of the net-average curvature

(NAC) model. To model the shape of µE droplets, it was assumed that they possess a cylindrical

core of length E

realdl, and hemispherical end caps of radius E

realdr, . According to this assumed

morphology, a smooth transition from spheres (where E

realdl, =0) to rods (where E

realdl, >> E

realdr, )

was observed upon approaching HLD=0. These µE droplet shape parameters were in turn used

by Dr. Edgar J. Acosta to predict small angle neutron scattering (SANS) profiles. Knowing how

the shape of µE droplets vary over the HLD spectrum, it was concluded that treating µE droplets

as a dilute mixture of rigid rods allowed for the accurate prediction of experimental ηµE peaks in

the nearby vicinity of HLD=0.

In Chapter 7, the NAC model was used to predict the experimental formation and

stability of the SDHS-toluene-water emulsions studied in Chapter 5 at HLD<0 and HLD>0. In

modeling the formation of these emulsions, their predicted average droplet size under turbulent

mixing conditions was matched to microscopic measurements. To model their stability, it was

assumed that the dispersed droplets collide via Brownian motion. The frequency of these

collisions was a function of the number concentration of emulsion droplets (which was taken as

a constant according to the simplified collision-coalescence-separation assumption) and ηµE.

The probability of a given collision leading to coalescence was controlled by an exponential

activation energy term. This activation energy term was modeled using the hole nucleation

theory. The key parameters of this theory include γow, the interfacial rigidity (Er), and the critical

film thickness (tH,crit) at which the formed hole fully opens rather than closes. While it has

already been demonstrated how to model γow and Er in the literature, a new approach was

introduced here for predicting tH,crit. The obtained modeling results fairly reproduced the

217

experimental emulsion stability trends. The model was also able to reproduce the stability

trends, and in some cases the absolute stability values, of other emulsions already in the

literature. Instances did however occur where modeling adjustments were made to account for

phenomena such as the added rigidity imparted by liquid crystals and surfactant multilayers.

In Chapter 8, the formation and stability of bitumen emulsions was finally revisited

using the HLD and NAC frameworks. The experimental formation of these emulsions with pure

asphaltenes and asphaltene and NaN surfactant mixtures was accurately reproduced using the

HLD model. To be able to accurately reproduce the formation of similar emulsions with

asphaltene and NA surfactant mixtures, partitioning effects need to be considered. The relative

stability of these emulsions was also accurately compared from HLD considerations alone. The

NAC model was able to further predict the γow and the emulsion droplet size for asphaltene and

NaN mixtures. Minimum stability times were also reported that did not take into account the

added effect of secondary asphaltene layers adsorbed on top of a primary adsorbed NaN layer.

218

CHAPTER 10:

RECOMMENDATIONS

219

10.1 RECOMMENDATIONS

Going forward, it is recommended that several studies be carried out to further the

conclusions drawn from this thesis. One such study involves experimentally assessing the effect

of fine clay solids (e.g. kaolinite and illite) on the stability of bitumen emulsions. These fine

clay solids are best described by the likes of Yan et al. and Jiang et al. as having a

heterogeneous surface charge distribution (and hence wettability)1,2

. It is therefore expected that

their molecular interactions with asphaltenes and naphthenic amphiphiles will vary significantly

as a function of solvent-bitumen-water ratios, solvent aromaticity, temperature, and pH.

Another study worthwhile pursuing is how the wettability of solid particles varies

according to the hydrophilic-lipophilic deviation (HLD) model. The wettability of solid particles

in a mixture of different liquids can be described as per the contact angle (θc). At θc<90º, solid

particles are hydrophilic and water-wet. At θc>90º, solid particles are conversely hydrophobic

and oil-wet. It is hypothesized that θc will vary according to HLD modifications as the

molecular interactions between the solid particles and different liquids and between the different

liquids themselves will also be impacted. These changes in molecular interactions can generally

be expressed via interfacial tensions according to Young’s equation. By better understanding

how to fine-tune the stability of solid stabilized emulsions, emulsification and/or demulsification

processes can more accurately be designed.

In next looking at the developed emulsion stability model, there are several parameters

that still require experimental validation. The most significant of these parameters is the

interfacial rigidity of the surfactant self-assembly (Er). A proposed method of indirectly

evaluating Er is by measuring the polydispersity (p2) of related microemulsion (µE) droplets. An

220

approach of this sort was formerly taken by Gradzielski et al. who alternatively related p2 to the

surfactant film’s bending (κ) and saddle-splay ( ) moduli as follows3:

Tfk

Tkp

B

B

228

2

(1)

The other parameters in this equation include the Boltzmann constant (kB), temperature

(T), and f( ), which is an entropic function of the µE droplet volume fraction. Exactly how Er is

related to κ and is not yet understood. Another possible approach for measuring Er is via the

surfactant film elasticity (ε). The reason for such is because both parameters are proportional to

the interfacial tension and available interfacial area.

An additional modeling parameter requiring experimental investigation is the critical

film thickness (tH,crit) at the onset of coalescence. This parameter is currently being

approximated as the thickness of the final layer of expelled µE droplets between approaching

emulsion droplets. It is suggested that a cryo-based imaging technique be applied to frozen

samples to validate the calculated range of tH,crit values from 3-11 nm. An example of such an

imaging technique, with a spot size resolution of 1-20 nm, is scanning electron microscopy

(SEM). SEM has been applied by Binks et al. to the nanoscale study of several emulsion types

in the past4,5

. Other possibilities include time of flight-secondary ion mass spectroscopy (ToF-

SIMS) and x-ray photoelectron spectroscopy (XPS). While the lateral resolution of these latter

imaging techniques is on the order of microns, they are able to profile a sample at a depth

resolution of 1-5 nm6,7

.

The final recommendation is to extend the emulsion stability model. The existing model

is only valid under conditions where emulsion droplet growth takes place over the entire

221

demulsification process. Under alternative conditions where aggregation, settling, film thinning,

and coalescence are less well connected, it is expected that a new set of governing

demulsification equations will apply. How to additionally incorporate complications arising

from multilayer adsorption (i.e. liquid crystals) and/or a change in the separation driving force

(i.e. centrifugation) is of even further interest. Understanding this would permit for a more

accurate prediction of the stability of bitumen emulsions.

10.2 REFERENCES

1 Yan, Z.; Elliott, J.A.W.; Masliyah, J.H. Roles of Various Bitumen Components in the Stability

of Water-in-Diluted Bitumen Emulsions. J. Colloid Interface Sci. 1999, 220, 329-337.

2 Jiang, T.; Hirasaki, G.J.; Miller, C.A.; Ng, S. Effects of Clay Wettability and Process Variables

on Separation of Diluted Bitumen Emulsions. Energy Fuels 2011, 25, 545-554.

3 Gradzielski, M.; Langevin, D.; Farago, B. Experimental Investigation of the Structure of

Nonionic Microemulsions and Their Relation to the Bending Elasticity of the Amphiphilic Film.

Phys. Rev. E: Stat. Phys., Plasmas, Fluids, Relat. Interdiscip. Top. 1996, 53, 3900-3919.

4 Binks, B.P.; Rodrigues, J.A. Enhanced Stabilization of Emulsions Due to Surfactant-Induced

Nanoparticle Flocculation. Langmuir 2007, 23, 7436-7439.

5 Binks, B.P.; Rodrigues, J.A. Double Inversion of Emulsions by Using Nanoparticles and a Di-

Chain Surfactant. Angew. Chem., Int. Ed. 2007, 46, 5389-5392,

222

6 Sodhi, R.N.S. Time-of-Flight Secondary Ion Mass Spectroscopy (ToF-SIMS): Versatility in

Chemical and Imaging Surface Analysis. Analyst 2004, 129, 483-487.

7 Cumpson, P.J. Angle-Resolved XPS and AES: Depth-Resolution Limits and a General

Comparison of Properties of Depth-Profile Reconstruction Methods. J. Electron Spectrosc.

Relat. Phenom. 1995, 73, 25-52.

223

APPENDIX 1:

THE REVISED NET CURVATURE

This chapter is derived from the following published manuscript:

Kiran, S.K.; Acosta, E.J. Predicting the Morphology and Viscosity of Microemulsions using the

HLD-NAC Model. Ind. Eng. Chem. Res. 2010, 49, 3424-3432.

224

The original definition of the net curvature (Hn) was introduced in the NAC model to

reflect the changes in the curvature of the interface with formulation conditions. Its relation to

the actual curvature of the interface was however never established1. In order to re-examine Hn,

it is useful to consider the expression for the radius of curvature (Rc) that Hwan et al. developed

for ionic microemulsions (µEs) based on changes in the electrical double layer surrounding the

head groups of ionic surfactants as a function of the electrolyte concentration2:

1

1

11

1

1

*

*

S

S

k

k

R

tail

c

(Eq. 1)

In this equation, k and k* are the inverse of the Debye length at a given electrolyte

concentration (S) and at the optimal formulation S* (where HLD=0), respectively, and δtail is the

surfactant tail length. In order to compare Hn and Rc, a normalized form of Hn is developed

based on the HLD equation, as a function of S/S*:

S

SHL n*

ln

11 (Eq. 2)

Using values of S/S* ranging from 1×10

-5 to 0.9, the normalized radius of curvature of

Hwan et al. (Rc/δtail) and the normalized net curvature (1/(L×Hn)) were calculated and are

presented in Figure 1.

225

0

5

10

15

20

0 2 4 6 8 10

Rc/δ

tail

1/(L×Hn)

Rc/δtail =2/(L Hn)

Figure 1: Comparison between the normalized radius of curvature predicted by Huang et al.

(Rc/δtail)2 and the normalized net curvature predicted as 1/(L×Hn).

According to these results, there is certainly a close correlation between the 2 normalized

curvatures which supports the idea that, as determined experimentally, the scaling factor L is

proportional to the surfactant tail length δtail. Furthermore, the comparison confirms that the

term “ln(S)” in the HLD equation reflects the effect of the electrolyte on the double layer

thickness. Finally, the solid line in Figure 1 is obtained by using the simple approximation

Rc/δtail=2/(L×Hn). Considering that L~δtail, then a revised net curvature term (H’n) can be

proposed:

c

E

w

E

o

nn

RL

HLD

rr

HH

1

2

11

2

1

2'

(Eq. 3)

According to this equation, H’n provides a better estimation of the curvature (1/Rc) of the

oil-water interface than Hn. However, it is important to keep in mind that the approximation

Rc/δtail=2/(L×Hn) is not accurate at low S/S* values. The value of Rc/δtail calculated using this

226

approximation is 20% larger than the value calculated using equation 1 when S/S*=0.4. As S/S*

tends to 1, this error approaches 0.

1 Acosta, E.; Szekeres, E.; Sabatini, D.A.; Harwell, J.H. Net-Average Curvature Model for

Solubilization and Supersolubilization in Surfactant Microemulsions. Langmuir 2003, 19, 186-

195.

2 Hwan, R.-N.; Miller, C.A.; Fort Jr., T. Determination of Microemulsion Phase Continuity and

Drop Size by Ultracentrifugation. J. Colloid Interface Sci. 1979, 68, 221-234.

227

APPENDIX 2:

DISJOINING PRESSURE EQUATIONS

228

The free energy penalty (ut) of expelling microemulsion (µE) droplet layers during the

film thinning process was modeled by Basheva et al. as follows1:

E

d

H

vdWE

d

H

oscE

d

E

d

E

d

H

tr

tu

r

tu

r

r

r

tu

2222 (Eq. 1)

In this equation, E

dr is the µE droplet radius, E

dr is the emulsion droplet radius, tH is the

emulsion droplet film thickness, and uosc and uvdW are the oscillatory and attractive van der

Waals contributions respectively. Expressions for uosc and uvdW are as follows:

1

221~4

~

21

21

20

E

d

H

oscE

d

H

E

d

H

E

d

H

oscr

tu

r

tP

r

t

r

tu

(Eq. 2)

Ed

H

Ed

Ho

r

t

r

tq

E

d

Ho

oE

d

Ho

o

oo

E

d

Hosc

w

r

t

r

tq

q

w

r

tu

21

12

1122

0 expexp2

sin2

cos12

(Eq. 3)

E

d

H

E

d

B

HHvdW

r

t

r

LTk

Atu

22

212

(Eq. 4)

In these equations, the dimensionless osmotic pressure ( P~

) is equal to

3

32

1

16

E

d

E

d

E

d

E

d

E

d

, the dimensionless excess free energy (~ ) is equal to

3

2

12

19

E

d

E

d

E

d

, w0 is equal to 265315.883439.057909.0 E

d

E

d

, ωo is equal to

3229751.830671.810586.745160.4 E

d

E

d

E

d

, qo is equal to

3259647.3037944.3764378.1978366.4 E

d

E

d

E

d

, 1 is equal to E

d

10336.240095.0 ,

229

w1 is equal to q

ow expcos~2 10 , δ is equal to 1

1

w

, π1 is equal to

oq

o

B

E

d

E

d

Tk

rp

expcos8exp6

20

3~

, the dimensionless chemical potential ( ~ ) is

equal to

32

1

398

E

d

E

d

E

d

E

d

, π0 is equal to 267381.7610572.306281.4 E

d

E

d

, 2 is

equal to 23027.23948.039687.0 E

d

E

d

, E

d

is the µE droplet volume fraction, and AH

is the effective Hamaker constant. The parameter p in the expression for π1 was left undefined

by the authors.

1 Basheva, E.S.; Kralchevsky, P.A.; Danov, K.D.; Ananthapadmanabhan, P.; Lips, A. The

Colloid Structural Forces as a Tool for Particle Characterization and Control of Dispersion

Stability. Phys. Chem. Chem. Phys. 2007, 9, 5183-5198.