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Paste extrusion of polytetrafluoroethylene (PTFE) fine powder resin : the effects of processing aid physical… Ochoa, Isaias 2006

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Paste extrusion of Polytetrafluoroethylene (PTFE)finepowder resin: The effects of processing aid physical properties by  Isaias Ochoa Master of Material and Polymers Science, University o f Sonora, 1997 Bachelor of Chemical Engineer, University o f Sonora, 1991  A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF  DOCTOR OF PHILOSOPHY  in  T H E F A C U L T Y OF G R A D U A T E STUDIES  (Chemical and Biological Engineering)  THE UNIVERSITY OF BRITISH C O L U M B I A A p r i l 2006 © Isaias Ochoa, 2006  Abstract The rheological properties o f a number o f polytetrafluoroethylene (PTFE) pastes were studied relevant to paste extrusion process. The effects o f the physical properties o f lubricants and the geometrical characteristics o f the extrusion die on the preforming, extrusion pressure and mechanical properties o f the final extrudates were studied. A number o f lubricants were identified as suitable for the paste extrusion o f polytetrafluoroethylene (PTFE). They were characterized in terms o f both flow and surface properties. It was found that it is possible to alter the flow and surface properties o f these lubricants independently and thus it became possible to study their relative effects on preforming and extrusion o f P T F E paste. Based on this study, it was concluded that preforming quality increases with increase o f lubricant viscosity and with improvement in the wetting characteristics o f the lubricant with P T F E .  The effects o f the  lubricant physical properties on the processing o f polytetrafluoroethylene (PTFE) fine powder resins were also studied by using dies o f various geometrical characteristics and resins having a variety o f molecular architecture. It was found that the wettability (surface tension) o f the lubricant strongly affects the pressure needed to extrude the P T F E pastes. The viscosity o f the lubricant was also found to play a significant role in the process since a lubricant with a low viscosity causes the paste to be extruded at a lower pressure. These effects  significantly  influence the mechanical properties o f the final extrudates. The latter are functions o f the degree of fibrillation which is significantly influenced by the wettability and viscosity o f lubricants. The effects o f die geometry on extrusion pressure and mechanical properties o f extrudates were assessed in order to determine the geometrical characteristics and operation conditions for the optimization o f the process. A rheological equation was developed and proposed for flow simulations o f the paste extrusion o f P T F E pastes, aiming at optimizing this process. Finally, various additives were tested in order to determine possible processing aids for the paste extrusion o f P T F E . It was found that moderate amount o f boron nitride and organically modified clays can have beneficial effects on mechanical properties o f the final extrudates.  ii  Table of contents Abstract  ii  Table of contents  iii  List of tables  vi  List of  figures  List of symbols  xvii  Acknowledgements 1  2  vii  xx  The Physics and Chemistry of Polytetrafluoroethylene 1.1. Introduction 1.2. Tetrafluoroethylene ( T F E ) polymerization techniques: types o f resins 1.3. Chemical and physical properties o f P T F E 1.4. P T F E fine powder resin processing and applications P T F E Paste Extrusion: General review 2.1. Introduction 2.2. Paste preparation, preforming and extrusion 2.2.1. Paste formulation 2.2.2. Preform ing 2.2.3. Phase migration and extrusion 2.2.4. Sintering 2.3. Constitutive equations proposed to predict pressure drop in capillary die flows 2.4. Basic equations governing the principle o f operation o f the experimental equipment 2.4.1. Capillary flow 2.4.2. Flow in a rectangular channel 2.4.3. Parallel plate flow 2.4.4. Extensional rheometer 2.5. Stress-strain curves  3  Scope of Work 3.1. Introduction 3.2. Thesis objectives 3.3. Thesis organization  4  Experimental Equipment, Materials and Procedures 4.1. Introduction 4.2. Material 4.2.1. PTFE fine powder resins 4.2.2. Lubricants 4.2.3. Other lubricants  1 2 3 6  13 14 15 16 17 18  20 22 22 24 26 29 29 30  iii  32 32 33 34  4.3. Experimental equipment 4.3.1. Preforming and extrusion 4.3.2. Mechanical and viscoelastic properties measurement 4.3.3. Other equipment 4.4. Experimental procedure 4.4.1. Paste preparation 4.4.2. Capillary and sessile drop experiments 4.4.3. Effect of surfactant on lubricant surface tension & wettability 4.4.4. Preform ing 4.4.5. Extrusion 4.4.6. Extrudate analysis 5  6  7  P T F E Paste Preforming 5.1. Introduction 5.2. Densification studies 5.3. Liquid migration 5.3.1. Axial liquid migration 5.3.2. Radial liquid migration 5.3.3. Effect of high preforming pressure and its duration on liquid migration 5.4. Liquid migration during extrusion 5.5. Summary  36 38 39 40 41 49 52 54 54  55 55 58 60 67 69 74  P T F E Paste Extrusion 6.1. Introduction 6.2. Pressure transient in P T F E extrusion 6.3. The effect o f the physical properties o f lubricants on P T F E paste extrusion 6.3.1. The effect of surface tension 6.3.2. The effect of viscosity 6.4. The effect o f geometrical characteristics o f die on extrusion pressure 6.4.1. Reduction ratio 6.4.2. Length-to-diameter ratio (L/D) 6.4.3. Entrance angle 6.4.4. Molecular structure of the resin 6.5. Effect o f temperature on P T F E paste extrusion 6.6. Appearance o f P T F E extrudates 6.7. Mechanical properties o f P T F E extrudates 6.8. Interpretation o f P T F E paste extrusion 6.9. Summary  88 89 91 93 95 98 106 109 111  Rheology of Preformed and Extruded P T F E Pastes 7.1. Introduction 7.2. Rheology o f dry P T F E powders and preformed pastes 7.3. Y i e l d stress o f P T F E pastes 7.4. Extensional rheology o f extrudates obtained from slit die extrusion 7.5. Yielding and deformation behaviour in extrusion 7.6. Modelling the extensional behaviour through the strain-energy function 7.7. Summary ;  112 113 119 134 142 145 153  iv  76 77 82 86  8  Extrusion of P T F E Blends and Effects of Various Additives 8.1. Introduction 8.2. P T F E blend extrusion 8.3. The effect o f lubricant blend on P T F E paste extrusion 8.4. Effects o f additives on paste extrusion 8.5. Summary  9  Conclusions, Recommendations and Contribution to Knowledge 9.1. Conclusions 9.2. Contribution to knowledge 9.3. Recommendations for future work  Bibliography Appendix  v  List of tables Table 4.1  Physical properties o f P T F E fine powder resin studied in this work, as provided by the supplier.  33  Table 4.2  Physical Properties o f Isopar® and HFE-7500 lubricants.  34  Table 4.3  Physical properties o f Isopar® G - A O T solutions at 25°C.  35  Table 4.4  Physical properties o f Isopar® G & Isopar® V solutions at 25°C.  35  Table 4.5  Comparison o f contact angles o f various lubricants with a P T F E substrate obtained by the Young-Dupre equation, the capillary rise and sessile drop methods.  49  Table 4.6  Contact Angle o f Isopar® G - A O T solutions on a P T F E substrate at 25°C.  52  Table 4.7  Physical properties o f Isopar® G - A O T solutions.  52  Table 5.1  Liquid concentration during four extrusion experiments at the conditions listed in Figure 5.20.  70  Table 5.2  Mechanical properties o f the extrudates obtained in different flow zones.  71  Table 6.1  Mechanical properties o f P T F E resins extruded under different conditions.  Table 6.2  Heat o f melting and melting point o f P T F E resins extruded at different  108  conditions.  110  Table 7.1  Y i e l d stress o f P T F E resins.  122  Table 7.2  Y i e l d stress o f various pastes determined by means o f fitting viscoplastic models to the experimental data depicted in Figures 7.23 to 7.26.  131  Table 7.3  The yield stress and strain o f P T F E extrudates at room temperature.  143  Table 7.4  Ogden's parameters for P T F E samples subjected to different Hencky strain rates. Mechanical properties o f PTFE-additive blends + Isopar® G extruded at different conditions.  149  Table 8.1  VI  166  List of figures Figure 1.1  Schematic diagram o f a chain segment o f P T F E molecule  3  Figure 1.2  Partial phase diagram o f P T F E (Sperati, 1989).  5  Figure 1.3  Tube extrusion equipment o f P T F E fine powder (Daikin technical bulletin).  9  Figure 1.4  Electric wire insulation extrusion process o f P T F E fine powder (Daikin technical bulletin).  10  Cross section o f electric wire insulation extruder die (Daikin technical bulletin).  11  Figure 1.5  Figure 1.6  Production o f unsintered tape from P T F E fine powder (Daikin technical bulletin).  '  12  Figure 2.1  Schematic diagram o f capillary rheometer.  21  Figure 2.2  Parallel plate rheometer.  23  Figure 2.3  Schematic o f Sentmanat Extensional Rheometer (SER).  24  Figure 2.4  A typical stress-strain curve o f a material subjected to tension.  26  Figure 2.5  Classification o f material based on the shape o f the stress-strain curves: (a) soft & weak; (b) hard & brittle; (c) hard & strong; (d) soft & tough; and (e) hard & tough.  27  Figure 4.1  S E M image o f F104 H M W fine powder resin.  33  Figure 4.2  Molecular structure o f dioctyl sulfosuccinate sodium salt ( A O T ) .  35  Figure 4.3  Set-up o f the Instron tensile tester machine for paste performing and extrusion.  37  Schematic diagram o f a typical cylindrical capillary die along with the definition o f the design parameters.  37  Schematic diagram o f the tapered slit die showing its contraction and the land zones.  38  Set-up o f C O M - T E N tester to measure the mechanical properties o f extrudates  39  M a x i m u m packing o f the solid phase in P T F E paste.  41  Figure 4.4  Figure 4.5  Figure 4.6  Figure 4.7  vn  Figure 4.8  Variation o f density o f a compacted P T F E resin as a function o f pressure.  42  Figure 4.9  Variation o f porosity, s, of a compacted P T F E resin as a function o f pressure.  43  Capillary rise method to determine the surface energy o f a powder or a liquid.  44  Capillary rise experiment o f Isopar® G through a tube filled with P T F E powder at 25°C.  46  Figure 4.10  Figure 4.11  Figure 4.12  Capillary rise experiment o f H F E 7500 through a tube filled with P T F E powder at 25°C.  46  Figure 4.13  Drops o f liquid placed on P T F E substrate, (a) Water, (b) Isopar® V .  47  Figure 4.14  Capillary rise experiment o f Isopar® G - A O T solutions through a tube filled with P T F E resin at 25°C.  51  (a) Preform paste slicing for determination o f the axial density variation and liquid migration, (b) Preform paste slicing for determination o f radial liquid migration, (c) Top view o f the preform sliced for radial liquid migration.  53  Figure 4.15  Figure 5.1  Figure 5.2  Figure 5.3  Figure 5.4  Figure 5.5  Figure 5.6  Figure 5.7  Figure 5.8  Variation o f preform density in axial direction resulting from an applied pressure o f 1 M P a for 30 s on F104 L M W resin +18 wt% o f lubricant.  56  Variation o f preform density in axial direction resulting from an applied pressure o f 2 M P a for 30 s on F104 L M W resin + 18 wt% o f lubricant.  57  Variation o f preform density in axial direction resulting from an applied pressure o f 3 M P a for 30 s on F104 L M W resin + 18 wt% o f lubricant.  57  L i q u i d migration in axial direction o f Isopar® G , M and V through onesided preformed F303 paste when a 2 M P a pressure was applied for 30 s at 25°C.  59  A x i a l liquid migration for Isopar® G , M and V for a two-sided F303 paste preformed by applying a pressure o f 2 M P a for 30 s at 25°C.  60  A x i a l and radial liquid distribution o f Isopar® G for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  61  A x i a l and radial liquid distribution o f Isopar® G for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  61  A x i a l and radial liquid distribution o f Isopar® M for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  62  Vlll  Figure 5.9  Figure 5.10  Figure 5.11  Figure 5.12  Figure 5.13  Figure 5.14  Figure 5.15  Figure 5.16  Figure 5.17  Figure 5.18  Figure 5.19  Figure 5.20  A x i a l and radial liquid distribution o f Isopar® M for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  62  A x i a l and radial liquid distribution o f Isopar® V for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  63  A x i a l and radial liquid distribution o f Isopar® V for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  63  Contour plot o f axial and radial liquid distribution o f Isopar® G for a onesided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  64  Contour plot o f axial and radial liquid distribution o f Isopar® G for a twosided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  65  Contour plot o f axial and radial liquid distribution o f Isopar® M for a onesided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  65  Contour plot o f axial and radial liquid distribution o f Isopar® M for a twosided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  66  Contour plot o f axial and radial liquid distribution o f Isopar® V for a onesided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  66  Contour plot o f axial and radial liquid distribution o f I s o p a r ® V for a twosided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  67  Effect o f preforming pressure and duration on the axial lubricant migration at 25°C on paste prepared with F301 and 38.8 V % o f lubricant.  68  A x i a l lubricant distribution in a one-sided preformed paste preparedwith F301 and various Isopar® G - A O T solutions preformed at 2 M P a for 30 s at 25°C.  68  Typical start up o f pressure transient during extrusion o f paste(F301 + 18 wt% o f Isopar® M ) using a die having 2a = 60°, L / D = 20 and R R = 352:1 at y  Figure 5.21  Figure 5.22  Figure 5.23  =2812 5 ' a n d 3 5 ° C . _  A  70  Stress vs. strain plot for extrudates obtained during flow zone 1 from Test 2 (see Table 5.2).  72  Stress vs. Strain plot for extrudates obtained during flow zone II from Test 3 (see Table 5.2).  72  Stress vs. strain plot for extrudates obtained during flow zone III from Test 1 (see Table 5.2).  73  ix  Figure 5.24  Stress vs. Strain plot for extrudates obtained during flow zone I V from Test 4 (see Table 5.2).  74  Figure 6.1  Typical start up o f pressure transient obtained i n P T F E paste extrusion.  77  Figure 6.2  Extrusion o f F104 H M W + 38.8 V % o f Isopar® M . The extrusion was stopped and restarted after (a) 1.5 minutes, (b) 10 minutes, (c) 45 minutes, (d) 40 hours.  79  Pressure transients during extrusion o f P T F E paste prepared with F104 H M W and Isopar G at different apparent shear rate values.  80  Pressure transients during extrusion o f P T F E paste prepared with F104 H M W and different lubricants at different apparent shear rate values.  81  Pressure transients during extrusion o f P T F E paste prepared with different resins and Isopar G .  81  Pressure transients during extrusion o f P T F E paste prepared with F104 H M W and Isopar G at different conditions and dies.  82  The effect o f lubricant wettability on the extrusion pressure o f paste prepared with resin F-104 H M W and two different lubricants having about the same viscosity and different wettability properties at 35°C.  83  The effect o f lubricant wettability on the extrusion pressure o f paste prepared with resin F-104 H M W and two different lubricants having about the same viscosity and different wettability properties at 35°C.  84  Tensile strength o f extrudates obtained from paste prepared with resin F104 H M W and two different lubricants having about the same viscosity and different wettability properties with F104 H M W at 35°C.  85  Tensile strength o f extrudates obtained from paste prepared with resin F 301 and two different lubricants having about the same viscosity and different wettability properties with F301 at 35°C.  86  Figure 6.11  Effect o f shear rate on extrusion pressure o f F104 H M W .  87  Figure 6.12  Tensile strength o f the dried extrudates o f pastes prepared with F-104 H M W and different lubricants having different viscosities and about the same wettability properties.  87  Figure 6.13  Effect o f die reduction ratio on extrusion pressure o f F104 H M W resin.  88  Figure 6.14  Effect of die reduction ratio on the tensile strength of dried and sintered extrudates obtained from paste prepared with homopolymer F104 H M W and various lubricants at 35°C.  89  Figure 6.3  Figure 6.4  Figure 6.5  Figure 6.6  Figure 6.7  Figure 6.8  Figure 6.9  Figure 6.10  9  Figure 6.15  Effect o f length-to-diameter ratio ( L / D ) on Extrusion Pressure o f pastes prepared with F301 and various lubricants at 35°C.  90  Effect o f L / D ratio on the tensile strength o f dried (lower plot) and sintered (upper plot) extrudates o f paste prepared with copolymer F301 and different lubricants.  91  Figure 6.17  Effect o f die entrance angle on extrusion pressure o f F301 resin.  92  Figure 6.18  Effect o f die entrance angle on tensile strength o f dried (lower plot) and sintered (upper plot) extrudates obtained from pastes prepared with copolymer F301 and different lubricants at 35°C.  93  Figure 6.19  Effect o f the molecular structure o f the resin on the extrusion pressure.  94  Figure 6.20  Effect o f the molecular structure o f the resin on the tensile strength o f dried  Figure 6.16  (lower plot) and sintered (upper plot) extrudates.  94  Figure 6.21  Effect o f temperature on extrusion pressure.  96  Figure 6.22  Effect o f temperature on tensile strength.  96  Figure 6.23  Effect o f temperature on extrusion pressure. Contour plot o f the results plotted in Figure 6.21.  97  Effect o f temperature on tensile strength. Contour plot o f the results plotted in Figure 6.22.  97  F104 H M W + Isopar® M at 38.8 v o l % . 2a = 15°, L / D = 20, R R = 352, y = 5859 s" . Extrusion Pressure: 59.3 M P a . Tensile Strength: 4.4 M P a .  98  Figure 6.24  Figure 6.25  1  A  Figure 6.26  F104 H M W + Isopar® M at 38.8 vol%. 2a = 30°, L / D = 20 R R = 352, y  Figure 6.27  = 5859 s" . Extrusion Pressure: 73.8 M P a , Tensile Strength: 2.3 M P a . 1  A  100  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = 0, R R = 352, y = 5859 s" . Extrusion Pressure: 73.0 M P a , Tensile Strength: 0.7 M P a .  101  1  1  A  Figure 6.30  99  F104 H M W + Isopar® M at 38.8 v o l % . 2a = 90°, L / D = 20, R R = 352, y = 5859 s' . Extrusion Pressure: 81.0 M P a , Tensile Strength: 3.2 M P a . A  Figure 6.29  99  F104 H M W + Isopar® M at 38.8 v o l % . 2a = 60°, L / D = 20, R R = 352, y  Figure 6.28  = 5859 s" . Extrusion Pressure: 53.8 M P a , Tensile Strength: 3.9 M P a . 1  A  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = 10, R R = 352, y = 5859 s" . Extrusion Pressure: 77.6 M P a , Tensile Strength: 3.2 M P a . 101 1  A  XI  Figure 6.31  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = 40, R R = 352, f = 5859 s" . Extrusion Pressure: 93.0 M P a , Tensile Strength: 2.8 M P a . 102 1  A  Figure 6.32  F104 H M W + Isopar® M at 38.8 v o l % . 2a = 60°, L / D = 20, R R = 352, f = 5859 s" . Extrusion Pressure: 73.8 M P a , Tensile Strength: 2.3 M P a . 102 1  A  Figure 6.33  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 60°, L / D = 20 R R =156, y = 5859 s" . Extrusion Pressure: 24.9 M P a , Tensile Strength: 2.4 M P a . 103 1  A  Figure 6.34  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 60°, L / D = 20, R R = 56, y = 5859 s" . Extrusion Pressure: 9.8 M P a , Tensile Strength: 1.2 M P a . 1  A  Figure 6.35  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 30°, L / D = 20, R R = 352, y  = 5859 s" , T = 15°C. Extrusion Pressure: 39.0 M P a , Tensile Strength: 1  A  2.2 M P a . Figure 6.36  104  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 30°, L / D = 20, R R = 352, f  = 5859 s" , T = 22°C. Extrusion Pressure: 55.9 M P a , Tensile Strength: 1  A  3.6 M P a . Figure 6.37  105  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 30°, L / D = 20, R R = 352, f  = 5859 s" , T = 35°C. Extrusion Pressure: 53.8 M P a , Tensile Strength: 1  A  3.9 M P a . Figure 6.38  103  105  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 30°, L / D = 20, R R = 352, y = 5859 s" , T = 65°C. Extrusion Pressure: 53.2 M P a , Tensile Strength: 7.0 M P a .  106  Figure 7.1  Strain sweep test o f F104 L M W powder at 1.5 H z and 35°C.  113  Figure 7.2  Frequency sweep test o f different resins (1% o f strain and 35°C).  114  Figure 7.3  Preparation o f sample for parallel plate rheological test.  115  Figure 7.4  Strain sweep o f F104 H M W resin (frequency o f 1.5 H z and 35°C).  115  Figure 7.5  Strain sweep test o f F104 H M W preformed (frequency o f 1.5 H z and 35°C).  1  A  Figure 7.6  Figure 7.7  at different  pressures 116  Frequency sweep o f F104 H M W resin preformed at 0.5 M P a and tested at different temperatures.  116  Frequency sweep o f F104 H M W resin preformed at 2 M P a and tested at different temperatures.  117  xii  Figure 7.8  Figure 7.9  Frequency sweep o f F104 H M W resin preformed at 10 M P a and tested at different temperatures.  117  Frequency sweep tests o f different P T F E resins preformed at0.5 M P a (1% strain and 35°C). Frequency sweep tests o f paste prepared with different resins and Isopar® M  118  at 38.8vol% and preformed at 0.5 M P a (1% strain and 35°C).  119  Figure 7.11  Y i e l d stress measurements for F104 H M W resin at 35°C.  120  Figure 7.12  Y i e l d stress measurements for F104 L M W at 35°C.  121  Figure 7.13  Y i e l d stress measurements for F301 at 35°C.  121  Figure 7.14  Y i e l d stress measurements for F303 at 35°C.  122  Figure 7.15  Creep test for F104 H M W fine powder resin at 35°C (cr= 600 Pa, t = 100 s).  123  Figure 7.16  Creep test for F104 L M W fine powder resin at 35°C (cr= 600 Pa, t = 100 s).  124  Figure 7.17  Creep test for F301 fine powder resin at 35°C (a = 600 Pa, t = 100 s).  124  Figure 7.18  Creep test for F303 fine powder resin at 35°C (cr= 600 Pa, t = 100 s).  125  Figure 7.19  Creep test for F104 H M W resin at different levels o f stress.  126  Figure 7.20  Creep test for F104 L M W resin at different levels o f stress.  126  Figure 7.21  Creep test for F301 resin at different levels o f stress.  127  Figure 7.22  Creep test for F303 resin at different levels o f stress.  127  Figure 7.23  Step stress test for F104 H M W fine powder resin preformed at 0.5 M P a .  128  Figure 7.24  Step stress test for F l 04 L M W fine powder resin preformed at 0.5 M P a .  129  Figure 7.25  Step stress test for F301 fine powder resin preformed at 0.5 M P a .  129  Figure 7.26  Step stress test for F303 fine powder resin preformed at 0.5 M P a .  130  Figure 7.27  Fitting viscoplastic models to the rheological data o f F104 H M W fine powder resin.  132  Fitting viscoplastic models to the rheological data o f F104 L M W fine powder resin.  132  Figure 7.10  Figure 7.28  Figure 7.29  Fitting viscoplastic models to the rheological data o f F301 fine powder resin. xiii  133  Figure 7.30  Fitting viscoplastic models to the rheological data o f F303 fine powder 133  resin. Figure 7.31  Figure 7.32  Figure 7.33  The extrusion pressure o f various paste extruded with a slit die as function o f the apparent shear rate at 35°C.  134  Tensile strength o f extrudates obtained from slit die at 35°C as a function o f the apparent shear rate.  135  Engineering ate  H  tensile  stress  response  of  F104  HMW  extrudates  =0.01135"'.  137  Figure 7.34  Engineering tensile stress o f F301 extrudates at £  Figure 7.35  Stress growth coefficient o f F104 H M W extrudates obtained from a slit die with L / H = 20 and different apparent shear rates at 35°C.  138  Stress growth coefficient o f F301 extrudates obtained from a slit die with L / H = 20 and different apparent shear rates at 35°C.  139  Figure 7.36  Figure 7.37  H  = 0.0113-s~'.  137  Comparison o f extensional behaviour o f slit die extrudates obtained at y =3750 s" and subjected to an extension at t = 0.0113 s" . 1  1  A  H  139 Figure 7.38  Stress growth coefficient for F104 H M W at different Hencky strain rates. Samples used were produced by extruding the resin aty  Figure 7.39  = 3750 s~ .  14Q  ]  A  Stress growth coefficient for F301 at different Hencky strain rates. Samples used were produced by extruding the resin at^^ = 3750 s~*.  Figure 7.40  Stress growth coefficient for F104 L M W at different Hencky strain rates. 14^  Samples used were produced by extruding the resin at^^ = 3750 s~ . ]  Figure 7.41  Stress growth coefficient for F303 at different H e n c k y strain rates. Samples used were produced by extruding the resin at y  =3750 s' .  142  x  A  .  143  .  144  Figure 7.42  Engineering tensile stress-strain curves for F104 H M W as a function ofs  Figure 7.43  Engineering tensile stress-strain curves for F104 L M W as a function o f s  Figure 7.44  Engineering tensile stress-strain curves for F301 as a function o£s .  144  Figure 7.45  Engineering tensile stress-strain curves for F303 as a function of s .  145  H  H  H  H  xiv  Figure 7.46  Figure 7.47  Figure 7.48  Figure 7.49  Figure 8.1  Figure 8.2  Figure 8.3  Figure 8.4  Figure 8.5  Figure 8.6  Figure 8.7  Figure 8.8  Figure 8.9  Uniaxial extension o f F104 H M W samples stretched at different Hencky strain rates. Continuous lines represent the model fittings.  151  Uniaxial extension o f F104 L M W samples stretched at different Hencky strain rates. Continuous lines represent the model fittings.  151  Uniaxial extension o f F301 samples stretched at different Hencky strain rates. Continuous lines represent the model fittings.  152  Uniaxial extension o f F303 samples stretched at different Hencky strain rates. Continuous lines represent the model fittings.  152  The extrusion pressure o f a blend o f a homopolymer (F104 H M W ) and a copolymer (F301) extruded at various apparent shear rates at 35°C.  155  The tensile strength o f a blend o f a homopolymer (F104 H M W ) and a copolymer (F301) extruded at various apparent shear rates at 35°C.  156  The extrusion pressure o f a blend o f two copolymers (F301 and F303) extruded at various apparent shear rates at 35°C.  156  The tensile strength o f a blend o f two copolymers (F301 and F303) extruded at various apparent shear rates at 35°C.  157  The density o f mixtures o f Isopar® G and Isopar® V as a function o f composition at 25°C.  158  The viscosity o f mixtures o f Isopar® G and Isopar® V as a function o f composition at 25°C.  158  The surface tension o f mixtures o f Isopar® G and Isopar® V as a function o f composition at 25°C.  159  Steady state extrusion pressure o f a paste prepared with homopolymer F104 H M W and a mixture o f Isopar G & V at 35°C.  159  Tensile strength o f a paste prepared with homopolymer F104 H M W and a mixture o f Isopar G & V at 35°C.  160  Figure 8.10  Extrusion pressure o f F104 H M W + B N blends extruded at 35°C.  161  Figure 8.11  Tensile strength o f F104 H M W + B N blends extruded at 35°C.  161  Figure 8.12  Extrusion pressure o f F104 H M W + Nanomer® I.44P blends extruded at 35°C.  162  Tensile strength o f F104 H M W + Nanomer® I.44P blends extruded at 35°C.  163  Figure 8.13  xv  Figure 8.14  Figure 8.15  Figure 8.16  Figure 8.17  Extrusion pressure o f F104 H M W + Nanomer® P G V blends extruded at 35°C.  163  Extrusion pressure o f F104 H M W + Nanomer® P G V blends extruded at 35°C.  164  The effect o f B N and clay addition on the extrusion pressure o f F104 H M W pastes.  164  The effect o f B N and clay addition on the tensile strength o f F104 H M W pastes.  165  Figure 8.18  Extrusion pressure o f blends o f F104 H M W + Isopar® G and Boron Nitride.  166  Figure 8.19  Tensile strength o f blends o f F104 H M W + Isopar® G and Boron Nitride.  167  xvi  List of symbols Chapter 1 A7 Chapter 2 C D Db Fd L m n P AP Q RR d  v a y f y n  Number average molecular weight. Proportionality constant for the viscous term in mathematical model D i e exit diameter D i e entrance diameter. A l s o refer to the diameter o f the barrel Piston force i n capillary flow Length o f the die capillary land Power law index for the viscous term in mathematical model. Power law index for the elastic term in mathematical model. Driving pressure in capillary flow. Pressure drop in capillary flow. Volumetric flow rate. Reduction ratio defined as the ratio o f the die entrance to exit cross sectional area. Velocity distribution in capillary flow. Die entrance angle, a = 90° refers to a flat die. Critical strain value at yielding. Recovarble strain tensor in mathematical model. Shear rate at the wall in capillary flow.  0  Proportionality constant for the viscous term i n mathematical model. A l s o refers to the viscosity coefficient i n Newton's law for viscoua flow. W a l l shear rate. Y i e l d stress extrapolated at zero velocity.  at c Db D / FR g h k L m nii r r R, R R  Constant that groups the physical properties o f a liquid in capillary rise. Constant for a specific liquid. Constant that groups the physical properties o f liquid i in capillary rise. Constant that account for a packed column. D i e entry diameter also the die o f the barrel. D i e exit diameter. Correction facto for du Notiy ring tensiometer. Force to detach the ring from a liquid surface. Gravity. Height reached by a liquid in a capillary rise. Tortuosity o f the porous i n a packed column. Length o f the die capillary land. Slope o f the Aw vs. t plot. Slope o f the Aw vs. t plot for liquid i. M e a n radius o f the capillary in a packed column. Radius o f the spherical particle. Inner radius o f a tube used for capillary rise. Radius o f the ring. Radius o f the ring's wire.  o a Chapter 4 a w  c  p  R  r  1  2  xvn  RR / V V V\ Vh Vhs V vol% w, Wf wt% W Wt a  Reduction ratio defined as the ratio o f the die entrance to exit cross sectional area. Time. V o l u m e o f the cubic unit in close packing. Volume occupied by the particles in the unit cell. Volume occupied by the lubricant in the unit cell. Volume o f the porous. Volume o f the porous material. Volume o f liquid rise above the flat level o f the liquid surface. Volume percentage o f lubricant i n the paste. Weight o f preform before drying. Weight o f preform after drying. Weight percentage o f lubricant in the paste. Work o f spreading. W o r k o f immersion. Die entrance angle, a = 90° refers to a flat die.  y  Surface tension o f a liquid.  yD  Non-polar component o f a liquid surface tension.  c  p  R  s  Polar component o f a liquid surface tension. Surface energy o f a solid. y t>  Non-polar component o f a solid surface energy.  yND  Polar component o f a solid surface energy.  n 6 Aw pi p p p e Chapter 6 Db D L Q RR  Viscosity o f a liquid. Contact angle. Mass o f a liquid that has penetrated the packed column in a capillary rise. Density o f the lubricant. Density o f the solid material. Density o f a porous material. Density o f air. Porosity o f a porous material.  s  hs  a  T vol% wt% a A//  D i e entrance diameter. A l s o refer to the dimeter o f the barrel. D i e exit diameter. Length o f the die capillary land. Volumetric flow rate. Reduction ratio defined as the ratio o f the die entrance to exit cross sectional area. Temperature. Volume percentage o f lubricant i n the paste. Weight percentage o f lubricant in the paste, Die entrance angle, a = 90° refers to a flat die. Second heat o f recrystallization.  A//  Heat o f melting.  AH AH_  R  C  Heat o f melting o f the reference. Heat o f melting o f the sample.  xviii  y £  £  Apparent shear rate. H e n c k y strain.  h  Hencky strain rate.  h  Chapter 7 Dt, D E G* G' G" J(t) L RR a y y e EH s  D i e entry diameter also the die o f the barrel. Die exit diameter. Y o u n g ' s modulus. Complex shear modulus. Shear storage modulus. Shear loss modulus. Shear creep compliance. Length o f the die capillary land. Reduction ratio defined as the ratio o f the die entrance to exit cross sectional area. D i e entrance angle, a = 90° refers to a flat die. Shear strain. Shear rate. Linear strain. Hencky strain. Hencky strain rate.  n rji rj  Viscosity. Instantaneous viscosity. Tensile stress growth function.  H  +  n* a OE as a <7 0  V  E  Complex viscosity. Shear stress. True tensile stress. Engineering tensile stress. Y i e l d stress i n shear. Y i e l d stress in extensional.  xix  Acknowledgements I wish to express m y sincere gratitude to my supervisor, Dr. Savvas G . Hatzikiriakos, for his truthful guidance and support during the course o f my studies and his patience throughout all these years.  Thanks to D a i k i n Industries L t d for the financial support and the supply o f the polymer samples. Special thanks to Daikin personnel for the comments, suggestions and technical information provided during their visit to our laboratory.  Thanks to m y colleagues and ex-colleagues from RheoLab at U B C for their helpful discussions and exchange o f ideas.  I would like to extend a special note o f thanks to my parents, brothers, sisters, nieces and nephews for their love and continuing support.  Finally, I thank to the National Council o f Science and Technology o f M e x i c o for its financial support that allowed me to continue my education.  xx  CHAPTER 1 The Physics and Chemistry of Polytetrafluoroethylene 1.1 Introduction The era o f fluoropolymers began with the important discovery made by Dr. R o y Plunkett on A p r i l , 1938 while experimenting with tetrafluoroehtylene ( T F E ) in the synthesis o f a new safe, non-flammable, non-toxic, colorless and odourless fluorinated refrigerant (CCIF2-CHF2). Polytetrafluoroethylene  (PTFE) was synthesized for the first time by accident, but once its  physical and chemical properties were disclosed, a wide gamut o f applications was envisioned. Those outstanding properties include a high melting point, exceedingly high molecular weight and melt viscosity, and very limited solubility. Thanks to its unique properties, P T F E has transformed the plastic industry and enhanced the application o f plastic where it was never thought possible. However,' these exceptional properties  also constituted  an obstacle for  processing. The development o f innovative fabrication processes resembling those used with metal powders was a crucial step in the emergence o f P T F E products. Processes currently in use include coating from aqueous dispersions, compression molding and ram extrusion o f granular powders and paste extrusion o f lubricated fine powder (Mazur, 1995). It is the last process o f paste extrusion that is being studied in this work. The rheology o f P T F E paste relevant to the paste extrusion o f P T F E is studied. During the process o f extrusion, the paste starts as a two phase system (lubricant and solid P T F E fine particles) and ends as a solid phase. It is known that many structural changes take place during the flow and thus rheological changes are significant. Therefore, in order to completely understand the process o f paste extrusion, a more fundamental understanding o f the rheological changes that occur during extrusion is needed. In particular, several questions should be considered: what is the influence of the resin structure on the rheological changes taking place during extrusion; what is the influence o f the lubricant concentration on these rheological properties; how is the process influenced by the physical properties o f the lubricant; and what are the effects o f the geometric characteristics o f the extrusion die used on the overall process as w e l l as on the mechanical properties o f the final extrudates. These questions define the scope o f the present work and they are discussed in more detail in chapter 3.  1  This chapter presents a general overview o f the main object o f this work, namely polytetrafluoroethylene  ( P T F E ) . The types o f resins produced, the chemistry o f P T F E and its  various physical properties are discussed. A n overview o f the various applications o f P T F E is also covered. Emphasis is placed on the industrial processes used to synthesize P T F E and in particular on the paste extrusion process.  1.2 Tetrafluoroethylene (TFE) polymerization techniques: types of resins Tetrafluoroethylene (TFE), the monomer for polytetrafluoroethylene (PTFE), is polymerized in a highly exothermic reaction. Two different regimes o f polymerization are common for production o f different types o f P T F E . Both o f them are carried out i n an aqueous medium involving an initiator, a surfactant, other additives and agitation at a high temperature and pressure. The main differences are the amount o f surfactant added to the polymerization reactor and the shear rate applied during the reaction. The first method is known as suspension polymerization.  This is the route to produce  granular resins which are processed as molding powder. In this method, little or no dispersing agent is used along with an initiator in a medium which is subjected to vigorous agitation. This initially causes a stable dispersion o f P T F E for a brief period o f time. The dispersing agent is rapidly consumed causing the polymer to precipitate. A pressure between 0.03 and 3.5 M P a is kept constant to control the molecular weight and its distribution whereas the temperature is held between 40-90°C (Ebnesajjad, 2000). The resulting dried polymer is 'stringy, irregular, and variable in shape' and is known as granular resin. The particle sizes o f granular resin and its powder flow property can be varied, depending on the end product requirements, by size reduction (cutting), or by mixing different grades (Gangal, 1994). The second technique o f polymerization is called emulsion or dispersion polymerization. Using this method, dispersion and fine powder P T F E products are manufactured. M i l d agitation, ample dispersant along with an emulsifying agent and an initiator are involved in this method. The pressure is set around 2.4 M P a , while the temperature is about 95°C. Gentle stirring is usually employed to ensure dispersion stability. The extent o f the dispersion stability is crucial at this stage since the P T F E particles should not coagulate prematurely and yet should be stable enough for transportation, storage and handling. The dispersion recovered from the reactor is finished by two different series o f processes depending on whether a dispersion or a fine powder is the desired final product. In the case o f the former, the diluted dispersion is brought up to 4065% polymer solids by weight in water. In the latter, the agglomerates are isolated by skimming 2  or filtration and then dried by vacuum, radio frequency or heated air such that the wet powder is not excessively fluidized to avoid premature mechanical damage by shearing between particles (Gangal, 1994; Ebnesajjad, 2000).  1.3 Chemical and physical properties of PTFE Fluoropolymer or perfluoropolymer are the names given to designate those polymers whose molecules mainly consist o f carbon (C) and fluorine (F) atoms. These names let us distinguish them from other polymers that are just partially fluorinated. A n example o f a linear fluoropolymer is polytetrafluoroethylene ( P T F E ) whose chemical formula is [(-CF2-CF2-) ]. It n  can be compared with polyethylene [(-CH -CH2-) ] where all the hydrogen atoms have been 2  n  substituted by fluorine atoms. O f course, polyethylene and P T F E are prepared in completely different ways. The basic properties o f fluoropolymers arise from the atomic structure o f fluorine and carbon and their covalent bonding in specific chemical structures. Figure 1.1 depicts the straight chain molecular configuration o f P T F E . The fluorine atoms, in cyan color, are placed helically around the carbon backbone (in grey color) providing a protective shield from virtually any chemical attack thus imparting chemical inertness and stability to the molecule (Gangal, 1994; Ebnesajjad, 2000). The helical conformation o f the fluorine atoms assures that the hysteric repulsion is minimized. The two types o f covalent bonds present in the P T F E molecule, C - F and C - C , are extremely strong (Cottrell, 1958; Sheppard and Sharts, 1969) causing P T F E to have excellent mechanical strength and resistance to heat.  Figure 1.1:  Schematic diagram o f a chain segment o f P T F E molecule.  3  The fluorine shield is also responsible for the low surface energy (18 dynes/cm) causing P T F E to have a low coefficient o f friction (0.05-0.8, static) and non-stick properties (Gangal, 1994). P T F E , with its thermal and chemical stability, makes an excellent electrical insulator. The slippery P T F E can not be dissolved in any solvent, acid, or base and upon melting forms a stiff clear gel without flow. Consequently, its molecular weight cannot be determined by conventional techniques. In practice, the number average molecular weight ( M ) is usually n  estimated from the standard specific gravity (SSG) o f the polymer. Higher S S G implies greater crystallinity and hence, lower molecular weight (Gangal, 1994; DuPont, 2001; Suwa, 1973). Due to the linearity o f P T F E molecules, the crystallinity o f a virgin P T F E resin may be as high as 92-98% (Gangal, 1989). A s a result, the S S G o f P T F E is high for a polymer, typically ranging from 2.1 to 2.3. Following the determination o f S S G using the standard ( A S T M D4895), the number average molecular weight, M ,  can be estimated from  n  M  n  = (0.597 x l O  ,  6  l  f,  procedure  A0.157 1 £-7  \  [1.1]  2.306 - SSG  Equation 1.1 applies to 100% homopolymer resins with S S G > 2.18 (DuPont technical information, 2001). The calculated molecular weights for P T F E with S S G < 2.18 are quite large (probably unrealistic), due to the asymptotic behavior o f Equation 1.1 i n this range. The number average molecular weight o f a 100% homopolymer has also been correlated to the second heat o f recrystallization (AH ). The second heat o f recrystallization is obtained by C  melting and crystallizing a sample o f P T F E twice by Differential Scanning Calorimetry ( D S C ) (DuPont technical information, 2001). It was found that M where AH  C  =2.1xl0 (A# )1 0  n  [1.2]  5 1 6  c  is in cal/g. The applicable cooling rate is 4-32°C/min, over which the heat o f  crystallization remained constant for a given polymer. Typically, M is in the 10 to 10 range 6  7  n  (Gangal, 1994). Comparison o f P T F E molecular weights, regardless o f whether or not the resins contain other comonomers, can be made by considering the resin melt creep viscosity instead. The melt creep viscosity, as detailed in U S Patent 3,819,594 (Holmes et al, 1974), is higher for a higher molecular weight P T F E resin (DuPont, 2001). The melting point o f virgin P T F E (first melting temperature) is 342°C (Sperati, 1989), which is high for a thermoplastic polymer. The second melting temperature is 327°C  4  (Ebnesajjad, 2000), which is the value often reported in the literature. It means that a previously melted P T F E does not recover its original crystallinity; making the resin less crystalline (Gangal, 1994). During melting, a volume increase o f 30% is typical (Sperati, 1989). The melt is stable, since even at 380°C, the melt viscosity is relatively high at approximately 10 GPa.s (Gangal, 1994). Besides the melting point, P T F E has other transition temperatures, two o f which are particularly important due to their proximity to the ambient temperature. These are shown in the partial phase diagram o f P T F E in Figure 1.2 (Sperati, 1989). Under ambient pressure conditions, the first transition occurs at 19°C. A t this temperature, the P T F E molecule chain segments change from a perfect three-dimensional order to a less ordered one undergoing a slight untwisting. Above 37°C, the second transition temperature, the extent o f disorder o f the rotational orientation o f molecules about their long axis is increased. In other words, below 19°C P T F E resin is strong enough to withstand premature mechanical damage. Above 19°C, the molecules are packed more loosely and become highly deformable.  Figure 1.2:  Partial phase diagram o f P T F E (Sperati, 1989).  5  1.4 PTFE fine powder resin processing and applications The following is a brief description o f the most important steps in the P T F E extrusion process indicating the most significant variables involved in each step (Daikin technical bulletin). i. Care and handling of the raw material. P T F E fine powder resins must be in a completely powdered form, so as to enable even pouring when it is blended with an extrusion aid. Strong vibrations and shock should be avoided as much as possible during transport, because this may cause the powder to form lumps. If the powder is to be stored, the ideal storage conditions are a dry place with a temperature range o f 5-20°C. In this way, the powder w i l l be less susceptible to lumping, and even i f it occurs, it is easy to restore the resin to its powdered form. If lumps exist in the powder prior to blending with the extrusion aid, the powder should be sieved using a N o . 4 mesh sieve. Care should be taken not to apply too much force to the powder while sieving. A n y lumps that do not pass through the sieve should be carefully removed, placed in a different container and, once the container has been filled up to 1/3 o f its capacity, shaken to break the lumps apart so that the powder can be sieved again. It is also important not to contaminate the powder during sieving since this may affect the quality o f the final product. i7. Extrusion aid. In the extrusion process for P T F E fine powder, an extrusion aid is used. This acts as a lubricant to enable smooth, even extrusion. The extrusion aid must be able to completely saturate the resin, and must be easily and rapidly removable from the product after extrusion without leaving any residue that would provide odour and/or colour. If the product is to be sintered, the extrusion aid must have a volatizing temperature lower than the sintering temperature thus ensuring that it w i l l not color the product. The extrusion aids ordinarily used are typically clear, light, aliphatic hydrocarbon liquids. Liquids with viscosity o f 0.0005 to 0.005 Pa s are preferred. The amount o f extrusion aid to be added to the resin varies accordingly to the application and the processing conditions. Ordinarily 15-25% in weight o f extrusion aid is used. iii. Extrusion aid blending. The mixing o f the powder and the liquid is crucial in ensuring a uniform flow o f the paste. To prevent shear damaging o f the resin, the blending operation must be performed with the resin temperature remaining below the transition temperature. The blending area should also provide controls for relative humidity, cleanliness, and safety. For blending, a clean, dry, wide-mouth container is filled with the powder (previously sifted) up to 2/3 o f its capacity. The prescribed amount o f processing aid is then  6  added and the container is covered and sealed so the processing aid does not volatilize. The container is then agitated, for example, by placing it horizontally on a two-roller jar mill for approximately 10-20 minutes at 30-45 rpm. iv. Aging. Before blending and prior to preforming, it is suggested that the container be left sealed for approximately 5-15 hours at room temperature (23°C) or above in order to allow the extrusion aid to completely permeate the surface o f any powder not sufficiently permeated by the blending process. v. Preforming. The next step after blending the powder/extrusion aid mixture is preforming. The object o f preforming is to remove most o f the air from the powder after it has been blended with the extrusion aid, and to mould it into a shape which can be inserted into the cylinder o f the extruder. It is beneficial to preform with the resin at the aging temperature. For preforming, the PTFE-lubricant mixture is poured in to the preforming mould and compressed until the volume o f the paste is reduced to approximately 1/3 o f its original volume to produce the preform. The pressure used in preforming should be 1.0-4.9 M P a , and the pressing speed should be 50 mm/min or less. Excessive shearing stress must not be applied, and no air should remain in the material after preforming. The powder is maintained in this pressed condition for 5-10 minutes, and then the pressure is gently released. For easy insertion, the diameter o f the preform must be 0.5-2 m m smaller than the diameter o f the extruder. The specific gravity o f the preform prepared in this way is about 1.3-1.6 g/ml. After preforming, the powder is removed from the preforming mould and inserted into the extruder. Extreme care must be taken at this time, to avoid m i x i n g any foreign matter into the preform. vi. Extrusion. In the extrusion step, the individual P T F E agglomerate particles in the preform are greatly deformed and forced together to create a continuous part such as tubing, rod, or tape. The extruder is capped with a die and the resin and lubricant are compressed and forced out o f the die opening. The function o f the die is to change the shape o f the preform into a moulded product o f the specified shape. The construction o f the die greatly influences the extrusion pressure and hence the quality o f the final product. The die angle varies inversely with the reduction ratio but 20-60° is suitable. The length o f the die land is ordinarily 3-10 times its diameter. It is important for this process to occur above the transition temperature o f the resin, which w i l l be greater than 19°C when under pressure. R o o m temperature is suitable for extrusion, but the die is usually heated to 50-60°C. M o r e than one preform can be loaded into a ram extruder, which is basically a cylindrical press. In this case, prior to extrusion, an extra 7  preforming is carried out within the extrusion cylinder to merge the preforms. From here, the process steps and conditions change accordingly to the desired final product. vii. Drying. In order to remove the extrusion aid contained in the product, the product is placed in an oven and heated to a temperature high enough to evaporate the extrusion aid. Inside the drying oven, the extrusion aid must reach the surface o f the product, and then be evaporated. The speed o f this process depends on. the temperature o f the oven, but it must be appropriately controlled in order to prevent blistering. The temperature o f the drying oven varies accordingly to the thickness and diameter o f the product, the extrusion speed, and the type o f extrusion aid used. Ordinarily, however, the temperatures at the entrance and the exit o f the oven are approximately 100°C and 250°C, respectively. The temperature o f the drying oven is adjusted by controlling the temperature o f the oven's heat source, and by varying the air current inside the oven. Because the extrusion aid is flammable, care must be taken to avoid fire, and sufficient ventilation must be provided. In practice, the drying oven must be 2-3 times longer than the sintering one. If the drying oven is too short, deformities w i l l occur due to inadequate drying. viii. Sintering. Immediately after the extrusion aid has been removed from the product, the temperature is raised to 360-390°C and the product is sintered. A s i n the drying oven, the sintering temperature depends on the extrusion speed and it must be increased i f the extrusion speed increases. However, care must be taken to avoid degradation. The purpose o f sintering is to coalesce the resin and eliminate porosity. A change in the volume o f the product occurs during the latter part o f the drying process and continues through the sintering process, shrinking the volume o f the final product by 25-30%. There is a directional nature to this shrinkage. The most significant shrinking occurs in the direction o f extrusion. This shrinkage is restricted to a large degree by the sintering equipment itself. Typical extrusion processes include tube extrusion, wire coating and calendering. For tube extrusion, the basic equipment is illustrated in Figure 1.3. The extruder consists o f a cylinder, a ram, a driving mechanism (hydraulic or screw type), a die, a mandrel, etc. The cylinders generally used i n extruders range from 50-200 m m i n diameter, and from 500-1800 m m in length. The processing for electric wire insulation (Figures 1.4 and 1.5) is similar to that o f tubes. The main difference is that in the former a clearance must be provided between the guide tube and the guide tip (x in Figure 1.5). If this clearance is too small, the flow section o f the resin is reduced causing the extrusion pressure to rise and the insulation to vary in thickness. If 8  the clearance is too large the flow speed o f the resin w i l l decrease and an excessive shearing force w i l l be exerted on the resin particles in contact with the core wire. This w i l l cause a loss o f resin fluidity, resulting in the cutting o f the core wire or producing defects in the insulation. The resin which is extruded from the die enters the drying and sintering zone together with the core die. Since wire insulation and spaghetti tubes can be curved, the drying and sintering ovens used differ from those used for tubes. A multiple-turn type can be used, which greatly increases the moulding speed.  p j T l . — - Exhaust  oj'°(o;  i I"  •6:  1  in  •Gotitf6l;pane1'  CO  :*3. I:  Counter-  «  Wife pay-off  C°>  Take-up^  Figure 1.3:  1  r  -Wire haul-off  -Sp'ark.tester  Wire  Tube extrusion equipment o f P T F E fine powder (Daikin technical bulletin).  9  Tube extrusion die Exhaust  Cylinder Ram head Seal ring*-  """-Dfe:.Temp. (5Q~6CC) (-122-140°F) KT-Banci heater (50-~60 'C) (122-140'F) "'Oie orifice  (Oh 30 - 60")  r  Core  R,R.  2  do'-dp*  Dc: Cylinder inside diameter Dm: Mandrel outside diameter do: Die orifice inside diameter dp: Core pin outside diameter  Molded product  Figure 1.4:  . PC*-Dm  Electric wire insulation extrusion process o f P T F E fine powder (Daikin technical bulletin).  P T F E fine powder can also be processed into unsintered tape by calendering. The process for ordinary unsintered tape is illustrated in Figure 1.6. The extrusion o f unsintered rod (cylindrical or square shape) is done in the same way as in tube extrusion. However, because calendering is done in an unsintered state, the lubricating properties o f the extrusion aid become more important than anything else. Generally, an extrusion cylinder 100-250 m m in diameter and a round die o f 10-30 m m in diameter (or a rectangular die 15x6 m m or 20x10 mm) are used. For a single step process, a roll o f 300-500 mm is used. The roll has a polished surface with no eccentricity and it is heated to 50-80°C and spins at 5-30 m/min. After these steps, the processing aid contained in the tape is removed by drying it in the calender roller and in a hot air furnace. Tape made in this way is called unsintered tape. It is generally a slit o f 13 m m in width, and wound in fixed lengths onto reels (see Figure 1.6 for details). 10  Figure 1.5:  Cross section o f electric wire insulation extruder die (Daikin technical bulletin).  In this work the paste extrusion o f P T F E through circular and slit dies having various geometrical characteristics is studied. In particular the rheological properties relevant to the processing o f paste are identified and an appropriate constitutive equation is formulated. This can be used to model several o f the processes described above. The effect o f the physical properties o f the lubricant such as viscosity and surface tension on the rheology and extrusion pressure is also studied systematically. A n essential part o f this study is the effect o f the lubricant type and resin as well as the die geometry on the mechanical properties o f the final  11  extrudates. The tensile strength o f the unsintered and sintered dried extrudates is taken as a measure o f their mechanical properties.  Exhaust, or recovery  Take-up roll  Take-up. roll  .Extrusion aid extraction:and:drawing Heating rolh.''00 ?5cc:  Figure 1.6:  Production o f unsintered tape from P T F E fine powder (Daikin technical bulletin).  12  CHAPTER 2 PTFE Paste Extrusion: General Review 2.1 Introduction Paste extrusion is a widely used process in many different industries, including the chemical, food and pharmaceutical industries. Less common, but not less important products include ceramic components, catalyst supports, bricks, and many others. The increasing demand of these products has attracted the attention o f researchers around the world and greatly increased interest in studying the paste extrusion process. M a n y complicated structures, such as thin-walled honeycomb catalytic supports, rely on the uniformity o f the extrudate to provide certain "high performance" properties. Ram extrusion has made possible the characterization o f the rheological properties o f pastes when other techniques can not be used. A m o n g the most important factors to be considered in ram extrusion o f pastes are paste formulation, paste densification, extrusion rate and die geometry. A l l these factors together w i l l allow a complete understanding o f paste extrusion for determining the optimum processing conditions for a given extruded product. So far, most o f the work on paste extrusion has been done with alumina pastes due to their importance in the catalyst and electronics industries. However, P T F E prepared by emulsion polymerization uses paste extrusion as a good alternative for product manufacturing. Thanks to paste extrusion it is possible to process thin hoses, thick tubes (liners), wire insulation and unsintered tapes o f P T F E but, like for other materials, P T F E paste extrusion is still under study and development. In this chapter, literature related to the subject o f P T F E rheology and its relation to the process o f paste extrusion is reviewed. The discussion is subdivided into paste flow and extrusion, and the modelling o f paste flow. In addition, some definitions used in other chapters are introduced and the fundamentals o f the operation o f the equipment used to study the rheology o f P T F E paste extrudates are also included.  13  2.2 Paste preparation, preforming and extrusion 2.2.1 Paste  formulation  In simple terms, paste is a mixture o f solid and liquid, the relative amounts being such that the resulting material can be moulded readily (Benbow and Bridgwater, 1993). In other words, the composition o f the paste is that to render a material soft and plastic, but the object so formed should be able to retain its shape to allow further processing. However, this definition is not definitive and other definitions, based on the perceived mechanical response, have been put forward (Khan, 2001). Fine powders o f polytetrafluoroethylene (PTFE) are also processed by paste extrusion. Here, P T F E resin is combined with a minimal quantity o f lubricant (an inert liquid hydrocarbon) and then extruded at modest temperature (typically 30-35°C) into preforms o f various shapes and dimensions with substantial mechanical integrity (Mazur, 1995). In paste extrusion, the liquid used as a processing aid has two purposes: to protect the solid particles from mechanical damage and to act as a lubricant. Thus, the particles that in the absence o f lubricant were susceptible to shear damage, are now covered with a thin layer o f lubricant that makes the paste become resistant to compressive load without increasing the interparticle contact area. A s a lubricant, the liquid separates the particles and lubricates both the relative particle movement and the motion along the wall o f the die land or other surfaces in the flow path during extrusion (Benbow and Bridgwater, 1993). Even though the liquid imparts plasticity to the solid phase; it does not act like plasticizer and is removed from the extrudate right after extrusion (Mazur, 1995). The amount o f lubricant and its properties critically affect the extrusion process and, hence, the quality o f the final product. The concentration o f the processing aid in the mixture depends on the type o f the product, equipment design and the desired extrusion pressure. Its content should be as low as possible but not so low that the extrusion pressure would be excessively high. The optimum range o f lubricant content was found to be between 15 and 25% of the total weight o f the compound, which typically corresponds to a volume fraction between 0.34 and 0.45 (Mazur, 1995; Ebnesajjad 2000). A s the amount o f liquid added to the powder increases above a critical value, the pressure required to extrude the mixture falls dramatically (Benbow, 1998; Ariawan, 2002). For a typical commercial fine powder, a 2% increase in lubricant causes a 40% decrease in extrusion pressure (Daikin technical bulletin). A s more 14  liquid is added, the material soon becomes too soft to retain its shape. O n the other hand, i f an inadequate amount o f lubricant is used, the extrudate tends to be rough and irregular (Mazur, 1995). Regarding the properties o f the processing aid, any difference in density and/or viscosity implies different rheological properties. The viscosity o f the lubricant has a large effect on the quality o f the paste. For example, the use o f a more viscous liquid as a lubricant results in a less uniform mixture (Ochoa and Hatzikiriakos, 2004). Consequently the paste would not extrude as a continuous body, and many microcracks are developed during the drying process  after  extrusion (Ebnesajjad, 2000). In addition, the extrusion pressure exhibits higher values when a processing aid with a higher viscosity is used (Benbow, 1998; Ochoa and Hatzikiriakos, 2005). Ideally, the lubricant should have a lower surface tension than the critical surface tension o f P T F E . The critical surface tension is the value o f surface tension o f a liquid below which the liquid w i l l spread on a solid. That increases the wettability o f the lubricant with the resin particles (Ebnesajjad, 2000; Ochoa and Hatzikiriakos, 2005). The extrusion aid must be easily removable from the extrudate without leaving a residue, which could alter the colour o f the final product. Other requirements o f lubricants include high purity, low odour, low polar components, high auto-ignition temperature, and low skin irritation. 2.2.2  Preforming Another aspect related to paste extrusion is preforming. During this step, the paste is  placed in a cylindrical billet and, by means o f a piston, the pressure is gradually increased to remove the air voids that would otherwise render a final product mechanically weak. In this way a cylindrical rod that is fitted into the extruder's barrel is formed. In P T F E paste processing, the preforming stage is carried out at room temperature although it is not temperature sensitive (Mazur, 1995). However, the application o f stress introduces another problem since it may cause the liquid component o f a paste to move through the solid matrix in the radial and axial directions causing a liquid maldistribution throughout the paste ( Y u , 1998). The level o f the preforming pressure and its duration significantly affect the quality o f the preform. In addition, the magnitude o f the pressure needed to produce a preform o f uniform density depends on the molecular weight (standard specific gravity) o f the resin (Ariawan, 2001). A lack o f adequate pressure w i l l result in a preform o f non-uniform density which w i l l extrude unsteadily, resulting in an unacceptable final product. During preforming, the applied pressure compacts the particles making those adjacent to the wall o f the preforming unit undergo plastic deformation that results 15  in a smooth film o f deformed powder surrounding the preform. Thanks to this layer, the rest o f the resin particles remain spherical even after high preforming pressure (Mazur, 1995). 2.2.3 Phase migration and  extrusion  Phase migration is a phenomenon that occurs not only during preforming but also during extrusion (Yu, 1998). It is caused by relative motion o f the liquid through the voids between the solid-phase particles. This migration eventually results in a non-uniform distribution o f lubricant in the mixture. This effect is enhanced with time, especially i n the presence o f high extrusion pressures. A s the paste becomes drier, the extrusion pressure rises and the liquid loss increases consequentially. Eventually, high frictional forces may occur due to direct contact between particles, and between the particles and containing walls (Benbow et al., 1998; Blackburn, 1993). The packing characteristics o f the paste depend on the particle size, shape and size distribution and there is a direct correlation between the permeability o f a consolidated paste and its porosity (Rough, 2002). Thus, since the solid and liquid phases move at significantly different rates under the application o f a pressure gradient, part o f the liquid escapes from the paste. If the permeability through the packed particles is high and the liquid viscosity is low, conditions for the liquid to move faster than the solid w i l l be promoted. The result w i l l be that the paste becomes effectively drier, the extrusion pressure rises, and the extrusion process may have to be halted in extreme case (Benbow and Bridgwater, 1993). After extrusion, the extrudates can exhibit surface fracture depending on the processing conditions. Benbow and Bridgwater (1993) have reported the effect o f die shape, operating conditions, and paste formulation on the surface defects i n the final products. Domanti and Bridgwater (2000) studied extensively the effect o f die land length, extrusion rate, die entry angle, extrusion ratio and water content on the surface fracture in the extrusion o f a-alumina paste mixed with Bentonite clay and carbohydrates. To reduce the severity o f the extrudate distortion several options are available such as decreasing the extrusion rate, increasing the lubricant concentration i n the paste mixture, using an extrusion dies with long length to diameter ratio and a small entrance angle (Benbow et al., 1987; Benbow and Bridgwater, 1993), altering the viscosity and yield properties o f the liquid phase, and blending fine and coarse powders in order to decrease the average pore size (Blackburn and Bohm, 1993). These alternatives have advantages and disadvantages,  or may not work for some pastes. A s far as P T F E paste  processing is concerned, there is an optimum value for the entrance angle and length to diameter 16  ratio o f the die as well as lubricant and viscosity concentration (Ariawan, 2001; Ochoa and Hatizikiriakos, 2005). A very interesting phenomenon that occurs during P T F E paste extrusion is fibrillation. It is the formation o f fibrils that interconnect the particles, and these essentially give the dimensional stability to the final product. Lewis and Winchester (1953) first reported that fibrillation occurs during paste flow through the contraction area o f the die. Later, Ariawan (2001) found the same through S E M analysis o f paste in the die entry region. Mazur (1995) explained this phenomenon by making reference to particles that reorganize themselves to pass through the die during the initial stage of extrusion. After passing that region, the particles are deformed due the shear/extensional  stresses, resulting in the formation o f fibrils which  contribute to the mechanical strength o f the extrudates. 2.2.4 Sintering Sintering is the process during which a granular material, such as a polymer powder o f ultra-high molecular weight polyethylene ( U H M W P E ) or P T F E , is heated to a temperature near its melting point (Hooper, 2000). In this process the particles o f the loose powder or pressed compact material weld together to form an interconnected solid (Mackenzie, 1949). A s a result, the density o f the compact changes. The coalescence o f contacting polymer particles is important to provide the final product with suitable and improved mechanical properties. Previous studies o f sintering revealed that the surface tension was the driving force for this phenomenon to occur. However, more recent reports have shown that the degree o f sintering is governed by the particle size, viscosity, interfacial tension, molecular architecture and molecular weight distribution (Hooper, 2000). In P T F E paste processing, the elastic phenomena dominate the sintering process (Mazur, 1995). During P T F E sintering, the net volume o f the material changes but these changes in linear dimensions is highly anisotropic. During the heating cycle, the sample contracts in the axial direction and expands in the radial direction yielding a net shrinkage of about 4% (Mazur, 1995). In fact, axial contraction is the resultant o f a contraction and expansion occurring simultaneously. Apparently, the axial contraction is driven by molecular orientation while the expansion is the response to release the stress accumulating during the former (Mazur, 1995). The resultant sintered extrudate exhibits a higher tensile strength than the unsintered sample, indicating improvement in its mechanical properties.  17  2.3 Constitutive equations proposed to predict pressure drop in capillary die flows Assuming steady state flow through an orifice die (length to diameter ratio, L/D = 0) o f conical entry angle 2a, and using the following rheological model: T = Cy" +T]  [2.1]  dt  Snelling and Lontz (1960) derived the following relationship for the total pressure drop during extrusion: 12Qsin a 3  AP=-^-[3i„ ™,r ^ (  +  3m n(\-  3(« + l)  [2.2]  cos a)D  3  where C, n, TJ, and m are constants to be evaluated experimentally, R R is the reduction ratio defined as (D//D) , and Q is the volumetric flow rate. T o arrive at Equation 2.2 through Equation 2.1, the authors used the radial flow hypothesis. This hypothesis assumes that paste particles at the same radial distance from the virtual apex o f the conical zone o f the die, move towards the die apex at the same velocity. This has been proven experimentally (Sneeling and Lontz, 1960; Ariawan et al., 2002b). Thus, the velocity o f a point on a spherical surface at distance r from the apex is: - = dt Doraiswamy  et al.  (1991)  have  ^ 2tt(1 - cos a)r  developed  [2.3]  a non-linear rheological model  for  concentrated pastes. This model considers the elastic, viscous, and yielding behaviour o f the material by introducing a recoverable strain term, y . In addition, this model has the advantage o f using data easily accessible by means o f a parallel plate rheometer. Thus, they suggested the following constitutive equation for a material that exhibits yield stress: T - Gy T=  \?\<Yc  [2.4]  \y\=y  [2.5]  =Y  \?\<Yc  t - ]  = 0  191= 7c  (Gy  • + K\yr  y  ]  c  J dy ~dt dy ~dt  18  2  <r  6  [ - l 2  7  where G is the elastic modulus, y is the critical strain value at yielding, y is the recoverable c  strain tensor, and K and n are power law constants. Note that the viscosity, evaluated as the term in brackets, approaches a Newtonian viscosity at low shear rates and a power law viscosity at high shear rates. Based on the above model, the pressure drop for an orifice die can be found to be: f  4Q~„ .  ~n + \ n  V  +  7iD _  Gy  K  K  D  c  "i  2  J  [2.8]  A simple equation for paste flow through dies o f various entry angles, cross-sectional shapes and L / D ratios have been derived analytically by Benbow and Bridgwater (1993). Particularly, for a steady state flow through a capillary die o f entry angle 2a, and performing separate force balances on the entry region o f the die and on the actual die land, the following relationship for the total pressure drop was derived: AP = \2(a  0  +ZV"  +T cot O  a  '  i-  I D  —T  f  where a , £ m, r , /?/ and n are parameters to be determined experimentally, VsQ/nD  2  0  0  and Db is  the barrel diameter. The first term accounts for the change i n cross sectional area in the conical entry {extensional and shear term), while the second term for the pressure drop in the die land {shear term). The rheological model used to obtain Equation 2.9 was cr = a +a-V  [2.10]  0  where <J is the yield stress extrapolated to zero velocity and V is the paste velocity in the die 0  land and a is a factor characterizing the effect o f velocity. The term a- V is, thus, analogous to the product r/ • y in liquid shear flow. Using the radial flow hypothesis proposed by Snelling and Lontz (1960), Ariawan et al. (2002b) developed a one-dimensional mathematical model to describe the effect o f extrusion conditions and geometric characteristic o f the die in the extrusion pressure o f P T F E paste. This model considers the paste as an elasto-viscoplastic material that exhibits both strain hardening and viscous resistance effects during flow. Thus, by using a power-law modified Kelvin's stress-strain relationship, Ariawan (2002) suggested the following constitutive equation:  ^-^  r  = CrL+w:  m  19  [2.ii]  where a  e  and a are the principal stresses in the 0 and r directions respectively and y r  nm  and  v are the maximum values o f the strain and strain rate, respectively. The extrusion pressure in / max the conical zone was found to be: P  .  =<j.  extrusion  =  C  r  rb  RR  B  ra  +2(1 + \  f B)\ c /  12Qsin a (3m + 2 5 ) ^ ( 1 - cos a)Db 3  V  [2.12]  3  J  where a „ is the stress at the die exit, RR is the reduction ratio defined as (Dt/D) , and C, 77, n, m 2  r  and/are material constants that have to be determined experimentally. This model was  found  to describe  adequately  most  experimental  observations.  Specifically, it was found to predict quantitatively and qualitatively the effects o f die entrance angle, reduction ratio, pressure, L / D ratio o f the die, and P T F E properties. However, it preassumes the velocity distribution by using the radial flow hypothesis.  2.4 Basic equations governing the principle of operation of the experimental equipment 2.4.1 Capillary flow The simplest and most popular type o f rheometer is the capillary rheometer shown in Figure 2.1 (Dealy, 1990). In its simplest configuration, the capillary rheometer consists o f a small tube through which paste is made to flow, either by means o f an imposed pressure or a piston moving at a fixed speed. The quantities normally measured are the flow rate Q, and the driving pressure, P^. If the flow is generated by a moving piston, it is usually the piston force, Fd, that is measured, and this is related to P as follows: d  P  d  4F = - ^  [2.13]  where Dt, is the diameter o f the barrel or reservoir (Dealy and Wissbrun, 1990). It can be shown that the absolute value o f the shear stress at the wall, a , is related to the w  pressure drop, AP, over a length o f tube, L, as follows: •  20  2L  p.14]  The pressure drop, AP, is always a negative quantity, because the flow is in the direction o f the axial coordinate, z. A s this is a partially controllable flow, the velocity profile depends on the rheological properties o f the fluid under study, and a general expression relating the volumetric flow rate, Q, to the wall shear rate cannot be derived. It is known that for a Newtonian fluid, the velocity distribution is given by the familiar parabolic law: Q v = 2TlR'  [2.15]  1-  This is the velocity profile for "fully developed flow" in which the effects of the entrance and exit are assumed negligible and there is thus no velocity component in the radial direction. The absolute magnitude of the shear rate at the wall, y , w  can be determined from  Equation 2.15, as follows: yw  4Q \dr)  21  TZR  3  r=R  [2.16]  For non-Newtonian fluids, Equation 2.16 no longer represents the true wall shear rate, instead it yields the apparent shear rate, y . A  The total pressure drop for flow from the reservoir, through the capillary and out to the ambient pressure (Equation 2.13) can be thought to consist o f three components: AP = AP ,+AP ent  where AP  enl  + AP=AP,+AP cop  exit  [2.171  end  cap  is the excess pressure drop due to entrance flow, AP  cnp  I  J  is the pressure drop for fully  developed flow in the capillary, and AP u is the excess pressure drop due to exit flow. The end ex  correction, AP <i, can be determined by using a technique outlined by Bagley. In this correction, en  the driving pressure, Pj, is measured for various values o f the flow rate using a variety o f dies with different length-to-diameter ratios. The value o f the driving pressure, Pa, for a length-todiameter value equal to zero is then obtained by extrapolation. This correction is known as the Bagley correction and Equation 2.14 can be modified to estimate the true wall shear stress as: W-AP ,„) 4(L/D) e  2.4.2 Flow in a rectangular  [2.18]  channel  Equations 2.14 and 2.16 apply for a fluid flowing in a cylindrical channel. When a fluid flows through a rectangular channel in which the width, W,\s much larger than the thickness, H, the edges make a negligible contribution to the pressure drop and this geometry can effectively be used for rheological measurements. For the steady flow o f an incompressible fluid in such a channel, the absolute value o f  --^  the shear stress at the wall, <j , is given by (Dealy and Wissbrun, 1990): w  where A P is the pressure drop over a length o f channel, L. The apparent shear rate in a slit, which is the true wall shear rate for a Newtonian fluid is given by: r  A  = - ^ -  [2.20]  2.4.3 Parallel plate flow Measurements o f rheological properties at low shear rates and deformations (linear viscoelasticity) are usually carried out in rotational rheometers such as parallel-plate rheometers (Figure 2.2). The two plates are mounted on a common axis o f symmetry, and the sample is 22  placed between them. The upper plate is rotated at a specified angular velocity co(t) and as a result the sample is subjected to shear. The shear rate in parallel plate rheometer experiments is given by the following expression (Dealy and Wessbrun, 1990) r • co Y=-  [2.21]  H  where co is a rotational speed, r is the distance from the center o f the plate, and H is the gap size between plates. The shear rate in the gap is not uniform which makes it impossible to calculate values o f material functions on the basis o f a single experiment, and differentiation data is required as indicated by the following equations, which are obtained by performing a force balance (Dealy and Wessbrun, 1990) 2F  1+  KR'  1  d\nF  2  d\ny  =  N (y )-N (y ) ]  R  2  R  [2.22]  R  3T 2nR'y  1+  1  d\nT  [2.23]  3 d In y  R  R  where R is the radius o f the plates, F and T are the force and the torque needed to rotate the upper plate, respectively, N/ and N  2  are the forces exerted by the material perpendicular and  along the upper plate, respectively, and T](YR) is the viscosity at the shear rate value calculated at r = R. en  H  "|  Fluid sampk  R Pressure transducer  Figure 2.2: , Parallel plate rheometer. However, in determining the linear viscoelastic properties o f a material (small amplitude oscillating experiments), the storage modulus, G\ and the loss modulus, G", can be calculated as follows (Dealy 1990): 23  G'=  2HM  cos 5  0  2HM  [2.24]  sin 8  0  [2.25]  where M is the torque oscillating amplitude, <p is an angular amplitude, and 8 is the loss angle. a  0  The complex viscosity, rj*, which approximately equals the real viscosity under small deformation, can be calculated as: ( G'~)2 VI ® J  +  [2.26] [co J  where co is the frequency o f oscillations. 2.4.4 Extensional  Rheometer  The Sentmanat Extensional Rheometer  ( S E R ) is suitable to perform  extensional  rehology studies o f strip-shape samples. A schematic o f this rheometer is shown in Figure 2.3. This rheometer is attached to the concentric disks rheometer described in section 2.4.3. The rheometer consists o f paired master and slave wind up drums mechanically coupled by a gear (Figure 2.3). The rotational motion o f the Bohlin V O R motor results in a rotation o f the master drum and an equal but opposite rotation o f the slave drum. A s a result, the sample, which is secured on the drums by means o f a pair o f clamps, is subjected to a uniform extensional deformation. A data acquisition system allows collecting data at a rate o f up to 5000 point per seconds. Lo  Master Drum  Slave Drum  Clips  Sample  Intermejshing Gears  Figure 2.3:  Schematic o f Sentmanat Extensional Rheometer (SER). 24  The Hencky strain, s , in an extensional flow is defined as: H  s  H  [2.27]  = In K oJ L  where L is the length o f the specimen at any time and L is the initial sample length. 0  The Hencky strain rate is then obtained by taking the derivative o f the Hencky strain with respect to time .  de^^dL  [  dt  2  2  g  ]  L dt  For small deformations, the length remians approximately constant at the initial length and Equation 2.28 becomes * * = - — " L dt  [2-29]  0  Since the change o f length with respect to time is essentially the linear velocity at which the sample is been stretched, it can be converted into angular velocity taking into account the drive shaft rotation rate, Q, and the radius o f the drums, R. Thus, Equation 2.29 becomes 2Q7?  £n =  [2.30]  The instantaneous torque reading, T(t), acquired from the instrument can be converted into instantaneous force, F(t), by: T(ty=2RF(t)  [2.31]  The instantaneous cross-sectional area, A(t), o f the stretched specimen changes with respect to the initial cross-sectional area, A , in an exponential fashion as follows: 0  A(t) = A exp(-e t) 0  [2.32]  H  The tensile stress, O~E, can be then estimated as  cr.-ga. A(t)  [2-33]  F ( , )  A exp(-£„0 a  Finally, for a constant Hencky strain rate, the tensile stress growth function, r/ (t), o f +  E  the stretched sample can be expressed as  ,;  ( 0  =-QiL=I_fW A(t)s  H  25  A  o  [2.34]  exp(-e t)e H  H  2.5 Stress-strain curves Stress-strain curves are extremely important and useful representations o f the mechanical properties o f a material. Included in the information that can be obtained from these plots are the elastic modulus, the elongation at yield, the yield stress, the elongation at break, the modulus o f resilience and the modulus o f toughness. Figure 2.4 shows a generalized tensile stress-strain curve for nearly any plastic material provided that: 1) the curve is obtained by a constant rate-ofstrain test, and 2) the rupture o f the test specimen may occur at any point on the curve (Carswell, 1944).  Elongation at break Elongation  Strain, s Figure 2.4: A typical stress-strain curve o f a material subjected to tension.  A ductile material subjected to tension, w i l l display a plot similar to Figure 2.4 whereas a non-ductile material w i l l show just the initial portion o f the curve i n this figure. From this, figure, the following mechanical properties o f a material can be defined: a.  The highest stress at which the stress is still directly proportional to the strain is called the proportional  limit. The ratio o f the stress to strain in this straight portion is known as  the modulus of elasticity or Young's modulus. This is usually a measure o f the stiffness of the material  26  b. The maximum stress the material can sustain without any permanent strain remaining upon the full release o f load is called the elastic limit. c.  The stress at the first knee in the curve is known as the yield point and it is an indication o f the strength o f the material and o f its resistance to permanent deformation. Before this point is reached, any deformation the material undergoes is reversible and thus is a measure o f the elastic  deformation.  After this point, any elongation the material  experiences is not easily recovered and, hence, is a measure o f the plastic  deformation.  d. The area under the stress-strain curve up to the elastic limit is the amount o f energy absorbed by the material i n the elastic range and is called the modulus of  resilience.  Material with a high yield stress and a low modulus o f elasticity w i l l have good resilience. e.  The area under the stress-strain curve, which represents the work required to fracture the test specimen, is a rough measure o f toughness and is known as modulus of toughness.  f.  The stress at the breaking point is known as the ultimate strength and it is a measure o f the force required to break the material completely. The parameters mentioned above influence the shape o f the plot and allow us to classify  the material based on its mechanical properties as shown i n Figure 2.5.  Figure 2.5:  Classification o f material based on the shape o f the stress-strain curves: (a) soft & weak; (b) hard & brittle; (c) hard & strong; (d) soft & tough; and (e) hard & tough. 27  In general, soft, weak materials show a low modulus, a l o w yield point and a low elongation at rupture. Hard, brittle materials have a high modulus, no w e l l defined yield point and a low elongation at break. Soft, tough materials are chracaterized b y a low modulus, a low yield point, a high elongation and a high stress at break. Hard, strong materials have a high modulus, a high yield point, a high elongation, and a high breaking stress.  28  CHAPTER 3 Scope of Work 3.1 Introduction P T F E is a technologically important material that has become essential to our lives. Because o f its outstanding properties, it is employed i n a wide variety o f applications ranging from wire insulation to body part replacement. It is undeniable that the processing techniques b y which P T F E is manufactured have been improved since their introduction. However there are still several issues to be understood i n order to further optimize these techniques. Because it is a thermoplastic, initial attempts were made to melt process P T F E as is done for normal polymers. However, due to its high melting point and high viscosity, such techniques are generally unsuccesful. The second possibility is to treat P T F E as paste. Techniques for processing pastes are known for other materials such as ceramic pastes, food stuffs and metallic powders  (Chevalier et al., 1997; Steffe  et al., 1996; Rough, 2000). Since the  PTFE  manufacturing process involves the production o f powders, it is not surprising that P T F E is processed using techniques  such as pressing, paste extrusion and sintering. While paste  extrusion has been studied for other materials to a great extent, it is only recently that P T F E paste extrusion has been the subject o f scientific studies (e.g. Huang et al., 2005; Ariawan et al., 2002b). The present work intends to contribute to our understanding o f several aspects o f P T F E paste rheology and its role i n extrusion. The various objectives o f this work are discussed in detail in the next section.  3.2 Thesis objectives The objectives o f this work can be summarized as follows: 1. To study the preform behavior o f several types o f P T F E fine powder pastes using two barrels o f different size and a variety o f lubricants having different physical properties, namely viscosity and surface tension. 2.  To study the rheology o f several types o f P T F E fine powder pastes using a variety o f  rheometers including parallel plates, capillary and extensional ( S E R ) rheometers.  29  3.  To assess the processability o f P T F E fine powder pastes by means o f capillary rheometry  that closely simulates the ram extrusion process. 4. To determine the effects o f die design, resin molecular structure, and processing conditions on the rheology o f P T F E pastes and the mechanical properties o f P T F E paste extrudates. 5. To study the shear/extensional-induced morphological changes during processing by means o f Scanning Electron Microscopy ( S E M ) and relate them to the rheology o f the pastes. 6.  To develop a constitutive equation that accurately describes the rheological data derived  from the paste extrusion process and rheological studies.  3.3 Thesis organization The first chapter o f the thesis discusses basic information related to tetrafluoroethylene (TFE) polymerization techniques as well as the basic physical and chemical properties o f P T F E that arise from its molecular conformation. The industrial processes relevant to P T F E are also discussed with particular attention to fine powder P T F E processes. Chapter 2 includes literature related to paste extrusion. The basic equations that underline the principles o f the operation o f the various pieces o f equipment used to Theologically characterize the materials under study are presented. Definitions related to shear and extensional rheology are also included in order to familiarize the reader with the terminology used in the rest o f this work. Chapter 3 includes the objectives o f the present work as well as the organization o f the thesis. In Chapter 4 the experimental procedure followed to achieve the objectives is described in detail. The physical properties o f the materials used for this study as well as the experimental equipment and procedure are also included here. Chapter 5 presents the performing aspect o f P T F E pastes. L i q u i d migration and in general the effects o f physical properties on P T F E performing are examined. This chapter is based on a journal paper that has already been published (Ochoa I. and S. G . Hatzikiriakos, " P T F E Paste Preforming: Viscosity and Surface Tension Effects," Powder Technology, 146, 7383 (2004)). Chapter 6 focuses on P T F E paste processing, namely paste extrusion. The effects o f various operating variables and physical properties o f lubricants are discussed. The quality o f the extrudates are examined and studied in terms o f their mechanical properties and the 30  quantity/quality o f the fibrils formed during extrusion. This chapter is also based on a journal paper that has already been published (Ochoa I. and S. G . Hatzikiriakos, "Paste Extrusion o f P T F E : Viscosity and Surface Tension Effects," Powder Technology, 153, 108-118 (2005)). Chapter 7 discusses the rheology o f unprocessed paste and the extensional rheology o f the final produced extrudates. A constitutive equation that is aimed at describing the rheology o f paste before and after the extrusion is also presented. Chapter 8 examines the extrusion o f various P T F E blends using different types o f fine powders and processing aids that have been found useful in the processing o f molten polymers. The effects o f blending pastes made from different resins on the mechanical properties o f the extrudates are also discussed. Finally, the conclusions and contributions to knowledge are discussed i n Chapter 9. A general summary o f the most significant experimental findings resulting from this work and some recommendations for future work are also presented.  31  CHAPTER 4 Experimental Equipment, Materials and Procedures 4.1 Introduction This chapter discusses the experimental equipment and materials used to study and characterise the rheological and processing properties o f P T F E pastes. It also describes in detail the conditions at which the different  stages o f P T F E  processing are conducted. The  measurements o f the physical properties such as surface tension, density and viscosity o f the different materials used i n this work are also explained in detail. The process o f paste extrusion is explained briefly. The various dies used to extrude the P T F E pastes as well as their detailed geometric characteristics are also presented. The slit die is a new device introduced i n this work to prepare samples i n order to study the extensional rheology o f the extrudates with the help o f a new extensional rheometer. Finally, the various instruments used to characterize the P T F E extrudates are mentioned. The properties measured with these pieces o f equipment are correlated with the mechanical properties o f the P T F E extrudates.  4.2 Material 4.2.1 PTFE fine powder resins Four different P T F E fine powder resins were used in this work. A l l the resins were supplied by Daikin Industries Ltd. The primary particles which form the basis o f these P T F E fine powder resins are extremely small, measuring approximately 0.2-0.4 urn (Daikin technical bulletin). They are almost spherical in shape and their properties are summarized in Table 4.1. In appearance, a large number o f these tiny particles aggregate to form secondary particles o f approximately 500 urn as shown in Figure 4.1. A s shown in Table 4.1, two resins have homopolymer structures and the other two have modified structures (copolymers). The resins have been classified according to A S T M D-1457-92, D-4895-98 standards and the physical properties were determined according to A S T M D-4895.  32  Figure 4.1: Table 4.1  S E M image o f F104 H M W fine powder resin.  Physical properties o f P T F E fine powder resin studied in this work, as provided by the supplier.  4.2.2  Resin  Type  F104 H M W F104 L M W F301 F303  Homopolymer Homopolymer Modified Modified  Particle Diameter (urn) 400-650 400-650 400-650 400-700  Apparent Density (g/ml) 0.45 0.45 0.45 0.45  Specific Gravity 2.17-2.20 2.16-2.18 2.15-2.18 2.14-2.16  Lubricants A n extrusion aid is added to P T F E fine powder as a lubricant to enable smooth, uniform  paste extrusion. The extrusion aid must be able to completely wet the resin particles, and must be easily removable from the product after extrusion. If a sintering stage is present in the process, the extrusion aid must not color the product. In other words, the volatilizing temperature o f the extrusion aid must be lower than the sintering temperature. The types and amounts o f the extrusion aid ordinarily used depend on the application o f the final product. Different lubricants were used as processing aid in this work in order to examine the effects o f their physical properties on the extrusion process. First, several isoparafinnic liquids under the trade name o f ISOPAR® were supplied by E x x o n M o b i l Chemicals. These lubricants are colorless, synthetically produced solvents with uniform composition, low reactivity, excellent stability and narrow boiling point ranges. The 33  odour o f ISOPAR® ranges from almost undetectable to very m i l d depending on the grade. Perhaps the most important property o f these lubricants is their low surface tension which promotes good spreadability o f the lubricant between the resin particles. The other liquid used as a lubricant in this work was a perfluorinated liquid with the chemical name 2-trifluoromethyl-3-ethoxydodecafluorohexane, and referred to as HFE-7500. This is also a clear, colorless liquid with a viscosity and surface tension even lower than those o f the Isopar® lubricants. A l l the lubricants were used as they were received. Their most important physical properties relevant to this work are listed i n Table 4.2. Table 4.2:  Physical Properties o f Isopar® and H F E - 7 5 0 0 lubricants.  E  0.70  0.72  0.74  0.79  0.82  1.61  21.2  22.5  23.5  26.6  30.8  16.2  98.0  52.0  14.0  3.1  0.3  15.7  0.51  0.62  1.00  2.70  7.50  1.24  2031.1  1886.5  1305.4  608.8  272.6  3386.5  Density, g/cm 25°C Surface Tension, dynes/cm 25°C Vapour Pressure, m m H g , 38°C Viscosity, mPa s 25°C {2)  (g s" cm" ) 2  1  5  V  C J  a  M  HFE7500  Isopar G  Property  (I)  (1) Value of vapour pressure reported at 25°C. (2) Group of physical properties given by a = pfyJ ) " 7  s e d i n  Equation 4.12  4.2.3 Other lubricants Dioctyl sulfosuccinate sodium salt ( A O T ) was used as a surfactant to alter the surface tension o f Isopar®G due its ability to create reverse micelles in nonpolar solvents. A O T is an anionic surfactant with the structure shown in Figure 4.2. The longest portion o f the molecule is the hydrocarbon (hydrophobic) part; whereas the other end is the hydrophilic part (polar). The main idea o f the incorporation o f this surfactant into Isopar®G is to modify the surface tension (wettability with P T F E ) o f the lubricant while maintaining the viscosity or viceversa. Solutions at different concentrations were prepared by dissolving known amounts o f A O T in Isopar  G.  The surface properties o f the mixture obtained in this way were measured by using the capillary rise method (Laskowski, 2001; Seth et al., 2001; Siebold et al., 1997), direct contact angle 34  measurements (Rathod, 2004; Anastasiadis and Hatzikiriakos, 1998; H u h and Reed, 1983) and the D u N o u y ring method. Their viscosities were measured b y using a Cannon-Fenske Opaque (Reverse-Flow) viscometer and their densities where measured with a 10 m l picnometer. These values are listed i n Table 4.3. It can be seen that the viscosity varies significantly while the surface tension remains almost constant.  Figure 4.2:  Table 4.3:  Molecular structure o f dioctyl sulfosuccinate sodium salt ( A O T )  Physical Properties o f Isopar® G - A O T solutions at 25°C.  Concentration (wt %) 0.0 10.0 20.2 30.3 39.4 56.5  Density (g/cm ) 0.75 0.77 0.80 0.82 0.86 0.90 3  Surface Tension (dynes/cm) 23.5 23.6 23.6 23.3 23.4 23.9  Viscosity (mPa*s) 1.00 1.22 1.67 2.33 4.48 10.05  Additional to the mixtures o f Isopar® G + A O T , mixtures o f Isopar  G + Isopar  V at  different weight concentrations were prepared. The physical properties o f these mixtures were also measured and they are listed in Table 4.4.  Table 4.4:  Physical Properties o f Isopar® G & Isopar® V solutions at 25°C.  Concentration ofIsopar® V (wt %) 0 20 40 60 80 100  Density (g/cm )  Viscosity (mPa*s)  Surface Tension (dynes/cm)  0.74 0.76 0.77 0.79 0.80 0.82  1.00 1.54 2.17 3.36 5.43 7.50  23.5 26.0 27.3 28.4 29.6 30.8  3  35  :  .  4.3 Experimental equipment 4.3.1 Preforming and  extrusion  A s previously discussed, paste extrusion is a two-step process: preforming followed by extrusion. During preforming, important phenomena such as liquid migration and densification occur and these are investigated in the present thesis. Preforming and extrusion were performed by using an Instron tensile tester machine model 1123. A s shown i n Figure 4.3, the instrument consists o f a barrel, a motor drive, a load cell and a data acquisition system. T w o interchangeable barrels o f 9.525 m m and 25.4 m m inner diameter were available. The smaller barrel was used for the paste extrusion study while both barrels were used in the liquid migration study. Both barrels have heating bands and temperature controllers. A load cell o f 2270 k g with a plunger attached to it is mounted on a mobile stage. The motor drive allows moving the stage at the specific speed entered into the control panel. The load cell senses the resistance to flow applied by the paste contained in the barrel through the plunger and sends it to the data acquisition system. The data acquisition board allows the experimental results to be recorded automatically and stored in the computer. The dies were attached at the lower end o f the barrel as shown in Figure 4.3. For preforming and liquid migration studies, a blind die was used. For extrusion purposes, different tapered dies made o f stainless steel were utilized. For these dies, the design variables o f interest were the die entrance angle, 2a, the die reduction ratio, RR, and the die length-to-diameter ratio, L/D, as shown in Figure 4.4. The die reduction ratio is defined as the square o f the ratio o f the diameter o f the barrel, Db, to the diameter o f the capillary, D, that is, the ratio o f the crosssectional areas before and after the contraction zone. RR =  [4.1]  D X  s  To prepare samples suitable for extensional rheological studies, a tapered slit die was used. L i k e the cylindrical dies, this die also has a conical zone which gradually turns from a round shape to a rectangular one, thereby merging into the slit-shaped land o f the die. Figure 4.5 depicts this transition and the geometrical characteristics o f the slit die. This die produces a rectangular shaped sample o f high aspect ratio that can be easily loaded onto the S E R extensional rheometer.  36-  Load C e l l  Temperature Controllers  Data Acquisition System  r  Thermocounle  Figure 4.3:  ^  l l IBM Compatible  1  Set-up o f the Instron tensile tester machine for paste preforming and extrusion.  Barrel diameter D b  T Contraction zone  Capillary zone L  Capillary diameter D Figure 4.4:  Schematic diagram o f a typical cylindrical capillary die along with the definition o f the design parameters.  37  Figure 4.5:  Schematic diagram o f the tapered slit die showing its contraction and the land zones.  4.3.2 Mechanical and viscoelasticproperties VOR  measurement  and C V O R parallel-plates Bohlin rheometers were used to characterize the  rheology o f the fine powder resins by measuring their viscoelastic properties. The former is a strain-controlled rheometer while the latter is a stress-controlled rheometer. A Sentmanat extensional rheometer (SER) was also attached to the V O R Bohlin rheometer in order to measure the extensional properties o f the extrudates obtained from the slit die. A schematic diagram o f the S E R rheometer has been shown previously in Figure 2.3. The principles and the details o f loading the sample and operation were discussed in Chapter 2. The mechanical properties o f the extrudates produced from the capillary dies were measured using a C O M - T E N (Compression & Tensile Strength) apparatus whose set-up is shown in Figure 4.6. This instrument comprises a motor drive with adjustable speed, two interchangeable load cells o f 20 N and 200 N , a digital monitor controller ( D M C ) and the data acquisition system ( D A S ) . The sample is placed and held with two clamps, the upper clamp being fixed at the shaft activated by the motor. The sample is then stretched at a constant speed until it fails. The load cell senses the force applied by the sample and sends it to the D A S via the DMC.  38  4.3.3 Other equipment After extrusion and prior to measuring the mechanical properties o f the extrudates, the lubricant had to be removed from the samples. In order to do this, the samples were dried in a vacuum oven for at least 16 hours at 120°C. The mechanical properties o f the sintered extrudates were also determined as mentioned before. After removing the lubricant from the extrudates, the samples were sintered at 370°C for about 60 s. For the sintering process, a vertical tubular furnace o f 2.54 c m inner diameter and 35 cm in length was used. The furnace consists o f a single heater zone built o f two ceramic circular brick halves that contain the heater elements and a single temperature control with a J type thermocouple. A differential scanning calorimeter ( D S C ) was also used to determine the thermal properties o f the fine powder resins before and after extrusion. In this way changes in the crystallinity o f the samples can be detected and correlated with the extent o f fibrillation during extrusion.  Motor  Load  Sample  Figures 4.6:  Set-up o f C O M - T E N tester to measure the mechanical properties o f extrudates.  39  4.4 Experimental procedure 4.4.1 Paste  preparation  Paste preparation is the first step in P T F E paste extrusion. The fine powder is mixed with the processing aid (lubricant) at the proper concentration. Even though this step sounds simple to perform, it is very important that the lubricant wets the particle resin properly. A typical lubricant concentration varies from 16 wt% to 25 wt% (Daikin technical bulletin; Ebnesajad, 2000). In a previous study it was found that a higher lubricant concentration produces preforms with less density variation although it does increase the extent o f lubricant migration, i.e., the preform w i l l have a larger lubricant concentration gradient (Ariawan et al., 2002a). Benbow and Brigwater (1993) have reported typical lubricant concentration o f 35 v o l % to 50 v o l % , depending on the physical properties o f the solid component, such as the particle shape and size distribution. However, the physical properties o f the lubricant play an important role in paste preparation and the effects are addressed in the present study. Assuming particles o f spherical shape in a cubic arrangement, as shown in Figure 4.7, the volume o f the void space that can be occupied by the lubricant can be calculated as follows. First, the volume o f the cubic unit cell, V , (see Figure 4.7) is: c  [4.2] where r is the radius o f the spherical particle. Thus, the volume occupied by the particles, V , p  p  is: VP =-7V 3  [4.3]  r\ P  Then the volume fraction that can be occupied by the lubricant is:  V _ j  Vp_ _ j _ n_  K  K ~  c  [4.4]  6  c  where Vi is the volume occupied by the lubricant. This shows that 47.6% o f the volume is void and it could be occupied by lubricant i f used in the right proportion. However, the amount o f lubricant used is less and this leaves voids inside the paste that must be removed during the preforming step. The amount o f lubricant in wt.% can be estimated as follow: wt%  _ p,  l-wt%  p  s  40  vol% l-vo/%  [4.5]  Mixtures o f various powders and lubricants were blended i n a jar m i l l at a speed o f 15 rpm for about 20 minutes. The resulting paste was then aged for at least 16 hours at room or lower temperature depending on the vapour pressure o f the lubricant. A g i n g the paste is important to allow uniform wetting o f the resin particles by the lubricant. It should be pointed out that the paste was prepared in a wide container made o f either glass or P E T in order to avoid changes in the lubricant concentration. After aging, the concentration o f the lubricant in the paste was measured by taking samples at different depths. These were weighed, dried and weighed again. The difference in weigh before and after drying was attributed to the presence o f the lubricant.  Figure 4.7:  M a x i m u m packing o f the solid phase in P T F E paste.  4.4.2 Capillary and sessile drop experiments To determine the surface properties o f the various lubricants used in the experimental work as well as their individual wetting characteristics with P T F E , (i) the capillary rise o f a liquid through a tube containing a compacted P T F E powder was measured and (ii) direct contact angle measurements o f the liquids on P T F E substrates were used. Prior to the capillary rise experiment, it was necessary to determine how the density and the porosity o f the material change with pressure. This could help to determine the conditions at which the P T F E powder resin should be packed. For this purpose, 5 g o f resin were poured into a copper tube (open from both ends) o f 14.1 m m inner diameter and 100 m m length and 41  compacted at different pressures by using the Instron capillary rheometer (Figure 4.3). The pressure was applied for 30 minutes. After that, the compacted powder was removed from the tube, weighed and the bulk density was estimated by using Archimedes' principle. The porosity o f the material is defined as: h e = -^V  [4.6]  hs where Vh and Vf, are the volume o f the pores and the porous material, respectively. s  Equation 4.6 can also be written as: s = \  where p and p s  hs  [4.7]  ^-  are the densities o f the solids and the porous material, respectively.  Figure 4.8 shows how the bulk density o f the compacted powder changes with applied pressure. Figure 4.9 depicts the change o f porosity with pressure. It can be seen from Figures 4.8 and 4.9 that the bulk density and porosity change less significantly at pressures greater than about 20 M P a .  1.2 0  5  10  15  20  25  30  35  40  45  Pressure (MPa)  Figure 4.8:  Variation o f density o f a compacted P T F E resin as a function o f pressure.  42  Even though the results shown i n Figures 4.8 and 4.9 seem quite reproducible, a pressure of 3 M P a was used to pack the powder. Thus, a copper tube was filled with 5 g o f P T F E fine powder and then compressed at a pressure o f 3 M P a for 30 minutes. The filled tube was hung from the bottom o f an analytical balance and its lower end was then placed in a reservoir containing the liquid whose wetting characteristics with P T F E were to be examined (see Figure 4.10). Depending on the surface properties o f the liquid, it permeated up into the powder at a different rate. The weight o f the column was recorded as a function o f time by means o f the electronic balance from which the tube was hung. The data acquisition system can be programmed to record data at a minimum rate o f 10 points per second. This modified technique was developed and used successfully in the past by Laskowski et.al. (1996; 2001) for mineral systems.  0.5  i  |  T —  1  ->—r  — r  0.4  'to (A O  o Q_  0.3  0.2 h  0.1 k  0.0 "—>• 0  5  10  15  20  25  30  35  40  45  Pressure (MPa)  Figure 4.9:  Variation o f porosity, 8, o f a compacted P T F E resin as a function o f pressure.  The data analysis o f the capillary rise method is based on Washburn's equation:  h = 2  kr y c  L  cos6  2^  [4.8]  .  where h is the rise height o f the liquid front in the capillary tube, r is the mean radius o f the c  capillary, k is a constant accounting for the tortuosity that depends on the particle size and  43  packing, y is the surface tension o f the liquid, 9 is the contact angle, 77 is the viscosity o f the L  liquid and t is the time.  Electronic Balance  Tube with sample Beaker with test liquid  Lab-jack  Figure 4.10:  Capillary rise method to determine the surface energy o f a powder or a liquid.  The height is related to the mass o f liquid, A w , which has penetrated the column by . Aw = sp,7rRfh  [4.9]  where e is the porosity o f the packed powder column, p, the density o f the liquid, and R, the inner radius o f the tube. Thus, the modified Washburn equation can be written as (Laskowski et. a l , 1996; Laskowski, 2001) p, y 2  Aw = c  L  cos0  [4.10]  where the term c is a geometric factor, which is a constant as long as the packing and the particle size remain the same. It has to be experimentally determined for each type o f particle packing. The last equation can be rewritten as Aw  =m-t  [4.11]  p, y, cosd where m = c—-—. Thus, from a plot o f Aw versus t a straight line should be obtained 2  whose slope m depends on the geometric factor c, and the contact angle 9. These parameters can 44  be found with the help o f a liquid that wets the packed solid particles completely {0 & 0°). H F E 7500 was assumed to be such a liquid, since it has a surface tension lower than that o f P T F E . Figures 4.11 and 4.12 show capillary rise experimental runs for lubricants Isopar® G and H F E 7500, respectively. Only the straight portion o f the individual curves was considered in the calculations o f the slopes. A n average slope was determined from the multiple experimental runs. It is noted that HFE-7500, which has a smaller surface tension compared to P T F E (implies excellent wetting properties with PTFE-assumed contact angle o f about 0°) has shown the smaller standard deviation together with Isopar® C . Liquids such as Isopar® G , M and V whose surface tensions are higher than that o f P T F E have shown larger standard deviations in repeated experiments. From Equation 4.11, the slope o f the straight portions o f the curves depicted in Figures 4.11 and 4.12 can be written as m = a-c- cosd  [4.12]  where c is the geometric factor discussed before and a = pfy' /TJ L  is a constant, which groups  the physical properties o f the liquid involved such as density pi, viscosity rj, and surface tension, YL- A s mentioned before, the geometric factor c depends on the packing and the particle size (these essentially remain constant). The parameter a changes since it depends on the operating conditions (T and P) and the type o f liquid used. Table 4.2 lists the physical properties o f the Isopar  and H F E - 7 5 0 0 lubricants as well as their corresponding values o f the parameter a  determined from these capillary rise experiments. Equation 4.12 applied to a lubricant "z" for the same packed powder can be written as: m, = a-c-  cos9  ;  [4.13]  or c =  a, cos 6,  N o w , combining Equation 4.14 for liquids  [4.14]  , "f and "k" results:  [4.15]  45  If the contact angle o f any o f the above liquids with a particular P T F E powder is known, then the contact angle o f the other liquids with this powder can be calculated.  —>  1  T"  1  1  1  •  /// Run#1 Run #2 Run #3  ™ / 's / /y '/ 0  500  •  •  1000  1500  2000  2500  3000  3500  4000  Time (s)  Figure 4.11:  Capillary rise experiment o f Isopar® G through a tube filled with P T F E powder at 25°C.  3.0  Time (s)  Figure 4.12:  Capillary rise experiment o f H F E 7500 through a tube filled with P T F E powder at 25°C. 46  To determine the validity o f Equation 4.15, sessile drop measurements using the Isopar" liquids on a P T F E substrate were also carried out. For a direct contact angle measurement, a drop i f the test liquid was placed on a substrate made o f compacted P T F E . The substrate was prepared by pouring 2 g o f powder into a 27.0 mm inner diameter ring and compressing at 5 M P a for 10 minutes. This way, a relatively smooth P T F E substrate can be prepared. The drop was placed on the substrate with a syringe kept in a vertical position and a few centimeters from the surface by a holder. Instead o f obtaining a still picture, a movie was shot with a C C D camera from the moment the drop left the syringe until it touched the surface o f the substrate. The movie was subsequently analysed frame by frame to find the precise moment at which the droplet equilibrates on the substrate. Figures 4.13a and 4.13b show typical pictures o f drops o f water and Isopar® V on P T F E substrates. The average contact angle for Isopar® V was found to be 54.5° ±1.8° and that for Isopar® M  49.7°±3.2°. These values are very close to those estimated by using the Young-Dupre  equation. This equation can be written as follows (Wu, 1982): r  t + cos0=2Jyf  [4.16] **• J  (a) F i g u r e 4.13:  {  n  )  (b)  Drops o f liquid placed on P T F E substrate, (a) Water, (b) Isopar® V .  where y® is the dispersive (non-polar) contribution to the surface energy o f the substrate ( P T F E is this case) and y^  D  is the non-dispersive (polar) contribution to the surface energy o f the  substrate. Since, P T F E is a non-polar polymer (repelling water), it is reasonable to assume t h a t ^ ^ = 0 . Thus, the second term in the right side o f Equation 4.16 can be neglected. In 47  addition, i f we assume that none o f the liquids have a polar contribution component to their surface tension (hydrocarbon liquids), t h e n / f = y and Equation 4.16 can be written as follows: L  cos 0 - 2  -1  [4.17]  Table 4.5 summarizes the surface tension and contact angle values for P T F E and several liquid lubricants calculated from Equations 4.15 and 4.17. The contact angle values measured from the sessile drop method are also listed. The agreement seems satisfactory. The work o f spreading, W , in Table 4.5 was calculated from: s  W =y (cos0-l) s  [4.18]  L  and the work o f immersion, Wt, calculated from: [4.19]  W - yi cos 6 i  The estimation o f W  s  and W; involved the values o f YL reported by the manufacturer and  6 estimated with Equation 4.17. According to Equation 4.18, 0=0 means that the liquid w i l l readily spread over the solid (W = 0). The more negative the value o f Ws is, the poorer is its wetting characteristic with the s  solid substrate. O n the other hand, according to Equation 4.19, a value o f 0< n/2 {W > 0) means t  that a particle w i l l spontaneously be incorporated into the liquid. It can also be seen from Table 4.5 that the value o f the contact angle o f Isopar® V predicted by Equation 4.15 matches well with that predicted by the Young-Dupre equation. The contact angles for Isopar® V and G predicted by Equation 4.15 fall within +10% o f the values predicted by Equation 4.17. In most cases, the various methods used for estimating the wetting characteristics o f the various lubricants agree relatively w e l l giving confidence about the consistency o f the experimental results. It is noted that direct contact angle and capillary rise experiments with the more volatile liquids could not be performed as these evaporate easily and the increase o f weight o f the liquid penetrating within the column can not be established with certainty. Using the experimental results listed in Table 4.5, it can be concluded that the wettability of Isobar ® lubricants decreases with an increase o f density and viscosity. While one might  48  expect that a higher viscosity would be desirable in paste extrusion (minimum liquid migration), the poor wettability would not lubricate properly the flow as w e l l as would not help in overcoming the friction developed between the individual particles ( M c L e o d , 1977). This might result into a small degree o f fibrillation that would have a detrimental effect on the mechanical properties o f the final P T F E extrudates. A s can also be seen from Table 4.5, lubricants having about the same viscosity and different surface tension (various wettabilities with P T F E ) are not available. Therefore, the effect o f surface tension cannot be assessed, independently, perhaps with the exception o f the pair o f HFE-7500 and Isopar® G . The effect o f viscosity can be assessed by using all Isopar® lubricants to some extent, which in some cases have significantly different viscosities and similar surface tensions, i.e. comparing the performance o f Isopar® C , E and G (although C is very volatile and difficult to use). Therefore, some other lubricants should be prepared in order to study these effects. B e l o w a surfactant ( A O T ) is used as an additive to Isopar® G in order to prepare additional lubricants having various viscosities and surface tension's' that result in various wettabilities with P T F E . Table 4.5:  Comparison o f contact angles o f various lubricants with a P T F E substrate obtained by the Young-Dupre equation, the capillary rise and sessile drop methods. **  YL  f dynes ^  I  cm  J  'dynes^  I  cm  w  Contact Angle, 9 (°)  YL  J  W  s  f  dynes  t  N  ( dynes "|  YoungDupre  Sessile Drop  Capillary rise  -  -  -  -  37.3  -3.33  17.87  36.1  -4.75  17.75  -5.87  17.63  I  cm  J  I  cm  PTFE  18.0  -  -  Isopar® C  21.2  20.5  32.6  Isopar® E  22.5  22.3  37.9  Isopar® G  23.5  23.7  41.4  25.5±4.8  40.0  Isopar® M  26.6  26.8  49.8  49.7±3.2  50.4  -9.44  17.16  30.8  28.1  58.1  54.5±1.8  63.1  -14.51  16.29  0  -  0  0  16.20  Isopar® V HFE-7500  16.2  17.0  J  * Values reported by the manufacturer ** Values measured by using du Noiiy Ring method.  4.4.3 Effect of surfactant on lubricant surface tension & wettability A s already discussed above, the main idea o f adding a surfactant to the lubricants is to see i f it is possible to modify their surface tensions so that the effects o f surface tension (wettability with P T F E ) and viscosity on paste extrusion could be studied independently. The 49  presence o f this chemical in Isopar® G affects not only the viscosity o f the lubricant, but also its density. Table 4.3 summarizes the physical properties o f Isopar® G solutions as a function o f A O T concentration. It can be seen that the viscosity increases exponentially while the density increases linearly with increase o f the surfactant concentration. The surface tension values shown in Table 4.3 represent average values o f several measurements using the du N o i i y ring tensiometer. This method falls in the detachment method classification o f measuring the surface and interfacial tension o f liquids. It is based on measuring the maximum equilibrium force required to detach a circular ring from a liquid surface. The surface tension, y , can be determined from L  F  ' R R  R  Y l  " AnR  R ^  R  [4.20]  f R  where F is the force required to detach a circular ring from a liquid surface; RR and R are the R  R  radius o f the ring and the ring's wire, respectively;/is a correction factor; VR is the volume o f the liquid raised above the flat level o f the liquid surface and is given by  F /(p, R  - p )g n  where,  pi and p are the density o f the liquid and the air, respectively and g is the gravity. The correction a  factor takes into account that the rupture o f the interface occurs at the plane where the film is thinnest leaving some liquid on the ring and that neither the inner nor the outer surface is vertical (Huh et al., 1975 and 1977). The values o f surface tension in Table 4.3 have been corrected by taking into account the effect o f the atmospheric pressure and hence the density o f the solutions (Zuidema et a l , 1941). In spite o f these corrections, it seems that the surfactant does not have a significant effect on the surface tension since the variability o f these values falls within experimental error. Using these values o f surface tension, the contact angles o f the various liquids with P T F E can be determined. In addition, the works o f spreading and immersion can be calculated by means o f Equations 4.16 and 4.17. For this purpose, capillary rise experiments were conducted. The results are plotted in Figure 4.14 and straight lines were fitted to the straight portions o f the curves to determine the contact angles via Equation 4.10. Table 4.6 lists the results o f the contact angle as well as the works o f immersion and spreading calculated using Equations 4.16 and 4.17 and taking H F E - 7 5 0 0 as a reference liquid (assumed contact angle o f 0°). It can be seen that the contact angle increases with increase o f the concentration, while the total surface tension remains about the same. Based on the results in 50  Table 4.6, the viscosity o f the lubricant can be fixed and the effect o f surface tension (wettability) can be assessed in a more comprehensive way. Table 4.7 shows the physical properties o f Isopar® G - A O T solutions prepared at different concentrations to mimic the viscosity o f other lubricants, while keeping the surface tension similar to Isopar® G .  Figure 4.14:  Capillary rise experiment o f Isopar® G - A O T solutions through a tube filled with P T F E resin at 25°C.  Returning again to Table 4.6, it can be seen that the total surface tension is the same at all concentrations, although the wettability (contact angle) with P T F E changes. Due to the fact, that A O T possesses a polar group in its molecular structure, it possibly alters the individual contributions to surface tension. It is also possible that A O T may deposit a film on the P T F E surface, and the concentration o f A O T has no effect on the surface tension o f the Isopar liquid (Wu, 1982; de Gennes, 1985; Chapius, 1984). While Isopar® G is a non-polar liquid, the presence o f A O T brings polarity into the total surface energy o f these solutions. For solutions o f A O T in Isopar®G, Equation 4.16 is more appropriate to use and it takes the following form,  l + cos6 = 2Jyf  51  [4.21]  It was assumed that y"  = 0 ( P T F E is non-polar). The difference between Equations  D  4.16 and 4.21 is that n o w ^ f & y . Equation 4.21 was used to estimate the dispersive component L  of surface tension component o f the solutions. This is the column labelled as yl in Table 4.4. A s can be seen, the dispersive component o f surface tension drops with increase o f A O T concentration while the non-dispersive component increases.  Table 4.6:  Contact Angle o f Isopar® G - A O T solutions on a P T F E substrate at 25°C.  Concentration w%  y (dynes/cm)  0.00 10.03 20.22 30.26 39.41 56.50  23.5 23.6 23.6 23.3 23.4 23.9  L  0  0.75 0.76 0.71 0.69 0.67 0.46  41.4 40.2 45.0 46.3 47.6 62.5  W (dynes/cm)  W (dynes/cm)  y (dynes/cm)  -5.87 -5.56 -6.92 -7.21 -7.63 -12.87  17.63 18.04 16.68 16.09 15.77 11.03  23.5 19.1 22.5 21.6 21.3 16.9  s  (  u  L  Having now available this variety o f lubricants, the effects o f their physical properties on paste perform (axial and radial lube migration and perform density) as w e l l as paste extrusion and rheology can be assessed. This work is carried out in the subsequent chapters.  Table 4.7:  Physical properties o f Isopar® G - A O T solutions.  wt% of A O T in Isopar® G Density (g/cc) @ 25°C  Concentration of A O T (wt%) 29.11% 11.23% Isopar G 0.83 0.77 0.74  50.47% 0.88  Viscosity (mPa*s) @ 25°C  1.00  1.24  2.54  7.33  wt% of I s o p a r ® G in paste  18.0  16.5  14.3  11.5  4.4.4  Preforming Preforming was carried out using the unit described above with reference to Figure 4.2.  A blind die was attached at the bottom o f the barrel and the paste was loaded. The tests were run at 25°C. The speed o f the motor was initially set at 0.85 mm/s and manually changed until the desired pressure was reached. Once the pressure was attained, it was kept there for a certain period o f time. 52  To determine the axial variation in density and lubricant concentration, the preform was removed from the barrel by moving the piston downward once the blind die was removed. The preform was then cut in portions o f about 20 m m in length as indicated in Figure 4.15a. For radial liquid migration studies, each portion was then sliced i n sections o f about 3 mm in thickness along the axis as shown in Figure 4.15b. Additional cuts were made along the axis as shown in Figure 4.15c. O n l y the portions in the center were analysed. The pieces were immediately weighed and dried in a vacuum oven at 100°C for 24 hour to ensure that the lubricant had been removed completely. From the difference between the initial and final weights, the lubricant concentration was determined as a percentage:  (w, wt%Lub = —  25.4 mm  -w ) f  [4.19]  ^-xlOO  ECDCE  3  h-H 20 mm  (a)  \  (c)  (b) Figure 4.15:  1 ir~7  ii  (a) Preform paste slicing for determination o f the axial density variation and liquid migration, (b) Preform paste slicing for determination o f radial liquid migration, (c) Top view o f the preform sliced for radial liquid migration.  For the axial density distribution study, similar experiments were conducted where the preform was sliced as shown in Figure 4.15a. After drying, the density o f each slice was determined by means o f the Archimedes' principle. The sample was hung on the bottom o f an analytical balance and immersed in water. The buoyancy was estimated from the difference between the weight o f the P T F E portion in air and water. Water was used because it does not wet P T F E due to its high polarity and surface tension. It has been assured that water does not flow into the pores o f the P T F E slices. 53  4.4.5 Extrusion Paste extrusion was performed using the same Instron capillary rheometer. A s usual, preforming was performed prior to extrusion by using a blind die. Once this step was completed, the blind die was replaced with a tapered capillary die and the paste was extruded at constant piston speed. To study the effect o f shear rate, extrusion runs were conducted at different piston speeds. A l s o various capillary dies having different geometrical characteristics were used. The extrudates obtained were collected for further analysis. 4.4.6 Extrudate  analysis  The extrudates obtained from extrusions, were dried as discussed above. Part o f the dried extrudates was used for sintering. Dried extrudates, sintered and unsintered, were tested to determine their tensile strength by using the C O M - T E N equipment following A S T M D1710-96 standard. The extensional rheology was studied by using the S E R rheometer (Figure 2.3) attached to the V O R B o h l i n rheometer. The thermal properties o f the extrudates determined using D S C according to A S T M D3418-82 standard.  54  were  CHAPTER 5 PTFE Paste Preforming 5.1 Introduction Paste preparation is the first step in P T F E paste extrusion. In this stage, the fine powder resin is mixed with an appropriate amount o f a processing aid (lubricant). A typical lube concentration varies from 16 to 25 wt% (Daikin technical bulletin), however the optimal amount of lubricant depends on its physical properties. In fact, the physical properties o f the lubricant such as density, viscosity and surface tension, play an important role in P T F E paste preparation and they also have an enormous effect on preforming quality and extrusion. During the extrusion process, the pressure undergoes fluctuations. These variations can be present even for the steady-state region and they are the consequence o f an uneven densification o f the paste before extrusion. Correspondingly, these variations in extrusion pressure w i l l affect the quality o f the extrudate (Ariawan, 2002). Therefore, it becomes important to study carefully the factors that influence the preforming o f paste. A number o f studies have been performed on liquid migration during extrusion in the past (Benbow and Bridgwater, 1993; Burbidge et a l , 1995; Bridgwater, 1989; Ariawan et al., 2001). The subject o f liquid migration through the soil and soil consolidation has been o f special interest in soil mechanics (Craig, 1997). However, none o f these reports have addressed the effect o f the physical properties o f the liquid on the quality o f the preform. In this work, P T F E paste preforming is studied, focusing on the liquid migration and the density distribution during the preforming stage. Emphasis is placed on the effect o f the physical properties o f the lubricant on liquid migration and densification during preforming. These properties are expected to have a significant impact on the process o f paste extrusion as viscosity would influence the liquid migration whereas surface tension affects the wettability between P T F E and the lubricant and thus the quality o f the final mixture. Furthermore, these characteristics might significantly affect the quality o f the final extrudates during the extrusion stage.  5.2 Densification studies Ariawan et al. (2001) have reported that a pressure o f 2 M P a applied for a period o f 30 s ensures an even density distribution along the preforms. They have reported that while longer 55  preforming may reduce the variation in preform density, it also significantly increases the extent of lubricant migration. Thus, maintaining an applied pressure o f 2 M P a for about 30 seconds is sufficient to yield a uniform lube distribution along preforms o f high molecular weight resins (Ariawan et al., 2001). To study the effect o f applied pressure on the preform density, different types o f pastes were prepared (see Chapter 4) with three lubricants: Isopar® V , M and G at concentration o f about 18 wt% and subjected to pressures o f 1, 2 and 3 M P a . The preform was removed from the barrel, sliced and analyzed for density and lube concentration as discussed before (see section 4.4.4). Figures 5.1, 5.2 and 5.3 depict typical results. It can be seen from Figure 5.1 that a pressure o f 1 M P a is not enough to achieve a uniform density in preforms prepared with Isopar® V and Isopar® M , due to their high viscosities. These resist rapid fluid motion due to volume reduction o f the paste. However, application o f a higher pressure (2 M P a and 3 M P a ) produces a preform o f more uniform density even for the liquid having the highest viscosity (see Figures 5.2 and 5.3). 1.5  1.4  E  1.3 h  (/) c  1.2 \-  0)  Q  • — Isopar V J  1.1 r-  - — Isopar M A  - • — Isopar G 1.0  20  40  60  80  100  120  Distance from bottom (mm)  Figure 5.1:  Variation o f preform density in axial direction resulting from an applied pressure of 1 M P a for 30 s on F104 L M W resin + 18 wt% o f lubricant.  56  1.5  '  1-  P = 2 MPa T = 25°C 1.4 h E o S  1.3 I-  (/) c  (D  Q  1.2  Isopar V Isopar M Isopar G  1.1  20  40  60  80  100  120  Distance from bottom (mm)  Figure 5.2:  Variation o f preform density in axial direction resulting from an applied pressure of 2 M P a for 30 s on F104 L M W resin + 18 wt% o f lubricant. 1.5  1.4 h  E o  "5)  1.3  v> c a>  a  1.2  1.1  — • — Isopar V — ± — Isopar M  P = 3 MPa  —®— Isopar G  T = 25°C  20  40  60  80  100  120  Distance from bottom (mm)  Figure 5.3:  Variation o f preform density in axial direction resulting from an applied pressure of 3 M P a for 30 s on F104 L M W resin + 18 wt% o f lubricant.  57  5.3 Liquid migration Preforming was performed using a typical capillary rheometer according to section 4.4.4. Two different barrels were used having diameters o f 0.952 and 2.54 cm in order to ensure that the results are size independent. To analyze the liquid migration in the axial and radial directions during preforming, several experiments were conducted. The liquid content in all experiments was kept constant and equal to about 18 wt%. This is particularly important in comparing the behavior o f the various lubricants within the context o f this work and with results reported by others. It is noted that all lubricants have densities (see Table 4.2) that differ significantly in some cases and therefore a constant weight percentage does not imply a constant volume percentage. 5.3.1 Axial liquid  migration  Extensive work on axial liquid migration under different operating conditions has been carried out in the past for a variety o f P T F E resins (Ariawan et al., 2002). In this work, the effect of the physical properties o f the lubricants on preforming is studied. For that purpose different lubricants were used to prepare the paste. In most cases, the paste was subjected to a pressure o f 2 M P a for about 30 s according to the previous section (Ariawan et al., 2001). A l s o , it was preformed in two different ways: namely, one-sided and two-sided. In the former, the paste was preformed by applying pressure only on one o f the ends. In the latter, the sample was flipped over after the first application o f pressure and pressure applied again on its other side. Figure 5.4 is a plot showing the axial liquid migration o f three lubricants (Isopar® G , M and V ) . In all cases the samples were preformed one-sided (2 M P a for 30 s) using the larger diameter barrel (2.54 cm). It can be seen that the higher viscosity o f the lubricant is, the lower is the degree o f lubricant redistribution. Variability o f lubricant concentration is only present at the end o f the preform (shown as T O P ) at which the pressure is applied. In addition, the difference between the horizontal lines labelled as initial lubricant concentration and the experimentally determined profiles indicate the loss o f lubricant through evaporation during the experimental procedure. These differences were even higher for the low viscosity and high vapour pressure lubricants, i.e. Isopar® C and E . Similar experiments using the barrel having a smaller diameter (0.952 cm) gave similar results. Therefore, it can be concluded that these viscosity effects are not geometrically dependent.  58  20 I  1  1  1  '  1  '  1  '  1  '  r  i  •  •  100  120  15 h 14 t" 0  i 20  i  i  •  •  40  60  80  •  1 140  D i s t a n c e f r o m b o t t o m (mm)  Figure 5.4:  L i q u i d migration in axial direction o f Isopar® G , M and V through one-sided preformed F303 paste when a 2 M P a pressure was applied for 30 s at 25°C.  Figure 5.5 shows the axial liquid migration o f the lubricants Isopar® G , M and V when the samples were preformed by applying pressure at both ends (two-sided preforming). Again, Isopar® V exhibits the highest lubricant concentration followed by Isopar® M . Isopar® G due to its relatively high volatility exhibits a lower amount compared to its initial concentration. A l l the lubricants show an even lube distribution in the middle. A s a result o f the two-sided preforming sample, the lubricant has migrated and accumulated in the middle. For lubricants Isopar® V and M , these variations are minor due to their high viscosities, which resist liquid motion. However, for Isopar® G these seem to be more significant. Therefore, it can be concluded that the higher the viscosity o f the lubricant is, the better is the quality o f perform. However, the high viscosity of the lubricant is expected to increase the extrusion pressure and this might have a deleterious effect on the quality o f the final products, i.e. influence the degree o f fibrillation (Ariawan et al., 2002b; Benbow et al., 1987; Mazur, 1995). In addition, the higher the viscosity o f the Isopar® is, the higher its surface tension is (see Table 1). This property has an effect on the wetting characteristics with P T F E , and as such it should be considered in selecting an appropriate lubricant.  59  20  19  Isopar G - b — Isopar M Isopar V  Initial lubricant concentration for Isopar M & Isopar V (18 w%)  14 20  40  60  80  100  120  140  D i s t a n c e f r o m b o t t o m (mm)  Figure 5.5:  A x i a l liquid migration for Isopar® G , M and V for a two-sided sample preformed F303 paste by applying a pressure o f 2 M P a for 30 s at 25°C.  5.3.2 Radial liquid  migration  Figure 5.6 shows the radial and axial liquid migration behavior for Isopar® G for one-sided preforming pastes. The higher concentrations o f lubricant are found near the bottom (opposite end from that where pressure is applied) when the paste is one-sided preformed (also seen before). Regarding the radial distribution, it can be seen that it is minimal. This behavior was also verified by the other lubricants examined, namely Isopar® M and V . Figure 5.7 depicts the radial and axial liquid migration for Isopar® G for a two-sided preformed sample. It can be easily seen that the profiles i n the axial directions are similar to those discussed with reference to Figure 5.5. The lube concentration is reduced near both ends and is increased i n the middle. In addition, the radial distribution is esentially uniform. Similar results were obtained for the other lubricants. Therefore, it can be concluded that the radial distribution is minor for all practical purposes.  60  Figure 5.6:  A x i a l and radial liquid distribution o f Isopar G for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 2 5 ° C .  Figure 5.7:  A x i a l and radial liquid distribution o f Isopar® G for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C. 61  Figures 5.8 to 5.11 are 3-D plots depicting the lubricant distribution in both the radial and axial direction for Isopar* M and V . In these figures it is possible to observe how the liquid concentration changes in both directions when the sample is one-sided preformed.  Figure 5.8:  A x i a l and radial liquid distribution o f Isopar M for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  Figure 5.9:  A x i a l and radial liquid distribution o f Isopar® M for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C. 62  O n the other hand, when the paste is two-sided preformed the lubricant accumulates in the middle o f the preform. In all cases, the concentration variations in the radial direction are minimal  Figure 5.11:  A x i a l and radial liquid distribution o f Isopar V for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C. 63  Perhaps a better way to see the liquid distribution within the preform is with the contour plots shown in Figures 5.12 to 5.17. Referring to Figures 5.12 and 5.13, it is possible to see how the lubricant with low viscosity (Isopar® G) distributes uniformly in the radial direction even when the pressure is applied just on one o f the two ends (one-sided). The radial distribution o f the lubricant is quite homogenous but the concentration o f the lubricant is lower at the top as a result o f the applied pressure on that point. When the paste is two-sided preformed,  the  distribution o f the lubricant is more even in both the radial and axial directions. Figures 5.14 and 5.15 are the corresponding contour plots for the axial and radial distribution o f Isopar® M within the paste for one-sided and two-sided preforms, respectively. In the latter, the lubricant is more evenly distributed but still accumulates in the middle. In Figures 5.16 and 5.17, it can be observed that the lubricant with the highest viscosity, Isopar® V , exhibits a higher resistance to flow and obviously the distribution is more homogeneous compared to other lubricants o f lower viscosity. The application o f pressure on both sides, Figure 5.17, helps to obtain a somewhat more even distribution o f the lubricant.  20  2  3  4  5  6  7  8  9  10  Distance from centre (mm)  Figure 5.12:  Contour plot o f axial and radial liquid distribution o f Isopar® G for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  64  I 2  , 3  , 4  , 5  , 6  7  8  9  10  Distance from centre (mm)  Figure 5.13:  Contour plot o f axial and radial liquid distribution of Isopar G for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  2  3  4  5  6  7  8  9  10  Distance from centre (mm)  Figure 5.14:  Contour plot o f axial and radial liquid distribution o f Isopar® M for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  65  2  3  4  5  6  7  8  9  10  Distance from centre (mm)  Figure 5.15:  Contour plot o f axial and radial liquid distribution o f Isopar M for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  2  3  4  5  6  7  8  9  10  Distance f r o m centre (mm)  Figure 5.16:  Contour plot o f axial and radial liquid distribution o f Isopar V for a one-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  66  2  3  4  5  6  7  8  9  10  Distance f r o m centre (mm)  Figure 5.17:  Contour plot o f axial and radial liquid distribution o f Isopar* V for a two-sided preformed sample by applying a pressure o f 2 M P a for 30 s at 25°C.  5.3.3. Effect of high preforming pressure and its duration on liquid  migration  Preforming was also studied under extreme conditions (high pressure and long preform times) in an attempt to determine possible effects o f the wettability o f lubricants on liquid migration at these extreme conditions. Figure 5.18 shows the effect o f high preforming pressure (IS (10 M P a ) applied for 10 minutes on liquid migration for Isopar  M , G , E and H F E 7500. It can  be seen that the application o f high pressure redistributes the liquid creating non-uniform distribution in all cases except Isopar® M (due to its relatively high viscosity). Surface tension also influences the liquid migration by changing the wetting characteristics with P T F E . For example, compare the behaviour o f H F E - 7 5 0 0 with that o f Isopar® G , which have similar viscosities. Having good wetting characteristics, the liquid HFE-7500 creates a thin coat around the P T F E particles. This coating cannot be displaced by the application o f a high pressure. O n the other hand, i f the wettability characteristics are poor, free liquid exists and this can be nonuniformly displaced to create nonuniformities within the sample.  67  1.00  t  •  t  r  r  1  "I  1  P = 10 MPa T = 25°C t = 10 min •  0.98 h  ~  0.96 k  _ O >  0.94  o >  0.92  0.90  — -•— -A— -+—1  Isopar M Isopar G Isopar E HFE 7500  I  20  _l_  L  40  _L  60  80  100  120  D i s t a n c e f r o m b o t t o m (mm)  Figure 5.18:  Effect o f preforming pressure and duration on the axial lubricant migration at 25°C on a paste prepared with F301 and 38.8 V % o f lubricant.  17 16 I15 c CO o XI 3  -5-  14 [•  - • — 50.47 wt% of A O T  13 12  •  29.11 wt% of A O T  -a  11.23 w t % o f A O T  fl 10  -L.  20  40  60  80  100  120  Distance from bottom (mm)  Figure 5.19:  A x i a l lubricant distribution in a one-sided preformed paste prepared with F301 and various Isopar® G - A O T solutions pastes preformed at 2 M P a for 30 s at 25°C. 68  To study further the effect o f wettability on liquid migration, different solutions o f Isopar® G with A O T were prepared. A s mentioned before, this surfactant does not change the total surface tension (although it increases its polar component) o f the lubricant, but it does change its viscosity, density and 3-phase contact angle. A s a result, the wettability o f the lubricant with P T F E changes accordingly (see Chapter 4). The concentration o f the solutions as well as their corresponding physical properties can be found in Table 4.3. Pastes with P T F E and these lubricants were prepared and preformed at a pressure o f 2 M P a for 30 seconds. A l l pastes were prepared with a lubricant concentration o f about 38.8 v o l % and the results are plotted in Figure 5.19. It can be clearly seen that the wettability o f these solutions have produced an excellent quality preform. It remains to be seen whether or not the examined viscosity and wettability effects on preform properties, are also reflected during extrusion (see Chapter 6).  5.4 Liquid migration during extrusion Although paste extrusion is investigated in detail i n the next chapter, it would be interesting to examine here liquid migration effects during a typical extrusion run and its consequences on the mechanical properties o f the extrudate. Figure 5.20 is a typical start up pressure transient plot obtained during P T F E paste extrusion. It has been divided into four zones in which the lubricant migration and extrudate mechanical properties w i l l be examined separately. The paste was prepared with F301 and 18.8 wt % ± 1.8% o f Isopar® M in the usual way. The paste was then extruded at a constant shear rate of 2812 s" through a tapered die with entrance angle o f 2a = 60°, L/D = 20 and RR = 352:1 at 1  35°C. Four similar experimental runs were carried out in order to check the repeatability. Fresh paste at the same concentration was prepared for each extrusion. The concentration o f the pastes was measured prior to preforming. During extrusion, samples from each marked zone were taken, dried and their lubricant concentration and tensile strengths measured. In addition, the paste remaining in the barrel was also retrieved and its lubricant concentration was also measured. Table 5.1 lists the results obtained for the lubricant concentration in each zone.  69  I  III  II  IV  Distance in the barrel  Figure 5.20:  Typical start up o f pressure transient during extrusion o f paste (F301 + 18 wt% of Isopar® M ) using a die having 2a = 60°, L / D = 20 and R R = 352:1 at y =2812 5"'and 3 5 C . 0  A  It can be seen from Table 5.1 that the lubricant concentrations in zones I to I V are about the same for all practical purposes. It is noted that zone V refers to the preform recovered from the barrel and this is usually a bit drier compared to the other zones, although the data does not substantiate this.  Table 5.1:  Lubi oncerltrat  a o u  u  L i q u i d concentration during four extrusion experiments at the conditions listed in Figure 5.20.  Test 1  Zone I 17. 8  Zone II 18.2  Zone III 18.6  Zone IV 18.4  Zone V 17.9  Test 2  18.2  18.7  19.2  18.6  18.2  Test 3  17.7  18.4  18.4  18.5  17.8  Test 4  17.7  18.5  18.6  19.2  18.3  Mean  17.9  18.5  18.7  18.7  18.1  Std Dev  0.3  0.2  0.3  0.4  0.2  70  The tensile strengths o f the extrudates were also measured and the results are listed in Table 5.2. The elastic modulus and the elongation at break are also included. The values correspond to the average value estimated from different specimens taken in each zone. The standard deviations are also included in parentheses.  Table 5.2: zone I II III IV  Mechanical properties o f the extrudates obtained in different flow zones. Tensile Strength (MPa) 3.3 (0.8) 2.6 (0.4) 2.7 (0.2) 2.8 (0.3)  Elastic Modulus (MPa) 64.7 (23.3) 52.2 (10.4) 53.3 (10.1) 60.1 (9.1)  Elongation at break (%) 117.8 (67.4) 140.4 (40.4) 136.3 (24.9) 131.8(20.8)  The results o f the tensile strength measurements are shown in Figures 5.21 to 5.24 for the four flow zones. M o s t o f the changes take place in zone I. Initially, as pressure builds, crystallites o f neighbouring particles mechanically interlock and thus fibrils are created. These give the good dimensional stability o f the extrudates. A s the pressure increases further, it seems that while fibrils might still be created, their extensibility is low. In the neighbourhood o f the peak pressure, the extensibility o f the samples decreases. In all the other zones the tensile strength and extensibility appear to be similar (see discussion below). Similar conclusions have been drawn by Ariawan (2002), who has reported that some fibrils may break at the point o f highest pressure. Figure 5.22 shows the stress vs. strain behaviour o f extrudates obtained during flow zone II. It can be seen that significant variations with respect to extensibility occur. A t the beginning o f zone II (after the pressure has reached its maximum) the extrudate is very weak. However, its quality improves when the extrusion pressure decreases toward its steady-state value.  71  5 I—•  i  i  •—i—'—i—i—i—i—i—i—i—i—i—i  i  '  i  i  E n d of Zone I  0  50  100  150  200  250  Strain (%)  Figure 5.21:  Stress vs. strain plot for extrudates obtained during flow zone I from Test 2 (see Table 5.2). 4 |  0  •  •  •  •  i  50  •  1  i  i  i  i  100  i  i  i  i  150  1  i  '  '  1  200  Strain (%) Figure 5.22:  Stress vs. Strain plot for extrudates obtained during flow zone II from Test 3 (see Table 5.2). 72  Figure 5.23 presents the results for extrudates obtained during zone III (the steady-state zone). It can be seen that the various extrudates show similar tensile strength and extensibility values. Regardless o f where the sample is taken from, it shows the same trend. Figure 5.24 depicts the stress-strain behaviour o f samples obtained during flow zone I V . This is the end o f the extrusion process where the pressure starts increasing slowly due to the fact that the paste becomes drier (Ariawan, 2002). It can be seen that the different parts o f the sample show consistently the same tensile strength and extensibility as those obtained at steady state. It can also be seen from Table 5.1 that the lubricant content i n the paste remaining in the barrel is not very different from the initial concentration and this is the reason that the pressure does not increase much towards the end o f the experiment.  Figure 5.23:  Stress vs. strain plot for extrudates obtained during flow zone III from Test 1 (see Table 5.2).  73  0  50  100  150  200  Strain (%) Figure 5.24:  Stress vs. Strain plot for extrudates obtained during flow zone I V from Test 4 (see Table 5.2).  5.5 Summary A number o f lubricants were used and examined as possible processing aids in the paste extrusion o f P T F E . They were characterized in terms o f both flow and surface properties. It was found that it is possible to alter flow and surface properties independently and thus it became possible to study their relative effects in preforming and extrusion. The following conclusions can be drawn so far from the present study: First,  the  preforming  pressure  and  duration  significantly affect the quality o f preform. Lack o f adequate pressure w i l l result in a preform o f non-uniform density which w i l l extrude unsteadily. Increasing the preforming time and pressure improves the uniformity o f preform density, indicating that the process o f preforming is time dependent. However, lubricant migration becomes important at longer times. Therefore, the applied pressure and its duration need to be optimized depending on the physical properties o f lubricant. A pressure o f 2 M P a applied for about 30 seconds seems to be optimum for most cases. Moreover, it has been shown that making a preform by compacting the paste on both ends improves the uniformity o f preform density without sacrificing its quality i n terms o f lubricant distribution.  74  The viscosity has a significant effect on producing a uniform preform. The use o f a lubricant o f high viscosity can produce a more uniform preform as liquid migration is minimized. For the lubricants used in this study a higher viscosity means also a lubricant o f lower vapour pressure and thus minimum liquid evaporation. Increasing the wettability o f lubricant with P T F E produces better mixture/pastes. Furthermore, this has an effect on the preform preparation. Excellent wetting would produce a uniform preform even under extreme conditions. The quality o f the extrudate was found to be very good during all moments o f an extrusion run except at the point where the maximum pressure is obtained.  75  CHAPTER 6 PTFE Paste Extrusion 6.1 Introduction Polytetrafluoroethylene (PTFE) cannot be processed by using melting techniques. Instead, it is processed by a technique known as paste extrusion, similar to that used in the processing o f ceramics (Ebnesajjad, 2000). The main difference is that when the P T F E particles are squeezed together during flow, fibrils are formed in the contraction zone o f the die. The degree o f fibrillation is closely related to the geometrical characteristics o f the die used for the extrusion. The physical properties o f the liquid used as a processing aid also play an important role in this process. A t atmospheric pressure conditions, P T F E has two transition temperatures at approximately 19°C and 30°C (see Figure 1.2 and Blanchet, 1997 for more details). Below 19°C, P T F E particles are very hard and shearing w i l l cause its crystals to slide past each other, retaining their identity. Above 19°C, P T F E molecules are packed more loosely and shearing w i l l cause the unwinding o f crystallites causing the creation o f fibrils (Ebnesajjad, 2000; Mazur, 1995). These fibrils connect the particles together, creating a network and giving dimensional stability to the extruded paste. A t temperatures greater than 30°C, a higher degree o f fibrillation can be achieved. This property has made it possible for P T F E paste extrusion to be performed near ambient temperature, producing a mechanically strong extrudate (Ebnesajjad, 2000). The present chapter studies the process o f P T F E paste extrusion under  different  operating conditions by using different die geometries and lubricants having different physical properties. First, it addresses the effects o f the lubricant physical properties on the extrusion pressure and the mechanical properties o f dried and sintered extrudates. Then, the effects o f the die geometry characteristics are assessed. The effect o f the temperature i n P T F E extrusion is also studied to a certain extent.  76  6.2 Pressure transient in PTFE extrusion Figure 6.1 depicts a typical pressure transient obtained during P T F E paste extrusion using a capillary rheometer. The horizontal axis labelled as distance in the barrel is equivalent to time once the constant piston speed is used to divide the distance. The curve has now been divided into three zones (Ariawan, 2001; M c L e o d , 1997). In zone I, the pressure gradually increases and goes through a maximum. This pressure peak has been thought to be due to the initial filling and wetting o f conical zone o f the die (Mazur, 1995; Ariawan, 2001; M c L e o d , 1997; Benbow & Bridgwater, 1987). Experiments performed by using dies having a 180° angle, have also shown the existence o f this peak. In other cases, the conical zone was filled before extrusion and the pressure peak still appeared (Ariawan, 2001; Ariawan et al., 2002b). Therefore, the presence o f the pressure peak is not due to the initial wetting and filling o f the die conical zone (Ariawan, 2001; Ariawan, et al., 2002b). The origin o f this maximum is discussed below.  i_ 3  (A (A <1>  Q_ C  o '55 3 l_  M  w  Distance in the barrel  Figure 6.1:  Typical start up o f pressure transient obtained in P T F E paste extrusion.  Zone II is taken to be the steady-state part o f the extrusion process. The recorded average pressure in this zone is reported as the extrusion pressure. Finally in zone III, the pressure gradually increases due to the fact that the final part o f the preform becomes drier due to liquid 77  migration. The network o f P T F E particles plays the role o f an apparently immobile screen. The net result o f this is that the lubricant is moving slightly faster than the assembly o f particles and therefore causes the last part o f the preform to become drier (lower lubricant concentration) and therefore to extrude at a higher pressure. The maximum in the extrusion pressure obtained in zone I during a typical pressure transient (defined as the yield pressure in Figure 6.1) deserves further discussion.  This  maximum is essentially due to the finite compressibility and the yield stress o f the paste in the barrel. Pastes in general are visco-elasto-plastic materials exhibiting a small but finite yield stress (Ebnesajjad, 2000; Ariawan et. al, 2000b; Snelling and Lontz, 1960). It is reasonable to argue that the existence o f a yield stress causes the appearance o f the yield pressure (peak pressure). Until this point is reached, the paste flows in the die at very low speed. The paste is been compressed in the barrel and as a result the static pressure increases gradually. During this compression period, the paste is in a state o f jamming, which is defined as the conversion of a liquid system into a solid by imposed stress (Haw, 2004). This essentially means that there is a number o f immobile clusters o f particles upstream o f the die entrance that is responsible for the jamming (Breedveld and Pine, 2003; Manoharan et al., 2003). Collapse o f these immobile clusters o f particles initiates the flow and this happens once the yield pressure is reached (Haw, 2004). Jamming and flow initiation at a yield point similar to that depicted in Figure 1, have been observed in the ram extrusion o f cold chocolate (Chen and Mackley, 2004). During this stage when the extrusion pressure is rising, the extrudate undergoes some distortion and in some cases severe fracture and breakage. After this point is reached, the flow experiences an acceleration and the pressure drops to its steady-state value. Further convincing evidence that the initial jamming is due to a few immobile clusters o f particles is provided in Figure 6.2. This figure depicts the extrusion o f paste prepared with F104 H M W and Isopar® M . During this test, the extrusion was stopped after reaching the steady-state region and the paste was allowed to relax for different periods o f time. For the small relaxation period o f 90 seconds, the yield pressure needed to be overcome in order to initiate the flow is very small compared to the initial yield pressure. This shows that the yield surfaces have only been recovered partially. A s the relaxation time increases the yield pressure needed to initiate the flow increases. Finally, for the large relaxation time o f 40 hours, the immobile clusters o f particles have been reinstated completely. In this case the extrusion pressure has to be increased to levels equal to the initial yield pressure in order to initiate the flow. While the yield pressure is a function o f relaxation time, the steady-state preassure is about the same. The small increase 78  in the steady-state preassure after the 40-hours relaxation time is mainly due to the partial evaporation o f the lubricant.  140 I  1  1  1  |  i  i  i  0 I 0  i  I 20  i  l 40  i  I 60  1  |  i  1  1  80  1  i  1  •  100  |  •  1  •  120  |  1  140  •  1  160  Distance (mm)  Figure 6.2:  Extrusion o f F104 H M W + 38.8 V % o f Isopar® M . The extrusion was stopped and restarted after (a) 1.5 minutes, (b) 10 minutes, (c) 45 minutes, (d) 40 hours.  It can also be argued that the yield pressure should depend on the contraction angle, the reduction ratio, and the type o f resin and lubricant and it should be independent o f the apparent shear rate. Figures 6.3 to 6.6 show typical pressure transients obtained by extruding various pastes (different lubricants and P T F E resins) using various die geometries and conditions in order to support these hypotheses. The yield pressure is independent o f the apparent shear rate, although the steady-state value scales with it (Figure 6.3). A statistical analysis o f the large body of data generated during this work supports this point o f view, whereas the results plotted in Figure 6.3 are only typical. Figure 6.4 shows that the yield pressure depends on the physical properties o f the lubricant. Viscosity and surface tension cause the immobile clusters o f particles to yield at different levels depending on the lubrication and wettability properties o f the PTFE/liquid interfaces. The yield pressure also depends on the type o f resin (Figure 6.5). Different P T F E molecular structures result in particles o f various sizes (primary and secondary) as well as particles having different crystallinity, hardness and roughness. A s pressure increases, the 79  particles are squeezed against each other and various levels o f mechanical interlocking between crystallites across the contact interface between particles occurs. This influences the nature and strength o f the immobile clusters and therefore the yield pressure at which flow is initiated. Note that resins F104 H M W and F104 L M W are similar, and thus exhibits the same yield and steadystate pressures. Figure 6.6 shows that the yield pressure is a function o f the reduction ratio (compare curves 3 and 4) and a function o f the contraction angle (compare curves 1 and 2). It is noted that the results plotted in Figures 6.3 to 6.6 are only representative results and the conclusions drawn are based on the large body o f experimental data generated during the course o f this study. Finally it should be stressed that the experimental observations discussed in this section can be put into the right perspective, i f one simulates a simple compressible viscoplastic fluid in capillary flow (including the conical zone into the simulation) to see whether or not a yield pressure is predicted. In addition it would be important to check through simulation whether or not the predicted yield pressure depends on L / D ratio, reduction ratio and contraction angle o f the die as well as whether or not it is independent o f the apparent shear rate. If these are true assumptions, then inverse simulation might be a good way o f determining the yield stress o f these types o f fluids in transient capillary flow. 80 I  i  1  i—i  i—i  i—i  •  1  •  r  Distance (mm)  Figure 6.3:  Pressure transients during extrusion o f P T F E paste prepared with F104 H M W and Isopar G at different apparent shear rate values.  80  90 I  1  1  1  1  1  1  1——i  1  1—•  r  80 J-  0  20  40  60  80  100  120  140  Distance (mm)  Figure 6.4:  Pressure transients during extrusion o f P T F E paste prepared with F104 H M W and different lubricants at different apparent shear rate values..  70 I — i — i — i — i — • — i — i — i — ' — i — i — i — • — r  0  20  40  60  80  100  120  140  160  Distance (mm)  Figure 6.5:  Pressure transients during extrusion o f P T F E paste prepared with different resins and Isopar G .  81  0  20  40  60  80  100  120  140  Distance (mm)  Figure 6.6:  Pressure transients during extrusion o f P T F E paste prepared with F104 H M W and Isopar G at different conditions and dies.  6.3 The effect of the physical properties of lubricants on PTFE paste extrusion 6.3.1 The effect of surface tension Figures 6.7 and 6.8 depict the effect o f lubricant surface tension and apparent shear rate on the extrusion pressure o f two different P T F E pastes. First, Figure 6.7 compares the behavior of Isopar® G + 11.2 wt% A O T with that o f HFE-7500 in the paste extrusion o f resin F104 H M W . The viscosities o f these lubricants at 25°C are equal to 1.24 mPa s. However, the H F E 7500 has a surface tension o f 16.2 x 10" N / m compared to 23.6 x 10" N / m o f Isopar® G + 11.2 3  3  wt% A O T . Given that the surface tension o f P T F E is about 18 x 10~ N / m , the wettability o f 3  HFE-7500 for P T F E should be complete. The contact angle o f H F E - 7 5 0 0 with P T F E has been reported to be 0°, whereas that o f Isopar® G + 11.2 wt% A O T with P T F E about 40° (Ochoa and Hatzikiriakos, 2004). It can be seen from Figure 6.7 that the enhanced wettability o f HFE-7500 with P T F E decreases significantly the extrusion pressure.  82  F104 HMW + 38.8 vol% Lubricant, T = 35 °C, 2a = 30°, L/D = 20, RR = 352:1 20 ^ * * ' ' * ' ' * ' ' * * * ' ' * * ' * * * * ' * * * * 1000 2000 3000 4000 5000 6000 7000 1  1  1  Apparent Shear Rate (s' ) 1  Figure 6.7:  The effect o f lubricant wettability on the extrusion pressure o f paste prepared with resin F-104 H M W and two different lubricants having about the same viscosity and different wettability properties at 35°C.  Figure 6.8 presents a similar case. It compares the behaviour o f Isopar® G + 50.5 wt% A O T with Isopar® V in the paste extrusion o f resin F301. The viscosities o f the two lubricants at 25°C are equal to 7.5 mPa s. However, the solution Isopar® G + 50.5 wt% A O T has a surface tension o f 23.6 x 10" N / m compared to 30.8 x 10~ N / m o f Isopar® V . The corresponding 3  3  contact angles o f these two lubricants with P T F E are about 56° and 60° respectively (Ochoa and Hatzikiriakos, 2004). The effect o f wettability (surface tension) on the steady-state extrusion is clear again. Better wettability decreases the extrusion pressure. It is also noted that the actual amount o f liquid Isopar® G in the Isopar® G - A O T mixture is only 11.5 wt% (the rest is A O T ) compared to about 19.2wt% o f Isopar® V . Using such a small amount o f lubricant, the pressure was expected to be extremely high (Horrobin, 1998). However, the effect o f wettability is dominant, keeping the pressure at low levels.  83  80  Q.  70 Isopar V  -JEl_  3  (A (A 0)  60  T  L-  0. c o "35 3 l_  !  v— 50  •5:-  Isopar G + 50,5 wt% AOT  F-301 + 38.8 vol% Lubricant,  "R  T = 35 °C, L/D = 0, 2a= 90°, R=352:1  UJ 40 i—'—>-  1000  2000  I i i 3000  4000  5000  i i i l i 6000  . 7000  Apparent Shear Rate (s~)  Figure 6.8:  The effect o f lubricant wettability on the extrusion pressure o f paste prepared with resin F-104 H M W and two different lubricants having about the same viscosity and different wettability properties at 35°C.  Figures 6.9 and 6.10 plot the tensile strength o f the extrudates obtained from the extrusion experiments plotted i n Figures 6.7 and 6.8. It can be seen that the tensile strength increases with a decrease in the extrusion pressure, an observation that is counterintuitive. The tensile strength o f the extrudates in the extrusion direction is due to the presence o f fibrils (Ebnesajjad, 2000; Mazur, 1995). Fibrillation occurs in the conical zone o f the die, where the particles are mechanically interlocked under the action o f high pressure. A s particles are accelerated into the conical zone, due to the nature o f the flow (extensionally dominated flow), mechanically locked crystallites across the area o f contact between particles are unwound and create fibrils. These fibrils give dimensional stability to the extrudates (Ariawan et al., 2002b). A higher pressure usually increases the degree o f fibrillation and therefore the tensile strength o f the extrudates (Ariawan et al., 2002). In the present case, the opposite is happening and this is certainly due to wettability. It seems that the enhanced wettability distributes the lubricant uniformly within the paste. This minimizes frictional effects. It is also apparent that the formation o f fibrils under minimized frictional effects makes them more stable. Although an excessive pressure due to a lack o f lubrication causes a significant degree o f fibrillation, many of these fibrils break (Ariawan et al., 2002b).  84  It can also be seen from Figures 6.9 and 6.10 that the tensile strength decreases with increase o f apparent shear rate. While excessive pressure might be the origin o f this decrease (note that pressure increases with apparent shear rate), the H e n c k y strain and strain rate undergone by the fibrils at the conical zone is the dominant mechanism for this. The total Hencky strain remains the same as this depends on the reduction ratio o f the die. For example, the total Hencky strain, s , H  experienced by a fluid element moving along the centreline o f the  contraction region from far upstream in the barrel to the die exit is given by (Rothstein, 1999) (z=L)  Vl  :  H  =  dt=  | ^ Vj(-oo)  = In  Db  = \n[RR]  D  (6.1)  where t is the time spent in the centreline, Dt/D is the barrel diameter to the die diameter ratio 0  and RR is the reduction ratio defined as (Dt/D) . However, a higher apparent shear rate 2  corresponds to a faster extensional rate (Hencky strain rate) at the entrance. Reaching the same extensional (Hencky) strain at a higher extensional (Hencky) strain rate causes breakage o f some o f the fibrils. A s a result, this causes a decrease in the tensile strength.  X _  re  HFE-7500 •  Q_ •  O)  c  +->  (/)  Isopar G + 11.2 wt% AOT  -  < ' /> c  F104 HMW + 38.8 vol% Lubricant, T = 35 °C, 2a = 30°, L/D = 20, RR = 352:1  1000  2000  3000  4000  5000  6000  7000  Apparent Shear Rate (s )  Figure 6.9:  Tensile strength o f extrudates obtained from paste prepared with resin F104 H M W and two different lubricants having about the same viscosity and different wettability properties with F104 H M W at 35°C.  85  1.2 1.0 h  re Q.  U)  c  <> l  Isopar G + 50.5 w t% A O T -T-  0.8  V7  0.6  Isopar V  —-w  w 0.4 '35 c  F-301 + 38.8 vol% Lubricant 0.2  T = 35 °C, L/D = 0, 2a= 90°, R=35i2:1 I I I I I I I  u.u 1000  2000  3000  4000  5000  6000  7000  Apparent Shear Rate (s ) 1  Figure 6.10:  Tensile strength o f extrudates obtained from paste prepared with resin F-301 and two different lubricants having about the same viscosity and different wettability properties with F301 at 35°C.  6.3.2 The effect of viscosity Figure 6.11 shows the effect o f the lubricant viscosity on the steady-state extrusion pressure o f P T F E pastes prepared with the homopolymer F104 H M W and lubricants having similar surface tensions but significantly different viscosities. It can be seen that the extrusion pressure generally increases with increase o f the lubricant viscosity and apparent shear rate. The higher pressure is due to the high viscosity o f the lubricants increasing the resistance to flow. The lubrication resistance between particles in contact and at the paste/wall interface also increases with lubricant viscosity. In a similar way, one may argue that the effective viscosity o f the mixture increases with lubricant viscosity and hence a higher extrusion pressure is needed to cause the paste to flow. Although results are presented for only one P T F E resin, a similar behaviour was found for the pastes prepared with all the other P T F E resins.  86  70  i i i i I i i i i I i i i i I i i i i I i i • • I i i • • F104 HMW + 38.8 vol% Lubricant,  65  T = 35 °C, 2a = 30°, L/D = 20, RR = 352:1 Isopar G + 50.5 wt% A O T  Q.  Ld>  i  60  i_  Isopar G + 11.2 wt% A O T  3  .  V>  S  55  O  50  '55  I  3  ^< 45  40 1000  i i i i i i i i  1  2000  3000  4000  5000  6000  7000  Apparent Shear Rate (s")  Figure 6.11:  Effect o f lubricant viscosity on extrusion pressure o f F104 H M W .  Figure 6.12 depicts the tensile strengths o f the dried extudates obtained from the extrusion experiments plotted in Figure 6.11. The tensile strengths not only exhibit a decrease with an increase o f the shear rate, but also with an increase i n viscosity due to the higher pressures encountered as the vicosity is raised which possibly cause breakage o f fibrils (discussed above).  10  1• • • • i • • • • i • • • • i • • • • i • • • ' •  F104 HMW + 38.8 v o l % Lubricant,  J  T = 35 °C, 2a = 30°, L/D = 20, RR = 352:1  J  9  re  U)  c  0 w 0)  '55 c  8  _  7  -  6  -  5Isopar G  •  5  A •  —-a*.  4 3  -  2  -  Isopar G + 50.5 wt% A O T "  1 1000  m  Isopar G + 11.2 wt% A O T J  Ki» -™-  2000  _ 1 •  3000  4000  _  _  mr  m-  5000  6000  ~  7000  A p p a r e n t S h e a r Rate (s")  Figure 6.12:  Tensile strength o f dried extrudates o f pastes prepared with F-104 H M W and different lubricants having different viscosities and about the same wettability properties. 87  6.4 The effect of geometrical characteristics of die on extrusion pressure 6.4.1 Reduction  ratio  Figure 6.13 depicts the effect o f the reduction ratio o f the die on the extrusion pressure of pastes prepared with resin F 1 0 4 H M W and various lubricants. The extrusions were performed at two different shear rates (375 s" and 1736 s" ) for each reduction ratio o f 56, 156 and 352. 1  1  The closed symbols correspond to the apparent shear rate o f 375 s" and the open symbols to that 1  of 1736 s" . It can be seen that the extrusion pressure increases with an increase in the reduction 1  ratio in a nonlinear fashion. The corresponding tensile strengths o f both dried and sintered extrudates are depicted in Figure 6.14. The volumetric flow rate has kept constant and the corresponding shear rates are indicated in the plot. First, it is noted that the tensile strength o f the sintered samples are about 8 to 10 times higher than those o f the dried ones. However, they both follow the same trend. It appears that there is an optimum reduction ratio that maximizes the tensile strength o f the extrudates i.e. at a RR around 150 the tensile strength o f the extrudates appears to obtain the highest value. Although the results correspond to different apparent shear rate values, it is noted that the effect o f apparent shear rate on the extrusion pressure is much weaker than that o f the reduction ratio. 70  • i ••••i '  -%  F104 H M W + 38.8 vol% Lubricant 60 T T = 35 °C, 2a = 60°, L/D = 20 CO Q.  0i 1_  3 (A (A CD  l_  Q_ C  •  50 40  -  30  o '(75  20  "R  10  j =1736 s"' A  — - • — -  3  LU  :  0  A  50  100  150  200  250  Isopar G :  o  Isopar M -  A  HFE-7500'  300  350  400  Die Reduction Ratio  Figure 6.13:  Effect o f die reduction ratio on the extrusion pressure o f F104 H M W resin.  88  50  i i i i i i i i i i i i i i i i i i i i i  I I I  i i i i i i  q *•  0  •  50  100  150  200  250  300  "•  350  400  Die Reduction Ratio  Figure 6.14:  Effect of die reduction ratio on the tensile strength of dried (lower plot) and sintered (upper plot) extrudates obtained from paste prepared with homopolymer F104 H M W and various lubricants at 35°C.  In general, it can be concluded that use o f extrusion dies having high reduction ratios, causes a significant increase in the extrusion pressure. This causes significant fibril breakages that result in weaker extrudates. Similar observations have been reported by Ariawan (2001) and Ariawan et al. (2002b).  6.4.2 Length-to-diameter  ratio (L/D)  Figure 6.15 illustrates the effect o f the length-to-diameter ratio o f the die, L/D, on the extrusion pressure o f pastes prepared with F301 and several lubricants. A s expected, the extrusion pressure increases linearly with increase o f the L/D ratio. Note that most o f the resistance to flow is due to the conical zone. The pressure needed to extrude the polymer through the conical zone is significantly higher (intercepts on the vertical axis) compared to that needed to extrude it through straight sections o f the capillary die (Bridgwater, 1998; Ariawan, 2001). In addition, it is noted that the smallest extrusion pressures were obtained with lubricant HFE-7500. This lubricant has the smallest surface tension and hence is expected to completely wet the P T F E particles (0° contact angle).  89  130 120  i  r  F301 + 38.8 vol% Lubricant T = 35 °C, f  A  =5860 s"\ 2a = 90°, RR = 352:1  110  re  Q_  100  L3 U) (/> 0)  90 '80  - • — Isopar G -•  Isopar M  A  Isopar V  <•  -A—HFE-7500  70 o "35  60  "R  50  3 — I  UJ  40 30  10  15  20  25  30  35  40  45  L/D  Figure 6.15:  Effect o f length-to-diameter ratio (L/D) on extrusion pressure o f pastes prepared with F301 and various lubricants at 35°C.  Figure 6.16 shows the effect o f L/D ratio on the tensile strengths o f dried and sintered extrudates. It can be seen that the tensile strength o f the extrudate increases with the increase o f the L/D ratio up to certain value (L/D = 20) after which there is a slight fall. Short capillaries cause excessive die swell due to unrelaxed high first normal stress effects. These weaken the mechanical properties o f the extrudates significantly. A certain length o f capillary helps fibrils to relax and orient themselves in the extrusion direction. This also causes the first normal stress to relax and therefore the die swell remains at relatively low levels. Further increase o f the capillary length w i l l cause an increase in the extrusion pressure. This has an adverse effect on the tensile strength as high enough pressure cause fibril breakage. Therefore, it appears that an L/D ratio o f about 20 optimizes the mechanical properties o f extrudates. Similar observations have been reported by Ariawan (2001) and Ariawan et al. (2002b).  90  0  5  10  15  20  25  30  35  40  45  L/D  Figure 6.16:  Effect o f L/D ratio on the tensile strength o f dried (lower plot) and sintered (upper plot) extrudates o f paste prepared with copolymerF301 and different lubricants.  6.4.3. Entrance angle The effect o f die entrance angle on the extrusion pressure o f pastes prepared with P T F E F301 and various lubricants is plotted in Figure 6.17. The same trend was found using all other three P T F E resins. It can be seen that the extrusion pressure initially decreases and subsequently increases with increase o f the die entrance angle. The decrease in the extrusion pressure at small entrance angle values is similar to the trend predicted for polymer melts and other viscous liquids using the lubrication approximation (Dealy and Wissbrun, 1990). A theoretical derivation similar to that employed in the lubrication approximation, has been used by Benbow and Bridgwater (1993) in their modeling work on paste flow. When the die entrance angle is sufficiently small, paste flow in the die conical zone follows essentially a plug flow pattern for the most part similar to flow in the barrel (Benbow and Bridgwater, 1993). The lubrication approximation approach is valid here. However, the lubrication approximation predicts a monotonic decrease in the extrusion pressure with increasing die entrance angle, which is not consistent with the experimental results plotted in Figure 6.17. Beyond a certain entrance angle, the extrusion pressure increases with increasing angle, as is commonly observed with the extrusion o f elastic solids (for example, see Horrobin and Nedderman (1998) and the references 91  therein). A s fibrils are created, the paste attains significant elastic extensional properties that cause a significant increase in the extrusion pressure at high contraction angles. Figure 6.18 illustrates the effect o f entrance angle on the tensile strengths o f the dried and sintered extrudates. It can be seen that the tensile strength initially decreases with increase o f the entrance angle up to about 60° and then increases slightly. In the case o f sintered extrudates it can be argued that a similar trend exists although o f a milder degree. A small decrease in the tensile strength with increase o f the contraction ratio is expected. A s discussed before, in all extrusion cases plotted in Figures 6.17 and 6.18, dies having the same reduction ratio were used ( R R = 352:1). Therefore, the total extensional (Hencky) strain undergone by the pastes was the same. Increase o f the contraction angle corresponds to an increase o f the extensional (Hencky) strain rate without increasing the apparent shear rate. Note that the latter depends on the final die diameter. A s argued above, the high extensional ratio causes breakage o f fibrils resulting in slightly weaker extrudates.  110 100 90 Q.  i_ 3  (/> (S)  80 -  I——i1  - • — Isopar G •  Isopar M  a  Isopar V  •  HFE 7500  70 60 50 -  C o '35 3  k.  •s  UJ  40 30 20  F301 + 38.8 voI% Lubricant T = 35 °C, J A  =  5  8  6  0  >  s_1  L  /  D  = 20, RR = 352:1  10 0  I i 10 20  1  30  40  50  60  70  80  90  100  Die Entrance Angle (2a)  Figure 6.17:  Effect o f die entrance angle on extrusion pressure o f F301 resin.  92  60  i  i  i  F301 + 38.8 vol% Lubricant T = 35 °C, y =5860 s" , L/D = 20, R R = 352:1  50  1  A  re Q.  40  ±= 30 o> c /  co o  4  '55 c  3 2  [ • • • ^—A—  Isopar Isopar M Isopar V  ^ *»«^0 =a  •  Hr-.HF.E-75ap 10  20  r  •  30  40  50  60  •  70  80  90  100  Die E n t r a n c e A n g l e (2a)  Figure 6.18:  Effect o f die entrance angle on tensile strength o f dried (lower plot) and sintered (upper plot) extrudates obtained from pastes prepared with copolymer F301 and different lubricants at 35°C.  6.4.4 Molecular structure of the resin Figure 6.19 shows the extrusion pressure as a function o f the apparent shear rate for pastes prepared with all four P T F E resins and Isopar® G . These have all been extruded under the same conditions by using the same die and lubricant. It can be seen that both homopolymers extrude at higher pressure compared to their copolymers counterparts. The increase in the extrusion pressure with the resin molecular weight is caused by the hardness o f the particles o f the higher molecular weight resins. In general, the higher the molecular weight, the higher the extrusion pressure is. Copolymers consist o f smaller and softer particles (primary particles) that might be the origin o f the lower extrusion pressure. Contact angle measurements could not distinguish any wettability differences o f lubricants with homopolymers and copolymers. Therefore,  these effects  probably depend  on the  hardness and  size o f particles.  The  corresponding tensile strength results o f sintered and dried samples are depicted in Figure 6.20. It can be seen that the homopolymers exhibit higher tensile strength values. Local roughness on the particle surface might also influence the mechanical interlocking o f particles which might lead to an increased amount o f fibrillation. 93  70  i i i i I i i i i I i i i i I i i i i I i i P T F E + 38.8 v o l % of Isopar G T = 35 °C, 2a = 30°, L/D = 20, RR = 352:1  60  re CL  F104-LMW 03  U  "I  50  3 M W 03 •_ CL C  F104-HMW  F301  -yfc  40  o •(/)  F303  3  30  X LU  20 I i i i i I i i 1000 2000  •  1  4000  3000  5000  *  *  7000  6000  Apparent Shear Rate (s )  Figure 6.19:  Effect o f the molecular structure o f the resin on the extrusion pressure.  60  i i i i  I  i i i i  l  i i i i I i i i i I i i i i I • i i i  PTFE + 38.8 vol% of Isopar G, Unsintered 50  re CL  T = 35 °C, 2a = 30, L/D = 20, RR = 352:1  *  40  * -  *  — • — PI 04 LMW _ .„ . . —• F104 HMW * F  :  3  0  1  30  B  c  03  .*-» w 03  '55 c  8 6 4 2 •  1000  *  •  1  •  2000  •  3000  4000  5000  6000  7000  Apparent Shear Rate (s )  Figure 6.20:  Effect o f the molecular structure o f the resin on the tensile strength o f dried (lower plot) and sintered (upper plot) extrudates.  94  6.5 Effect of temperature on PTFE paste extrusion To study the effect o f temperature on the P T F E extrusion process, paste with F104 H M W and Isopar M at 38.8 v o l % was prepared. Isopar M was selected because it has a boiling point o f 224° C (vapour pressure o f 3.1 m m H g at 38°C) which makes it suitable for processing at high temperatures (minimal evaporation). The extrusions were carried out i n the regular way by using the Instron capillary rheometer. In order to extrude at temperatures below room temperature, copper tubing was wound around the barrel and cold water was circulated by using a refrigerator-bath circulator. Different shear rates varying from 1875 s" to 5859 s" were 1  1  attained by changing the plunger speed (from 0.34 to 1.06 mm/s) and using a cylindrical die with an entrance angle 2 a o f 30°, L/D ratio o f 20 and a RR o f 352. The extrudates were then dried and the tensile strengths were measured. Figure 6.21 depicts the extrusion pressure o f the paste at different temperatures. A t low temperature, the extrusion pressure was very low mainly due to limited fibril formation. P T F E experiences its first transition temperature at 19°C. Below this temperature, the P T F E particles are strong enough to resist any deformation and therefore few fibrils are formed. Since the viscosity and surface tension are functions o f temperature, the lubricant at low temperature can neither wet the P T F E particles properly nor move freely within the paste and its distribution is expected to be very poor. Thus, even though the extrusion pressure is very low, the extrudates appear very weak as is confirmed by their low tensile strengths. A s the temperature increases, the extrusion pressure also increases and the appearance o f the extrudate improves. However, at temperatures beyond 45°C, the extrusion pressure experiences a sudden decrement and there is no significant difference in the extrusion pressure at 55°C and 65°C. Figure 6.22 shows the tensile strengths  o f the extrudates  obtained at different  temperatures. The plots follow the usual behaviour found previously: "The higher the shear rate, the lower the tensile strength." However, it can also be seen that the tensile strength increases with an increase in the temperature. There is no significant difference in the tensile strength o f the extrudates obtained at 45°C and 55°C, in fact the tensile strength at 55°C seems to be lower than at 45°C. A t 65°C the tensile strength experiences an increase.  95  65  60 CL  55 ri_  3 <f> 1/5 50 h 0> i_ CL C  2 o 45 h '(/)  x UJ  F104 H M W + 38.8 v o l % Isopar M  40 r-  2a = 30°, R R = 352, L/D = 20 35 ' ' • • • * • ' • ' 1000 2000 3000  4000  5000  7000  6000  Apparent Shear Rate (s )  Figure 6.21:  Effect o f temperature on extrusion pressure.  16  I i i i i F104 HMW + 38.8 v o l % Isopar M  14  2a = 30°, RR = 352, L/D = 20  n  12  - • - T = 15°C •  T = 18"C  A  T = 22°C  - T - T • 35°C T = 4S°C  •L  +  10  T = SS'C  +-*  u> c m CO  8 6 4 2 _L  1000  2000  3000  4000  5000  6000  7000  A p p a r e n t S h e a r Rate (s")  F i g u r e 6.22:  Effect o f temperature on tensile strength.  Figure 6.23 and 6.24 are contour plots which allows a more comprehensive visualization of the effects o f temperature on extrusion pressure and tensile strength, respectively.  96  1875 2375 2875  3375 3875 4375 4875  5375 5875  Apparent Shear Rate (1/s)  Figure 6.23:  Effect o f temperature on extrusion pressure. Contour plot o f the results plotted in Figure 6.21.  1875 2375  2875  3375  3875 4375 4875  5375  5875  Apparent Shear Rate (1/s)  F i g u r e 6.24:  Effect o f temperature on tensile strength. Contour plot o f the results plotted in Figure 6.22. 97  6.6 Appearance of PTFE extrudates S E M micrograph pictures o f P T F E extrudates obtained at different conditions are examined below. These pictures are not intended to quantify the degree o f fibrillation during P T F E paste extrusion. However, it is possible to draw conclusions by making observations about the quality o f the fibrils. The latter may be correlated with the mechanical properties discussed above. Figures 6.25 to 6.28 are S E M micrograph o f extrudates obtained at y  = 5859 s" using 1  A  dies having different entrance angles (2a = 15, 30, 60 and 90°) and the same L/D = 20 and RR = 352:1. The paste used was F104 H M W mixed with 38.8 V % o f Isopar® M . It is possible to observe differences in the degree o f fibrillation as a function o f the entrance angle. The highest tensile strength o f 4.39 M P a is obtained at 2a = 15° and this is reflected in the thickness and density o f the fibrils. The excessive pressures at 2a = 60° and 90° make the fibrils weak and perhaps many o f these break due to the high Hencky strain rate that occurs for the dies having higher contraction angles.  Figure 6.25  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 15°, L / D - 20, R R = 352, y  = 5859 s" . Extrusion Pressure: 59.3 M P a . Tensile Strength: 4.4 M P a . 1  A  98  Figure 6.26:  F104 H M W + Isopar® M at 38.8 v o l % . 2a = 30°, L / D = 20, R R = 352, y = 5859 s" . Extrusion Pressure: 53.8 M P a . Tensile Strength: 3.9 M P a . . 1  A  99  Figure 6.28:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = 20, R R = 352, y = 5859 s" . Extrusion Pressure: 81.0 M P a . Tensile Strength: 3.2 M P a . 1  A  Figures 6.29 to 6.31 show S E M pictures o f extrudates obtained with dies having different L/D ratios from 0 to 40 at y  = 5859 s" and 35°C. The case o f L/D = 20 is shown in 1  A  Figure 6.28. The thickest fibrils in Figure 6.30 are deformed fibrils which were formed during sample preparation and should not be considered as an effect o f die geometry. In all cases, it is possible to see the fibrils connecting the P T F E particles. More specifically, at L/D o f 0 and 10, empty spaces can be seen among the particles which make the extrudates weaker. A t L/D - 20 (Figure 6.28) the size o f these spaces have been reduced and it is possible to observe that the fibrils connect the particles in both the axial and transverse directions. A t L/D = 40, the amount of empty space is minimal, but less fibrils are formed due to the high extrusion pressure.  100  Figure 6.29:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = 0, R R = 352, y = 5859 s" . Extrusion Pressure: 73.0 M P a . Tensile Strength: 0.7 M P a . 1  A  Figure 6.30:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = y  A  10,  R R = 352,  = 5859 s" . Extrusion Pressure: 77.6 M P a . Tensile Strength: 3.2 M P a . 1  101  Figure 6.31:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 90°, L / D = 40, R R = 352, y = 5859 s" . Extrusion Pressure: 93.0 M P a . Tensile Strength: 2.8 M P a . 1  A  Figures 6.32, 6.33 and 6.34 are S E M pictures o f extrudates processed with dies having different reduction ratios at y  = 5859 s" . It can be easily seen that the degree o f fibrillation 1  A  increases with reduction ratio causing an increase in the tensile strength.  Figure 6.32:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a - 60°, L / D = 20, R R = 352, y = 5859 s" . Extrusion Pressure: 73.8 M P a . Tensile Strength: 2.3 M P a . 1  A  102  Figure 6.33:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 60°, L / D = 20, RR = y = 5859 s" . Extrusion Pressure: 24.9 M P a . Tensile Strength: 2.4 M P a .  156,  1  A  Figure 6.34:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 60°, L / D = 20, R R = f = 5859 s" . Extrusion Pressure: 9.8 M P a . Tensile Strength: 1.2 M P a . 1  103  56,  Figures 6.35 to 6.38 are S E M pictures o f selected extrudates obtained from pastes prepared with F104 H M W and 38.8 v o l % Isopar® M and extruded at y = 5859 s" with a die 1  A  having 2a = 30°, L/D = 20, and RR = 352. The extrusions were performed at temperatures varying from 15°C to 65°C. The S E M micrographs show the effect o f temperature on P T F E extrudate appearance. Referring to Figure 6.22, when the extrusion temperature is increased from 15°C to 65°C, the extrusion pressure initially increases, before it starts decreasing at about 45°C. The tensile strength, however, shows a monotonic increase with increase o f temperature. It is possible to observe from Figures 6.35 to 6.38 that the quality and quantity o f fibrils also increase with temperature. The fibrils connecting the particles are becoming thicker and longer and the density o f the sample seems to increase since the empty spaces exhibited by the former samples are reduced and filled with interconnected particles.  104  Figure 6.36:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 30°, L / D = 20, R R = 352, y  = 5859 s" , T = 22°C. Extrusion Pressure: 55.9 M P a . Tensile Strength: 3.6 1  A  MPa.  Figure 6.37:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a = 30°, L / D = 20, R R = 352, y  = 5859 s" , T = 35°C. Extrusion Pressure: 53.8 M P a . Tensile Strength: 3.9 1  A  MPa..  105  Figure 6.38:  F104 H M W + Isopar® M at 38.8 v o l % . 2 a 352,  = 30°, L / D = 20, R R =  = 5859s" , T = 65°C. Extrusion Pressure: 53.2 M P a . Tensile Strength: 1  7.0 M P a .  6.7 Mechanical properties of PTFE extrudates Tensile strength is not the only mechanical property that should be taken into account when selecting a material for a certain application. Depending on the final application o f the product, there are other mechanical properties such as the elastic modulus, extensibility, modulus o f resilience and modulus o f toughness, which define a material as hard, tough, brittle, ductile, etc. A l l these properties are functions o f the conditions at which the extrudates are produced as well as functions o f the die geometry and physical properties o f the lubricants used for their extrusion. Table 6.1 lists some o f the mechanical properties o f the extrudates obtained for the various pastes extruded under different conditions using different die geometries and lubricants. The first column lists the tensile strength values already discussed above. The other two columns are the elastic modulus and the % elongation at break which are related to the elastic and plastic properties, repsectively, o f the material. A l l these properties have been defined in section 2.6 and are useful in classifying the materials as hard, soft, tough, brittle, and so on (Carswell, 1944). It can be seen that the homopolymers exhibit the highest elastic modulus.  106  These also possess a higher extensibility. Regarding the copolymers, F301 has a higher elastic modulus and extensibility compared to F303. A similar classification was concluded previously in terms o f tensile strength o f extrudates obtained under various conditions. The homopolymers always exhibited a higher tensile strength compared to the copolymers. The physical properties o f lubricants also affect the mechanical properties o f the extrudate. Lubricants having a lower surface tension (good wettability with P T F E ) , and lower viscosity increase the tensile strength and the elastic modulus o f the extrudates. The effect o f apparent shear rate on the mechanical properties seems to be negligible in general. Finally, temperature seems to have a strong effect on the mechanical properties. A s temperature increases, tensile strength, elastic modulus, and elongations at break increase significantly (see Table 6.1).  107  Table 6.1:  Mechanical properties o f P T F E resins extruded under different conditions.  Resin/Condition  Tensile Strength (MPa)  % Elongation at Break  Elastic Modulus (MPa)  Resin+ Isopar M , L/D =20, la = 30, ^ = 5859 s\RR  = 352:1  100.3 (13.9) 89.6 (3.7) 5.4 (0.1) F104 H M W 100.1 (8.8) 82.6 (3.9) 5.7(0.1) F104 L M W 62.7(15.2) 77.6(11.2) 4.4 (0.2) F301 38.8 (5.2) 48.6 (4.0) 3.6 (0.2) F303 r Lubricant, L/D =20, 2a = 30, RR = 352:1, ^ = 5859 s F104 H M W 90.5 (5.4) 97.0(5.5) 5.6 (0.2) Isopar® G 100.3 (13.9) 89.6 (3.7) 5.4 (0.1) Isopar® M 162.6 (13.2) 78.3 (3.3) 4.5 (0.2) Isopar® V 81.9(8.6) 160.4 (58.5) 5.6(1.3) HFE-7500 F104 H M W + Isopar® M , L/D =20, 2a = 60, y = 5859 s" 1  1  A  3.9 (0.2) RR = 352:1 2.3 (0.1) RR = 156:1 1.2 (0.1) RR = 56:1 =20, F104 H M W + Isopar® M , L/D = 2« 2a 2« la  = = = =  15 30 60 90 F104 H M W  5.0 (0.1) 3.8 (0.1) 2.3 (0.1) 3.2(0.1)  131.9(18.2) 347.3 (36.2) 373.5 (57.6)  82.5 (6.5) 51.2 (3.8) 42.2 (2.4)  RR = 352:1, y = 5859 s  -1  A  79.1 66.0 62.1 72.5  136.1 (4.3) 149.4 (23.6) 171.8 (7.3) 157.6 (19.8)  (4.4) (2.2) (3.8) (4.7)  + Isopar® M , 2a =9 0 , ^ = 352:1, y = 5859 s 171.2 (24.8) 38.0(1.3) 0.7(0.1) L/D= 0 133.0 (20.0) 64.8 (6.5) 2.0 (0.1) L/D= 5 178.6(16.5) 73.9 (2.7) 3.2(0.1) L/D = 10 157.6(19.8) 72.5 (4.7) 3.2 (0.1) L/D = 20 129.7 (20.2) 79.6 (4.0) 3.0 (0.2) L/D = 40 = 352:1 F104 H M W + Isopar® M , L/D = 20, 2a = 30, RR 99.70 (4.1) 94.7 (4.8) 6.1 (0.2) y = 1875 s" 102.22 (7.6) 88.7 (2.7) 5.9(0.1) y = l%\l s" 1  A  1  A  1  A  y = 3750 s  _1  A  y = 4688 s"  1  A  ^ = 5859 s"  1  5.7(0.1)  84.9 (4.9)  114.17 (11.5)  5.5 (0.1)  83.4 (5.6)  103.45 (7.6)  5.4 (0.1)  89.6 (3.7)  100.26(13.9)  F104 H M W + Isopar® M , L/D = 20, 2a = 30, RR = 352:1, y = 5859 s" 30.7 (9.9) 63.0 (3.9) 2.7 (0.4) T= 15°C 45.3 (7.3) 69.2 (2.9) 4.1 (0.2) r=i9°c 105.8(15.4) 61.2 (3.4) 3.7(0.1) r=22°C 100.3 (13.9) 89.6(3.7) 5.4 (0.1) 7-=35°C 86.0 (5.8) 92.0(11.2) 7.5 (0.2) J=45°C 110.4 (5.3) 93.1 (2.8) 7.0 (0.4) r=55°C 92.0 (4.6) 106.5 (3.4) 8.2 (0.2) r=65°C 1  A  108  6.8 Interpretation of PTFE paste extrusion It has been pointed out that fibril formation during P T F E paste extrusion is a critical factor. The quality o f these fibrils has been described in terms o f their degree o f orientation by using Raman spectroscopy (Lehnert et al., 1997; Ariawan, 2001). Ariawan (2001) has reported typical Raman spectra o f unprocessed and processed powder. In the case o f unprocessed powder, no difference in the scattering intensity at all Raman shifts was observed between the two polarization geometries (i.e. parallel or perpendicular to the extrusion direction), indicating no preferred orientation (isotropicity). In the case o f processed powder, however, a difference in the scattering intensity at major Raman shifts between the two polarization geometries was observed. This anisotropicity was interpreted as fibril orientation due to flow. Paradoxically, the high fibril orientation seen in the Raman spectra during extrusion is lost when the pressure reaches its highest point (yield pressure or maximum pressure). This was attributed to the increased level o f fibril breakages which now are no longer oriented by the flow. It is well known that the extrusion pressure can be used as a quantitative measure o f the extent o f fibrillation occurring during extrusion. High extrusion pressure usually means the formation o f a higher number o f fibrils. However, it is also true that an excessive pressure can cause the fibrils to break. The degree o f fibrillation i n an extrudate has been quantified in the past by using D S C analysis o f processed and unprocessed samples (Ariawan, 2002). Since the first heat o f melting o f a polymer is directly proportional to the degree o f resin crystallinity, the extent o f fibrillation and the quantity o f the fibrils can be assessed for different P T F E extrudates by measuring the difference in the first heat o f melting o f the paste before and after extrusion. Such a difference w i l l suggest that the resin crystallinity after extrusion has changed compared to the virgin resin and this is an indication o f the amount o f fibrils formed during extrusion. The larger the difference, the higher the degree o f fibrillation and hence the better the mechanical properties o f the product. It is noted that fibrils are crystallites which have been unwound due to the flow and have assumed an amorphous state. The D S C results are shown in Table 6.2. The values obtained by D S C analysis in many cases can be correlated with the results observed during extrusion and tensile strength measurement.  109  Table 6.2:  Heat o f melting and melting point o f P T F E resins extruded at different conditions.  Resin/Condition  Melting Point (°C)  Heat of melting  ^m,R  ~ m,S AH  (J/g)  m,R  Reference -78.3 (0.9) 333.2 (0.4) F104 H M W -75.8 (0.4) 332.5 (0.4) F104 L M W 331.2 (0.3) -74.1 (1.2) F301 -68.6(1.7) 334.0 (0.8) F303 F l 04 H M W + Lubricant, L / D =20, 2a = 30, R R = 352:1, y = 5859 s" 1  A  0.20 0.17 0.21  -62.9 333.0 Isopar® G -65.3 332.8 Isopar® M -61.6 332.7 Isopar® V F104 H M W , Isopar® M , L / D =20, 2a = 60, y = 5859 s'  1  A  0.02 -76.6 332.8 • R R = 352:1 0.15 -66.2 333.2 R R = 156:1 0.78 -61.4 333.1 R R = 56:1 F104 H M W , Isopar® M , L / D =20, R R = 352:1, y = 5859 s" 0.88 -69.3 333.7 2a = 15 0.17 -65.3 332.8 2a = 30 0.98 -76.6 332.8 2a = 60 0.09 -71.5 332.8 2a = 90 1  A  F104 H M W , Isopar® M , 2a = 90, R R = 352:1, y = 5859 s"  1  A  -77.6 331.8 L/D= 0 -54.2 333.2 L / D = 10 -71.5 332.8 L / D = 20 -72.8 332.9 L / D = 40 F104 H M W + Isopar® M , L / D = 20, 2a = 3 0 , R R == 352:1  0.02 0.31 0.09 0.07  ^ = 1875 s'  1  332.4  -72.5  0.07  y = 2812 s'  1  333.0  -71.5  0.09  y = 3750 s"  332.7  -64.7  0.17  ^ = 4688 s"  333.5  -70.0  0.11  ^ = 5859 s"  332.8  -65.3  0.17  A  1  A  1  1  F104 H M W + Isopar® M , L / D = 20, 2a = 30, R R = 352:1 , ^ = 5859 s' 0.04 -75.2 331.6 T = 15°C 0.10 -70.8 332.4 T = 22°G 0.17 -65.3 332.8 T = 35°C 0.16 -66.0 332.3 T = 55°C  1  110  6.9 Summary In this chapter, the effects  o f die design, resin molecular structure and physical  properties o f lubricants on the paste extrusion pressure and the mechanical properties o f extrudates obtained during extrusion o f pastes prepared with a variety o f P T F E powders and lubricants have been studied. It was found that a balance between fibril quantity and quality is necessary to ensure commercially acceptable products. Specifically, it was found that the physical properties o f the lubricant play a significant role in the extrusion o f P T F E pastes. Increasing the wettability o f the lubricant with P T F E and decreasing the lubricant viscosity causes a reduction in the extrusion pressure and an increase in the tensile strength o f the extrudates. It was also found that the geometrical characteristics o f dies also play a significant role in the process. The extrusion pressure increases with L/D ratio and reduction ratio, while it exhibits a minimum with the contraction angle at about 60°. These effects are also reflected in the mechanical properties o f extrudates. The tensile strengths o f dried and sintered extrudates exhibit optimum values at an L/D equal to about 20, and reduction ratio equal to about 150, whereas they exhibit a minimum value at a contraction angle o f about 60°. Finally, it was found that a higher molecular weight resin produces stronger fibrils that account for the mechanical superiority o f the final extrudates. Therefore, i f the objective is to prepare a product having a high tensile strength, a homopolymer o f high molecular weight should be used in combination with a lubricant having a low surface tension such as HFE-7500 and minimum viscosity greater than a certain value. Too low a viscosity would produce a nonuniform preform. A s far as die selection is concerned, this should have an L/D ratio o f about 20, reduction ratio o f around 150 and a contraction angle o f no more than 60°.  Ill  CHAPTER 7 Rheology of Preformed and Extruded PTFE Pastes 7.1 Introduction A s previously discussed, when P T F E paste flows from a large reservoir into a die o f significantly smaller diameter and consequently into the die land, fibrils between particles form that continuously change the rheology o f the paste. It is desirable to formulate a constitutive equation that would be suitable for describing the rheological behaviour o f the paste taking into account the continuous change o f its structure. Initially the preformed paste is a two phase material where the individual P T F E particles retain their identity. In other words, they are independent o f each other having a thin coat o f a lubricant around them. A s these particles flow through the contraction area o f the extrusion die, mechanical binding occurs that locks the particles together and the extensional nature o f the flow (contraction) causes unwinding o f crystallites. Thus, fibrils are formed that interconnect the particles giving a dimensional rigidity to the extrudate (see Figures 6.25 to 6.38). To describe the rheology o f such a complicated system, the following constitutive equation is proposed: T  where X  p  =  (1 -  £ )  X  p  +  [7-1]  is the contribution to the stress tensor from the presence o f the unfibrillated particles,  and X f is the contribution due to the fibrillated particles and £ represents the fraction o f the fibrillated particles, where £ = 0 corresponds to an unfibrillated paste, i.e. just preformed paste before it is subjected to flow, and £ = 1 to a fully fibrillated paste. This constitutive equation can be used for flow simulations in conjunction with an equation that can describe the dynamics of In the present chapter, two individual equations are formulated for the two extreme cases: (i) X  p  for the rheology o f preformed paste (absence o f fibrils) and (ii) T  f  for the  rheology o f a "fully" fibrillated paste, although it is difficult to determine experimentally, when a PTFE  paste has become fully fibrillated.  112  7.2 Rheology of dry PTFE powders and preformed pastes The study o f the rheology o f preformed pastes using various P T F E resins was carried out using a rate-controlled rheometer equipped with the parallel-plates geometry ( V O R Bohlin). A two-sided tape (the same size as the plate) was glued on the lower plate in order to prevent slippage o f the powder or paste. Similarly, the upper plate was serrated in order to avoid slippage during the test. Initially 1 g o f resin was placed on the lower plate and the plates were brought together to adjust the sample thickness. While pressing the powder between the plates, some resin was squeezed out and this excess material was removed. Figure 7.1 shows a strain sweep test in order to find the limits o f linear viscoelasticity. The resin used was F104 L M W and the test was carried out at 1.5 H z and 35°C. It seems that strains up to 1% can be used safely for linear viscoelasticity measurements.  Strain,y  Figure 7.1:  Strain sweep test o f F l 04 L M W powder at 1.5 H z and 35°C.  Frequency sweep tests were performed for the four different P T F E resins. The same amount o f resin was always used (1 g) and the small amplitude oscillatory shear tests were performed using 1% strain at 35°C. The results are shown i n Figure 7.2 for pastes prepared with the four resins and 38.8 v o l % o f Isopar M . Only one run for each paste is shown in this figure although several runs were carried out. A l l the runs for each paste were within 10% o f experimental error. A s it can be seen, there are no significant differences between the linear  113  viscoelastic properties o f these resins under these conditions. Note that the strain used is very small (1%) and the structural strength of the various powders in the presence or absence o f a lubricant is preserved. It is also possible to observe from Figure 7.2 that G ' »  G " indicating the  solid-like behaviour o f the samples (Larson, 1999).  Q_  (0  b  Frequency, ra (Hz)  Figure 7.2:  Frequency sweep test o f different resins (1% o f strain and 35°C).  The schematic depicted in Figure 7.3 shows another way by which the samples were prepared and loaded onto the parallel plate rheometer. One gram o f the resin was placed in a metallic ring and consequently compacted at the desired pressure for 60 s. This procedure assured a sample having a constant diameter o f 25.4 m m and a thickness o f about 1.6 mm. There was some concern about the development o f strains inside the sample due to compression. For this reason, some samples were tested in a strain sweep test under the same conditions at various times after their compression in order to examine the effect o f these stresses. Figure 7.4 shows the results. The samples exhibit almost the same viscosity within the experimental error. This result is important especially for the test involving the presence o f lubricant (paste). Due to its volatility, a significant amount o f lubricant may evaporate during the time taken to release the stress. These results show that the samples can be tested immediately after compression as the effects o f stresses developed during this step (compressions) are insignificant. 114  Figure 7.5 shows the effect o f preforming pressure. It can be seen that samples preformed at lower pressure exhibit higher viscosities. For pressures higher than 2 M P a the effect becomes insignificant. It should be mentioned that using a higher pressure to preform the paste with the method depicted in Figure 7.3 causes a significant loss o f lubricant.  Figure 7.3:  Preparation o f sample for parallel plate rheological test.  Once the region o f linear viscoelasticity was determined, frequency sweep tests were carried out at 1 % o f strain. Figures 7.6-7.8 show the effect o f temperature for a pure resin preformed at different pressures. It can be seen that the temperature has a small effect on the viscoelastic properties o f pastes.  10  5  > — • — Just r.irefc >rmec — « — After 15 rr lin &  —  After 50 rr lin After 24 h i  i  10"  10"  3  2  10"  1  Strain,y  Figure 7.4:  Strain sweep o f F104 H M W resin (frequency o f 1.5 H z and 35°C). 115  Figure 7.5:  Strain sweep test o f F104 H M W preformed at different pressures (frequency o f 1.5 H z a n d 3 5 ° C ) .  10°  10  1  Frequency, co (Hz) Figure 7.6:  Frequency sweep o f F104 H M W resin preformed at 0.5 M P a and tested at different temperatures.  116  Figure 7.7:  Frequency sweep o f F104 H M W resin preformed at 2 M P a and tested at different temperatures.  10  3  -! 10-  1 2  10"  1 1  10°  1 10  1  » 10  2  Frequency, co (Hz) Figure 7.8:  Frequency sweep o f F104 H M W resin preformed at 10 M P a and tested at different temperatures.  117  Figure 7.9 plots frequency sweep tests performed on different resins (dry powder). The resins were preformed at 0.5 M P a and tested at 35°C. N o differences among the resins are observed at such small strain (1%). Figure 7.10 compares the behaviour o f the pastes prepared with different P T F E resins and 38.8 v o l % Isopar® M . A l l the samples were preformed at 0.5 M P a for 30 seconds and the tests were performed using 1% strain at 35°C. There are practically no significant differences between the pastes. In fact, there is no effect o f the presence o f lubricant when the pastes are compared with the pure resins. The linear viscoelastic measurement presented in this section show that the preformed powders and pastes show a solid-like behaviour that preserves their strength at small values o f strain, i.e. up to strains o f 1%. This is also an indication o f the existence o f yield stress and, once this is exceeded, the powder and pastes w i l l exhibit signs o f fluid-like behaviour. Finally, the temperature and the presence o f lubricant do not have any significant effect on these viscoelastic measurements. These might be more important at larger deformations where the powder and pastes start flowing. However, these need other techniques or other types o f rheometers such as a sliding plate rehometer.  10-  2  10-  1  10°  10  1  10  2  F r e q u e n c y , GO ( H Z )  Figure 7.9:  Frequency sweep tests of different P T F E resins preformed at 0.5 M P a (1% strain and 35°C).  118  If) CO  •*  CL  CO CL  "co~  CL  b  10"  10°  1  F r e q u e n c y , co (Hz)  Figure 7.10:  Frequency sweep tests o f paste prepared with different resins and Isopar 38.8vol% and preformed at 0.5 M P a (1% strain and 35°C).  M at  7.3 Yield stress of PTFE pastes Concentrated solid-liquid suspension systems having strong interparticle interactions often exhibit unique plastic flow behaviour and the presence o f a yield stress (Dzuy and Boger, 1983). Under the application o f a small stress these systems deform elastically with finite rigidity, but when the applied stress exceeds the yield value continuous deformation occurs with the material flowing like a viscous fluid. The yield stress can thus be considered as a material property denoting a transition between solid-like and liquid-like behaviour. The yield stress is then the minimum shear stress corresponding to the first evidence o f flow, i.e., the value o f the shear stress at zero velocity gradient (Dzuy and Boger, 1983). Bingham and Green i n 1920 first recognized and introduced the concept o f a yield stress in fluid-like materials (Dzuy and Boger, 1983). M a n y workers have studied the measurement o f yield stress, but there is a general disagreement in results between the different methods used. In fact, there is doubt whether yield stress exist at all. Some people state that yield stress exists as a well-defined quantity and its value is unique for a given material. Others maintain that yield stress does not really exist (Cheng, 1986).  119  In this work, the yield stress o f four different P T F E resins was measured by direct and indirect methods. A l l the different methods were carried out in a stress-controlled rheometer equipped with parallel serrated plates (Bholin C - V O R ) . The first method employed was a ramp stress test. This test is a quick way to find the yield stress o f a material but it is not accurate i f insufficient time to reach steady state is allowed. The test involves applying a gradually  increasing  stress and  monitoring  the  instantaneous viscosity estimated as: % - 7  [7-2]  where a is the applied shear stress, andy is the measured shear rate upon the application o f the stress. A plot o f the instantaneous viscosity versus shear stress is obtained and the yield stress is read directly from the plot as the maximum instantaneous viscosity. Figures 7.11 to 7.14 show such plots for homopolymers F104 H M W and F104 L M W and copolymers F301 and F303, respectively. The samples were dry powders prepared by compression at 0.5 M P a for 30 seconds as depicted in Figure 7.3.  30x10 i  1  6  0  I  1  0  500  1  1  1000  1500  1  1  1  2000  2500  3000  Shear Stress, a (Pa) Figure 7.11:  Y i e l d stress measurements for F104 H M W resin at 35°C.  120  30x10  6  Run#1CT = 1028.3 Pa| o  Run#2o = 840.7 Pa  (0  CL-  o  25x10  Run #3CT = 934.5 Pa  6  o  Run #4 a = 887.6 Pa o  Run#5o- = 840.7 Pa 0  > IO O O CO  20x10  Run#7a = 699.1 Pa  6  c  Run #8 o\. = 652.2 Pa  15x10  f  CO  3  o  <D  C TO C  ™ CO  10x10  5x10  v\ \  7/  6  6  €  c  500  1000  1500  2000  2500  3000  Shear Stress, a (Pa)  F i g u r e 7.12:  40x10  Y i e l d stress measurements for F104 L M W at 35°C.  6  Shear Stress, a (Pa)  F i g u r e 7.13:  Y i e l d stress measurements for F301 at 35°C. 121  40x10 Run #1 a = 1028.3 Pa  to  0  *  Run#2ci = 0  CO  30x10  981.4 Pa  R u n # 3 a = 746.9 Pa D  6  Run #4CT„= 887.6 Pa  '</)  o o  g  20x10  6  10x10  6  to 3 O  o c jS c  rc to c  +j  500  1000  1500  2000  2500  3000  Shear stress, a (Pa)  Figure 7.14:  Y i e l d stress measurements for F303 at 35°C.  Table 7.1 summarizes the results for the mean yield stress values determined from the above plots with the standard deviation from various tests in parenthesis.  Table 7.1:  Yield stress of PTFE resins. Y i e l d Stress (Pa) 869(103) 840(130) 1044 (97) 911(124)  Resin F104 H M W F104 L M W F301 F303 The  determination o f yield stress usually implies steady shear under prolonged  application o f the shear stress. When the stress is first applied, however, the material shows a creep response. For low values o f stress, the elastic component o f the material w i l l play a more important role than the viscous part. The strain eventually w i l l attain a constant value depending on the level o f stress. However for high values o f stress, the strain w i l l increase indefinitely since the viscous component o f the material is more dominant and it is the shear rate which w i l l reach a constant value. 122  Figures 7.15 to 7.18 show creep tests for P T F E fine powder resins. The samples were prepared using the same method described before (see Figure 7.3). F o r these tests, a stress o f 600 Pa for 100 seconds was applied. The level o f the applied stress was selected to be smaller than the yield stress determined from the first method (see values i n Table 7.1). If this is true then according to the values listed in Table 7.1, a constant value for the strain (or alternatively for the compliance) should be reached. However, it can be seen from Figures 7.15 to 7.18 that the compliance keeps increasing, indicating that either the stress applied is above the yield stress or the elapsed time is not large enough. In addition, the repeatability o f each test suffers significantly. In order to determine the yield stress, a lower stress should be applied. However, there is a practical complication in fixing the duration o f the test and trying to decide whether a sample has attained constant strain or is still deforming. In general, the longer the observation period the lower the yield stress. This means that the yield stress is a time dependent property (Cheng, 1986).  12x10-  6  £  2x10" 6  0  I 0  1 20  1 40  1 60  1 80  1 100  T i m e (s)  Figure 7.15:  Creep test for F104 H M W fine powder resin at 35°C (cr= 600 Pa, r = 100 s).  123  50x10"  (  0  I  0  1 20  1 40  1 60  1 80  ' 100  ' 120  T i m e (s)  ure 7.16:  Creep test for F104 L M W fine powder resin at 35°C (cr= 600 Pa, r = 100  12X10"  6  2  2x10" 6  0  20  40  60  80  100  120  T i m e (s)  Figure 7.17:  Creep test for F301 fine powder resin at 35°C (cx= 600 Pa, t = 100 s  124  10X10-  6  o O a. to u  2X1G"  6  -  O  0  I  1  1  1  1  0  20  40  60  80  1 100  120  T i m e (s)  Figure 7.18:  Creep test for F303 fine powder resin at 35°C (<j= 600 Pa, t = 100 s).  Figures 7.19 to 7.22 show the results for creep tests with a duration o f about 1600 s. For each resin, different values o f stress were selected. In each run, the sample was subjected to the given stress and the sample was discarded after each run. A high value o f stress was taken as the initial guess according to Table 7.1. It can be seen how the slope o f the creep compliance keeps increasing with the application o f a high stress to the sample indicating that this high value o f stress does not correspond to the yield stress o f the resin. The stress was then decreased until the compliance remained almost constant for a certain period o f time until the end o f the test. Due to the nature o f the test, it can be said that F104 H M W exhibits a yield stress between 90 and 100 Pa; F104 L M W shows a yield stress between 120 andl25 Pa; for F301, the yield stress is around 160 Pa; while for F303, the yield stress falls between 160 and 170 Pa.  125  35x10"  6  i  i  30x10-  6  ro Ql  25x10"  6  I  20x10"  6  O  V  c ro  E o o a  i  • ©  15X10"  6  90 Pa 100 Pa 120 Pa 125 Pa 127.5 Pa 130 Pa 150 Pa  a •  10x10"  6  v  a> 0)  O  5x10  J  200  400  600  800  1000  1200  1400  1600  1800  T i m e (s)  Figure 7.19:  Creep test for F104 H M W resin at different levels o f stress.  30x10"  6  ©  ro Q_ 25X10-  6  ©  s  • 20X10"  6  A  0  A  0  V A  A  V  A  •  0 0  o  <$>  •  c ro 15X10"  6  *  O  Q.  E o o a  10x10"  a) ,«r  5x10-'  6  •  n •  c •  •  A  V •  c ©  200  400  600  800  1000  1200  1400  120 125 130 140 160 170 180  1600  Pa Pa Pa Pa Pa Pa Pa  1800  T i m e (s)  Figure 7.20:  Creep test for F104 L M W resin at different levels o f stress.  126  50x10-  6  i  • Q. 40x10" S:  6  V  •  <D O C  ro "EL  E o O a  160 Pa 165 Pa 170 Pa 190 Pa 200 Pa 400 Pa  30x10"  6  20x10"  6  2  8 •  0  10X10"  6  o  200  400  600  800  1000  1200  1400  1600  1800  T i m e (s)  Figure 7.21:  Creep test for F301 resin at different levels o f stress.  16x10"  6  •  o o  .a •  A A A *  S  i  £- 4x10" <x>  6  l_  °  • a A  2x10"  6  200  400  600  800  1000  1200  1400  160 Pa 170 Pa 180 Pa  1600  1800  T i m e (s)  F i g u r e 7.22:  Creep test for F303 resin at different levels o f stress.  127  Another way to determine the yield stress is by performing a step stress test. This test is basically a multiple creep test that allows measurement o f the yield stress in a more accurate way although the test lasts longer. For this method, a set o f stress values are selected. Each stress level is applied and held for a pre-defined time, while the strain response is measured. The stress is gradually stepped up until a measurable flow is obtained. Subsequently, the shear rate is plotted as a function o f shear stress and the yield stress is determined when a change in the shear rate is attained. Figures 7.23 to 7.26 depict the results o f the step stress tests applied to the P T F E samples preformed at 0.5 M P a for 60 seconds. The plotted data suggest the existence o f a yield stress. The initial portion o f the curves where the shear rate remains constant is clear and so is that portion when the shear rate starts increasing, i.e., when the sample starts to flow.  30x10"  6  25x10-6 L  « . 20x10"  6  of 15x10  • •••  6  ra £ 10X10"  6  o  CO  5x10-'  0  200  400  600  800  1000  1200  1400  1600  S h e a r S t r e s s , a (Pa)  Figure 7.23:  Step stress test for F104 H M W fine powder resin preformed at 0.5 M P a .  128  S  30x10-  6  25x10-  6  20x10"  6  "JS  15X10"  6  co £CO  10X10"  6  CD  V  (/) 5x10"'  0  200  400  600  800  1000  1200  1400  1600  S h e a r S t r e s s , a (Pa)  Figure 7.24:  Step stress test for F104 L M W fine powder resin preformed at 0.5 M P a  30x10  CO 10X10"  6  CO  200  400  600  800  1000  S h e a r S t r e s s , a (Pa)  Figure 7.25:  Step stress test for F301 fine powder resin preformed at 0.5 M P a .  129  30x10"  6  25x10"  6  to, 20x10"  6  0  15x10  6  L  or ra o 10x10"* I £-C  4  CO  *J" - -** ° 5  -  5x10-  6  0  200  400  600  800  1000  1200  1400  1600  S h e a r S t r e s s , a (Pa)  Figure 7.26:  Step stress test for F303 fine powder resin preformed at 0.5 M P a .  A n indirect method can be used to determine the yield stress from Figures 7.23 to 7.26, i.e., by fitting to the data with well known models where the yield stress is a parameter o f the model. Here, the following models were used: Bingham (j = cr +T]r  [7.3]  <T = CT +riy  [7.4]  0  Herschel-Bulkley o  Casson V. . v.  a" = <j' ' +tj"y"  [7.5]  0  Figures 7.27 to 7.30 show the shear stress-shear rate plots with fits o f the three viscoplastic constitutive equations presented above. Only the straight portions o f the plots have been fitted and the parameters o f the models are listed in Table 7.2. Even though the HerschelBulkley model fits the data better i n some cases than the other two models (see R i n Table 7.2), 2  the yield stress values determined with this model are always negative. Therefore, it seems that this model is not suitable to represent the whole range o f data. From the other two models, the Casson model seems to perform better and therefore might be a good choice to represent the  130  rheology o f the paste as a first approximation. The values o f the yield stress determined from this model are in general smaller than those reported in Table 7.1 but higher than those obtained from the creep test. Therefore, one may conclude that the yield stresses o f the powders are very small and this is the reason why fitting procedures some times yield negative values for the yield stress. Similar conclusion has been drawn by Ariawan (2001) who has neglected the yield stress.  Figure 7.2:  Y i e l d stress o f various pastes determined by means o f fitting viscoplastic models to the experimental data depicted in Figures 7.23 to 7.26.  Parameter F104 H M W Oo (Pa) r] (Pa s") k (dimensionless) R F104 L M W o (Pa) 77 (Pa s ) k (dimensionless) R F301 (Jo (Pa) rj (Pa s ) k (dimensionless) R F303 (Jo (Pa) TJ (Pa s ) k (dimensionless) R 2  0  m  2  m  2  m  2  Herschel-Bulkley  Casson  -1.35xl0 1.35xl0 2.983xl0" 0.959  330.5 1.62xl0  7  7  5  Bingham  -  -  0.971  0.965  -2.126xl0 2013xl0 , 1.172xl0" 0.995  744.0 1.041xl0  -4.786xl0 4.791xl0 6.25xl0" 0.976  6  595.2 3.247xl0  -1.053xl0 1.054x10 3.251xl0" 0.986  7  6  6  4  6  5  7  5  570.3 3.723xl0  7  6  945.5 4.648xl0  -  -  0.967  0.930  6  825.5 LlOlxlO  -  -  0.935  0.891  518.6 9.021xl0  6  759.4 2.739xl0  -  -  0.977  0.957  131  7  7  7  7  1600 1400 ro 0-  1200 1000  to to  o  1_ •«->  CO  800  l_  600  CO  400  re o  Experimental Data Herschel-Bulkley Casson Bingham  —  200  5x10"  10x10"  6  15x10"  20x10  6  25x10  6  6  S h e a r Rate, x(s' ) 1  F i g u r e 7.27:  Fitting viscoplastic models to the rheological data o f F104 H M W fine powder resin. 1600 1400 V  ro  to CD  CO ro cu  .c  CO  •  Experimental Data Casson  Herschel-Burkley Bingham  0  20x10"  6  40x10"  6  60x10"  6  80x10  6  100x10"  6  120X10-  1  Shear R a t e , / ( s ) 1  F i g u r e 7.28:  Fitting viscoplastic models to the rheological data o f F104 L M W fine powder resin. 132  1600 1400 h  2  1200 r  b to 1000 r 10  0  CO  800 h  CO  a> £  600 I-  .f  •  Experimental Data Casson Herschel-Burkley "] Bingham  400 r _.° 200 20x10-  40x10-  6  60x10"  6  6  80x10"  6  Shear Rate,^ (s' ) 1  F i g u r e 7.29:  Fitting viscoplastic models to the rheological data o f F301 fine powder resin. 1600 1400  to  1200  CL  b  1000  10  CO 0)  800  CO  co  600  (J)  .c  W  Experimental Data Bingham Herschel-Burkley Casson  400 200  10x10"  20x10"  6  6  30x10"  6  Shear Rate,^ (s' ) 1  F i g u r e 7.30:  Fitting viscoplastic models to the rheological data o f F303 fine powder resin.  133  7.4 Extensional reology of extrudates obtained from slit die extrusion Two different pastes with F104 HMW-Isopar M and F301-Isopar M at 38.8 v o l % o f lubricant were prepared in the usual way. A slit die was used to extrude the pastes in order to produce rectangular shaped samples to load them onto the S E R extensional rheometer. The slit die has an entrance angle o f 2a = 60° and L/H = 20 whereas the slit is 2.62 mm wide and 0.508 mm high. The pastes were extruded at shear rates varying from 854 to 5859 s" at 35°C. 1  Figures 7.31 and 7.32 show the steady-state extrusion pressure and the tensile strength, respectively, at different shear rates. The tensile strength was measured with a Corn-Ten apparatus by stretching the sample at a constant speed o f 13 mm/s until the sample fails. The dimensions o f the sample were: 80 mm in length, 2.54 mm in width and 0.34 mm in thickness. A s with the cylindrical dies, the extrusion pressure increases with increase o f the apparent shear rate (Figure 7.31) while the tensile strength decreases with increase o f the shear rate (Figure 7.32). It means that better quality products are obtained at lower extrusion rates.  ig I . . . . i . . . . i . . . . i . . . . i . . . . i . . . . i . . . . i 0  1000  2000  3000  4000  5000  6000  7000  Apparent Shear Rate (s~ ) 1  Figure 7.31:  The extrusion pressure o f various paste extruded with a slit die as function o f the apparent shear rate at 35°C.  134  3.0  2.5  i I i i i i I i i i i I i i i i I i i i i I i i i i I i • • i —•—F104 HMW •  h  F301  •  F303  CO  a  2.0  h  W m  1.5  h  CO .2 (/> C 0)  1.0  F104LMW  —A—  .*-*  r-  0.5 f0.0 I i i i i I 1000  2000  3000  i i i I i i i i I 4000 5000  6000  7000  Apparent Shear Rate (s ) 1  Figure 7.32:  Tensile strength o f extrudates obtained from slit die at 35°C as a function o f the apparent shear rate.  The samples were subjected to steady Hencky strain rate extensional rheology tests. A s discussed before, a Senmanat  Extensional Rheometer ( S E R ) was attached  to the strain  controlled rheometer for these measurments. A l l tests were performed at 35°C. The following equations were applied in the data analysis and interpretation. The linear strain, e, is defined as follows: £  I  I  AL  L  [7.6]  L  0  0  where L is the initial length o f the sample, L is the length o f the sample at any time and AL is 0  the increment o f length. The Hencky strain, £H, is defined as: L^  f  s  H  =ln  [7.7]  v A) j which can also be written as: e  H  =ln  L  f AT \ AL + 1 , = ln —  L + L Q  L  0  0  [7.8]  JJ  Combining Equation 7.6 with 7.8, s  H  135  = \n(s +1)  [7.9]  The linear strain, e, can be expressed in term o f the Hencky strain, en, as follows: e = exp(s )-l  [7.10]  H  The engineering tensile stress, <J , was calculated with the following formula: S  cr =^— 2RA  [7.11]  s s  0  where T is the torque, R is the drum radius and Ao is the initial sample cross sectional area. The true tensile stress, <7g, was calculated as follows: a= E  2RA Qxp(-ts ) 0  =  H  ^ exp(-t£ )  [7.12]  H  The tensile stress growth coefficient was estimated with the following generalized formula for SER: T ?  E  = ^ -  [7  .13]  First, the collected extrudate samples obtained at different apparent shear rates were stretched at the same Hencky strain rate o f 0.0113 s" . The engineering tensile stress was 1  calculated and plotted as a function o f the linear strain. The results for F104 H M W and F301 resins are shown i n Figure 7.33 and 7.34, respectively. The lines i n these figures correspond to averages from several tests carried out with different specimens from the same extrudate. The response for both materials are similar and resemble the responses obtained with the Corn-Ten tester; that is the tensile strength (maximum tensile stress value in the corresponding curve) increases with a decrease in the shear rate at which the corresponding specimen was extruded. This means that more fibrils are present in the specimens extruded at smaller shear rates.  136  Figure 7.33:  Engineering tensile stress response o f F104 H M W extrudates ats  H  Linear Strain, e Figure 7.34:  Engineering tensile stress o f F301 extrudates at e  H  137  = 0.0113s  = 0.011  The stress growth coefficients o f these samples were calculated by using Equation 7.13. The results are plotted in Figures 7.35 and 7.36 for the F104 H M W and F301 extrudates, respectively. There are no significant differences among the extensional viscosities for samples extruded at different shear rates. It seems that the extensional behaviour o f all samples of the same resin are essentially the same, i.e. the degree and orientation o f fibrillation are similar. However, the difference is perceptible for different resins as shown in Figure 7.37. The homopolymer F104 exhibits higher values than the copolymer F301. It seems that F303 exhibits the lowest extensional transient viscosity.  T i m e (s)  Figure 7.35:  Stress growth coefficient o f F104 H M W extrudates obtained from a slit die with L/H= 20 and different apparent shear rates at 35°C.  138  Figure 7.36:  Stress growth coefficient o f F301 extrudates obtained from a slit die with L / H = 20 and different apparent shear rates at 35°C.  10-  1  10°  10  10  1  2  10  3  T i m e (s)  Figure 7.37:  Comparison o f extensional behaviour o f slit die extrudates obtained at y =3750 A  s" and subjected to an extension at z = 0.0113 s" . 1  1  H  139  Another series o f tests was carried out on extrudates obtained when the pastes were extruded at an apparent shear rate o f 3750 s" . The samples were stretched at Hencky strain rates 1  varying from 0.00113 s" to 11.3 s~\ The stress growth coefficients were calculated from 1  Equation 7.13 and the results are shown in Figures 7.38 to 7.41. Again, each line in these figures is the average calculated from several tests performed at the same Hencky strain rate. The response is typically that obtained from a linear molten polymer and the lack o f plateau in the plots is an indication o f a viscoelastic material.  io-  2  id-  1  io°  io  1  10  2  10  10  3  4  io  5  T i m e (s)  Figure 7.38:  Stress growth coefficient for F104 H M W at different Hencky strain rates. Samples used were produced by extruding the resin at y  A  140  = 3750 s~ .  10-  2  1CV  10°  1  10  1  10  10  2  10  3  4  10  5  T i m e (s)  Figure 7.39:  Stress growth coefficient for F301 at different Hencky strain rates. Samples used were produced by extruding the resin at y  A  10" -I 2  10  —i 2  :  10"  1  ——i  10°  1 10  10  2  10  -i  1  —i  i 1  = 3750 s" .  10  3  4  10  5  T i m e (s)  Figure 7.40:  Stress growth coefficient for F104 L M W at different Hencky strain rates. Samples used were produced by extruding the resin at y = 3750 A  141  s~.  10-  2  10-  10°  1  10  10  1  10  2  3  10  10  4  5  T i m e (s)  Figure 7.41:  Stress growth coefficient for F303 at different Hencky strain rates. Samples used were produced by extruding the resin at y  A  = 3750  s' . 1  7.5. Yielding and deformation behaviour in extension Figures 7.42 to 7.45 are plots o f the engineering tensile stress, <Js, versus strain, s, for extrudates subjected to uniaxial extension at different strain rates and at room temperature. The same scale has been used in all the plots for comparison. The strain was estimated from the Hencky strain by means o f Equation 7.10. From each curve two fundamental quantities may be obtained. First the yield stress, <r , as defined in section 2.5 as the value o f the stress at the first y  knee in the stress-strain curve. It is obtained at a certain strain value which is defined as the yield strain,e . It can be seen from Figures 7.42 to 7.45 how the yield stress (<7 ) changes as a y  y  function o f the Hencky strain rate {s ). H  The strain rate effects follow the same trend as for  molten polymers (Crist, 1993), i.e. increasing e  H  about 50 %. In addition, the yield strain ( e  by one order o f magnitude increases a by y  ) and the tensile modulus (E) are essentially  independent o f the strain rate except for a very high value o f e . H  142  In summary, the changes in qj,,  s  y  and E at the highest s  H  resemble the behaviour o f a glassy polymers near T (Crist, 1993). g  The values o f <j , and s are listed in Table 7.3. y  Table 7.3:  v  The yield stress and strain o f P T F E extrudates at room temperature. F104 H M W y (MPa) 1.4 1.8 2.6 4.2 CT  (•"•) 0.00113 0.0113 0.113 1.13  F104 L M W y (MPa) 1.0 1.8 2.2 3.6 CT  1.0 0.8 0.8 1.2  y  £  1.0 1.0 1.0 1.3  F301 a (MPa) 1.0 1.6 2.0 2.8 y  F303 h 1.3 0.9 1.2 1.0  y (MPa) 1.6 1.7 2.0 2.8 CT  h 0.4 0.4 0.4 0.6  The data listed in Table 7.3 indicate that F301 shows the higher extensibility followed by the two homopolymers and finally by F303, which has the smallest extensibility.  Figure 7.42:  Engineering tensile stress-strain curves for F104 H M W as a function o f  143  e. H  ro CL in b  CO CO  o  k.  CO  Q)  'co c 0  hO)  c  0 0 C  '5) c  LU 3  4  5  Linear Strain, £  Figure 7.43:  Engineering tensile stress-strain curves for F104 L M W as a function of e  H  ro CL id  b  Hencky strain rate, £  4r  (s")  0.00113 0.0113 0.113 1.113  US to  0  ft  H  3  0  'co §  2 •  c  X  X  03 0  c  "5> c  LU 3  4  5  Linear strain, e  Figure 7.44:  Engineering tensile stress-strain curves for F301 as a function of s  144  ^ ro  5  ' ' ' ' I i i i i I i i i i I ii  i  CL  11111111111111111 Hencky strain rate, £  (s ) 1  w  v> 4 0.00113 0.0113 0.113 1.13  CO  "5 to  3 r  /  '</>  § IUi c "C  2  co  CO  c c LU  0 I i i i •i ••• 1 2  3  4  5  • 6  7  8  Linear Strain, s  Figure 7.45:  Engineering tensile stress-strain curves for F303 as a function o f f ,  7.6. Modelling the extensional behaviour through the strain-energy function Stress, defined as a force per unit area may be tensile, compressive, or shear. Nine separate quantities are required to completely describe the state o f stress i n a material. Stress is indicated as ay where the first subscript refers to the orientation o f the face upon which the force acts and the second subscript refers to the direction o f the force. Usually, in rheology numbered coordinate directions are used and the components o f the stress tensor are written as: r,, °21  ex,, a  22  cr, '23  [7.13]  '33 J  The diagonal components in Equation 7.13 correspond to extensional stresses whereas the other components correspond to shear stresses. In addition, the stress matrix is symmetrical, i.e., a j = ay when i t  j; meaning there are only six independent components in the stress tensor  represented by Equation 7.13. In a similar way, the deformation experienced by the material under stress can be expressed with as strain tensor. Considering the change in the deformation o f a material element in present state, t, with respect to the past state,/', defines the displacement gradient tensor, Ftf. 145  where x x' represent the present and past position, respectively and the subscripts represent coordinate directions. A more common strain tensor used in rheology is the Finger tensor, By, defined as the dot product o f the Fy and its trasposse: dx. B«=F -F =^r ik  dx 3x  ;  [7.15]  J  ik  3x  4  k  When dealing with viscous material, it is common to consider the material at rest when it can support only uniform normal stresses (hydrostatic pressure p) and when in motion. Thus the total stress, t , is the sum o f these two terms. In vector and tensor notation this is written as: T = -pI + a  [7.16]  where / is the identity tensor and j j is the viscous strees tensor or just stress tensor (Equation 7.13). Only stress differences have meaning. Two such normal stress differences are used: l  N  =  r  i l  — 2*22  where N/ and  ~ 22  = C J  T  ^~33  11 ^"22  —  -°"22  ^  ^33  define the first and second normal stress differences.  The stress-strain data obtained from the uniaxial extension tests and presented in sections 7.4 and 7.5 indicate that these materials behave like a viscoelastic solid since they essentially fail after yielding. The materials have been subjected to uniaxial extension. The stress applied during the process is:  A where F is the force applied on a plane o f area A perpendicular to the direction o f deformation. The area, A, changes during the process and then an is refered as the true tensile stress, cr . If A £  0  is the initial area, then the stress is refered as the engineering tensile stress, as. In extension, it is common to use the extensional ratio, ot\, defined as:  JLv;  1,0  to denote deformation where L  i0  and Z, are the length in the / direction at t = 0 and t - t,  respectively. 146  The mechanical properties  o f a perfectly  elastic material may conveniently be  represented in terms o f the strain energy W. For a perfect elastic solid at equilibrium, the stress can only be a function o f the change i n the internal energy U o f the sample away from its reference state due to a deformation ^r j T =p  [7.20] SB  where B is the Finger strain tensor deifned before (Equation 7.15) For an incompressible material, Equation 7.20 is expressed as (Macosko, 1994):  T  =-pI  2 ^ B - 2 ^ B - < di dii  +  B  [7.21]  B  where Wis the strain energy function. Here the material functions are derivatives o f the strain-energy function with respect to the invariants o f B . I f the material is considered incompressible with respect to volume, the principal extension ratios (Dzuy and Boger, 1983) are related by: a,a a 2  3  =1  [7.22]  Then the two strain invariants can be expressed by: I  B  where  IIIB —  and  =a] + O C 2 + C C 3  II  = a\  2  B  +aj  + a~  2  [7.23]  2  1 since the material is considered incompressible.  In each case, W is not known but must be determined experimentally. Since W is a function o f h and II , Valanis and Landel (1967) proposed that Wis separable into the sum o f B  the same function o f each o f the principal extensions ccj W = vv(a,) + vv(a ) +vv(oc ) 2  [7.24]  3  They found that an exponential function fit data w(a ) i  = ma  [7.25]  ni  i  i  In a uniaxial extension test the main extension ratio is just a/ and the following simplifications are found 2 a =a =a ~ 2  3  1 / 2  1  1  I =a -f—  II =2a +  2  B  B  a!  j  —  [7.26]  a,  The extensional ration, cti, can be related with the linear strain (Equation 7.6), e, and the Hencky strain (Equation 7.7), 8 h , as follows: a , = s +1  and  147  e  H  = ln(a ) l  [7.27]  Thus, the principal stress cr,, is given by. a  = a w'(a )-a ' w'(a~ ) U2  n  ]  l  [7.28]  U2  1  where \v'(ct\) is the derivative o f PTwith respect to « / . Once an analytical expression for w' is found, W can be determined. Extensional and compression experiments or biaxial extension tests should be conducted in order to apply the separation method to find this analytical form o f ^ ( V a l a n i s and Landel, 1967). However, once the function is found, it can be used to fit simple extension data (Kearsley and Zapas, 1980). Ogden (1972) proposed a strain-energy function which is a linear combination o f strain invariants discarding the requirement that W shall be an even-powered function o f the extension ratios (Jones and Treloar, 1975) and finding an expression for W as VH-  I  H;  W = T—W i n  U-  +a  l 2  71-  +a  l 3  -3  [7.29]  t  The Ogden model fits experimental data much better than the M o o n e y - R i v l i n model (Jones and Treloar, 1975; H u and Desai 2004) without the constraint imposed by the latter about the material constants. The number o f terms included i n the summation is to be determined by comparison with the experimental data. From the incompressibility o f the material (Equation 7.26), Equation 7.29 can be written as: W=>Z  m  :  a" + 2 « , 1  2  - 3  and the derivative with respect to a\ is dW da,  =5>. a ^ - a ^  [7.30]  2  When Equation 7.30 is substituted into Equation 7.28 the following expression for a , , results: a  1 x  - a.  2  [7.31]  Equation 7.31 was used to fit the data shown i n Figures 7.42 to 7.45 obtained from extensional rheology. Just two parameters were used to fit the data since no improvement was noticed when four parameters or more were used. Thus, the extensional data can be modeled b y an equation having the following form: 148  <r =m(a' ' n  ]  [7.32]  -a;" ) 12  The tensile or Y o u n g ' s modulus, E, for uniaxial extension can be defined as: E = 1 im  £->0  Recalling that the extensional ratio, a , can be expressed in term o f the linear strain, e, t  Equation 7.32 can be used to estimate E as follows: — —m-n 2  [7.33]  Table 7.4 lists the values o f the constants for the extrudates obtained from pastes prepared with the four different P T F E resins and Isopar M . The engineering tensile stress and Young's modulus obtained from Figures 7.42 to 7.45 as well as those obtained by Equations 7.32 and 7.33 are also included. The fittings were performed with a program that is listed in the Appendix.  Table 7.4:  Ogden's parameters for P T F E samples subjected to different Hencky strain rates.  Hencky strain rate (V ) 0.00113 0.0113 r £ 0.113 1.13 0.00113 0.0113 T 0.113 O E 1.13 0.00113 0.0113 © 0.113 1.13 0.00113 fO 0.0113 © 0.113 1.13  Tensile Stress (MPa) m (MPa)  n  12.76 16.63 74.24xl0 87.20xl0 22.17 24.90 8.271xl0 80.00xl0 39.11 19.42 146.9xl0  0.22 0.22 7.0xl0" 8.5xl0" 8.9xl0" 0.14 5.2xl0" 8.0xl0" 4.9xl0" 0.15 2.6x10" 1.3xl0' 0.24 9.2xl0" 3.8xl0" 2.9xl0"  Measured  1  41.25xl0 19.71 4739 115.7xl0 17.50xl0  2  3  2  3  4  5  2  3  4  2  2  3  2  3  4  4  4  4  4  1.5, 1.9 2.6 4.2 1.0 1.9 2.2 3.7 1.0 1.5 2.0 2.9 1.4 1.7 2.1 2.8  Equation 7.35 1.63 2.15 2.87 4.09 1.11 1.98 2.38 3.53 1.07 1.62 2.07 2.82 2.78 2.41 2.41 2.89  Elastic Modulus (MPa) Measured Equation 7.37 10.10 . 4.17 5.49 7.80 11.12 2.95 11.56 5.18 6.45 9.60 2.87 6.96 4.24 5.62 7.67 7.06 9.54 6.54 6.56 7.84  Figures 7.46 to 7.49 show the fits obtained by using Equation 7.32 for the samples stretched at different Hencky strains. The scale in each plot has been kept the same for the sake  149  of comparison. It can be seen that Equation 7.32 fits the data reasonably well at low Hencky strain rates, but exhibits some problems for high Hencky strains due mainly to the abrupt fall o f the curve. Resin F303 displays the lower extensibility and the model is incapable o f fitting these values since it was developed with a rubber-like material in mind. One possible solution to this discrepancy could be to use Equation 7.32 with more parameters, but as was mentioned earlier, there is no significant improvement. Ogden (1972) recommends splitting the plot into as many regions as needed and finding the parameter for each region. Then by a linear combination o f those equations and adjustment o f the parameters, the final expression for the model can be found. This can be carried out by trial and error. The values o f the parameters for a high Hencky strain rate look out o f proportion, but those values are necessary for the model to follow the trends shown by the data. In addition, the tensile stresses calculated with these parameters by means o f Equation 7.32 are very close to those read directly from the corresponding plots. O n the other hand, the elastic modulus calculated from Equation 7.33, continues to vary as a function o f the Hencky strain rate which is not true as can be seen from Figures 7.42 to 7.45. It can also be seen from Figures 7.46 to 7.49, how the maximum o f the fitting data is attained at the same extensional ratio, a\ = e. This is not surprising since, according to Equation 7.32, the extensional ratio, a.\, at which the maximum stress is reached, is a function o f the parameter n only with the form:  [7.34] which tends to the value e when n tends to zero. In order for the engineering tensile stress, a, to be a maximum, n has to be between 0 and 1 in Equation 7.34. In general, and i n spite o f few discrepancies, it can be said that the model works fairly well and is especially useful as it can classify the various P T F E pastes according to their mechanical properties.  150  Linear Strain, e  Figure 7.46:  Uniaxial extension o f F104 H M W samples stretched at different Hencky strain rates. Continuous line represents the model fitting.  0  1  2  3  4  5  6  7  8  9  10  Linear Strain, e  Figure 7.47:  Uniaxial extension o f F104 L M W samples stretched at different Hencky strain rates. Continuous lines represent the model fittings.  151  Figure 7.48:  Uniaxial extension o f F301 samples stretched at different Hencky strain rates. Continuos lines represent the model fittings. 20  i iiii iiiiiiiiiiiiiiiiiii  i • • • • i  iiiiiiiiiiiiiii  Hencky strain rate, £  (s ) 1  H  CO 0L *  Hi b w  0.00113 0.0113 0.113 1.13  15  (0  £  10  a> < ' /> c a> a; 3  0  w  0  1  i * * '  2  3  4  5  6  7  8  9  10  Linear Strain, s  Figure 7.49:  Uniaxial extension o f F303 samples stretched at different Hencky strain rates. Continuous lines represent the model fittings.  152  7.7 Summary In this chapter the rheology o f processed and unprocessed P T F E paste has been studied. The formation o f fibrils during P T F E paste extrusion introduces a complication in the study o f this material since their continuous formation and breakage change the rheology o f the material. In an attempt o f modelling the rheological behavour o f P T F E paste during extrusion, an appropiate constitutive equation was proposed as follows: T  = (1-  £)T  + tJT  p  f  Regarding r , it seems that the unprocessed paste follows the Casson model. The yield p  stress values predicted by this model fall between those values obtained from two different tests and perhaps they must be considered reasonable as a first approximation. It is worthwhile to mention that the presence o f the lubricant does not seem to influence the value o f the yield stress. Experiments at large strains and strain rates should be performed to evaluate the other parameters with certainty. Regarding ry, the extensional rheology o f the processed resin has shown that extrudates behave as linear elastic polymers. It has been observed that the yield stress and strain are independent o f the rate at which the samples are stretched, but they depend on the rate at which the paste is extruded. It is noted that the latter influences the degree o f fibrillation. It has been discussed in this chapter that the processed paste can be modelled by means o f the Ogden model rather than M o o n e y - R i v l i n model. Therefore the constitutive equation can be written as above with T to be obtained from: P  r  X  =  C  T  X  +  ^ X  <  -a,  [7-5]  and Tf from  •f  2  [7.32]  Finally, an additional equation for df should be formulated. The parameter £ depends on the kinetics o f the flow. This equation coupled with the Cauchy equation for flow and the constitutive equation developed in this chapter can be solved to predict important parameters such as pressure, velocity profiles and fibrillation profiles.  153  flow  CHAPTER 8 Extrusion of PTFE Blends and Effects of Various Additives 8.1 Introduction P T F E paste can be extruded at room temperature and the extrudates display good mechanical properties even without sintering, e.g. unsintered tape. During P T F E extrusion, the distribution o f the processing aid in the paste matrix confined in the barrel changes with the position and time due to the high pressure extrusion, especially when a long extrusion time is involved. In certain instances, the product can exhibit fractures from m i l d to severe depending on the processing conditions. To help minimize or eliminate these defects  in the extrudate,  several  alternative  processing aids are studied and proposed in this chapter. Moreover, blending o f P T F E resins is examined in order to exploit the possibility o f producing a resin that can be processed better than any o f its individual components. Another option to enhance processability is to utilize mixture o f lubricants in order to fix physical properties to certain values. Some processing aids such as fluoropolymers, boron nitride and nanoclays commonly used in the extrusion o f polyethylenes are tested here to determine their effect on P T F E paste extrusion.  8.2 PTFE blend extrusion Blends o f various types o f P T F E (homopolymer and copolymer) were prepared and extruded in order to study whether or not synergistic effects exist in fibril formation. In other words, it is desirable to study the effect o f blending on the mechanical properties o f extrudates and more specifically to determine i f these (mechanical properties) are better than those o f the individual components o f the blend when they are extruded under identical conditions. The various extrusion experiments were performed using a die having an entrance angle, 2a, o f 30°, a length-to-diameter ratio, L/D, o f 20, a reduction ratio, RR, o f 352 and different shear rates at 35°C. A s usual, all the extrudates were dried for 24 h at 120°C and the tensile strengths were measured in a Corn-Ten apparatus equipped with an 18 N load cell. During testing, the samples were stretched at a rate o f 13 mm/s.  154  The various blends were prepared i n two different ways. In the first method, pastes were prepared with each resin separately. These were subsequently mixed at a 1:1 ratio. In the second method, the powders were first mixed at a 1:1 ratio and subsequently the lubricant was added at the desired concentration. The results for a blend o f a homopolymer (F104 H M W ) with a copolymer (F301) are summarized i n Figures 8.1 and 8.2. The results for pastes prepared with the individual components o f the blend are also included for the sake o f comparison. First, the effect o f the method used to prepare the blend is insignificant and the differences are within the experimental error. Secondly, the extrusion pressures and the tensile strengths o f the blend are generally falling between those o f the pure components and closer to the component that exhibits the lower values. A similar blend was prepared with the two copolymers, F301 and F303, at a 1:1 ratio. Figures 8.3 and 8.4 depict the extrusion pressure and the tensile strength as a function o f the apparent shear rate, respectively. The results for pastes prepared with the individual components of the blend are also included. Once again the results are not surprising as both the extrusion pressure and tensile strength are falling between those o f the pure components and/or closer to that component that exhibits the lower values.  60  i  i i i I i—i i i | i • • • | '————I—•"" 2a=30°; L/D = 20; RR = 352:1; T = 35°C 1  1  1  i i I i i i  55 CO  0.  * 0)  50  3 CO  to  45  c o 'to  40  — • — F104 HMW-lsopar M + F301-Isopar M  3  i_  *J X  LU  •  35  — • — F104 HMW + Isopar M —A—  30 1000  F104 HMW & F301 + Isopar M  2000  3000  F301 + Isopar M 4000  _i l—i—i i i  5000  6000  7000  A p p a r e n t S h e a r R a t e (s' ) 1  Figure 8.1:  The extrusion pressure o f a blend o f a homopolymer (F104 H M W ) and a copolymer (F301) extruded at various apparent shear rates at 3 5 ° C . 155  7.0  i i i i I i i i i I i i i i | i  i i I i i i i I  2a=30°; L/D = 20; RR = 352:1; T = 35°C 6.5 6.0  co  Q.  5.5 5.0  c CD  4.5  CO  "35 c  - • — F104 HMW-lsopar M + F301-lsopar M  4.0  - • — F104 HMW & F301 + Isopar M  CD  - • — F104 HMW + Isopar M  3.5  A  F301 + Isopar M  i . . . . i—i i i i —  3.0 1000  2000  3000  4000  7000  6000  5000  A p p a r e n t S h e a r R a t e (s' ) 1  Figure 8.2:  The tensile strength o f a blend o f a homopolymer (F104 H M W ) and a copolymer (F301) extruded at various apparent shear rates at 35°C.  50  i i i i I i i i i I i i i i I i i i i I i i ' 2a=30°; L/D = 20; RR = 352:1; T = 35°C  o.  45 h  CD L.  3  (/) (/> CD  40 \-  CL C  o 3  35  — • — F301 -Isopar M + F303-lsopar M  X LU  30 1000  i  i  i  i i  2000  •  F301 + Isopar M  A-  F303 + Isopar M i  3000  4000  i  i  I i  6000  5000  i i  7000  Apparent Shear Rate (s' ) 1  Figure 8.3:  The extrusion pressure o f a blend o f two copolymers (F301 and F303) extruded at various apparent shear rates at 35°C. 156  7.0  I i i i i I i i i i I i i i i I i i i i — • — F301-Isopar M + F303-lsopar M  6.5  CL  — • — F301 + Isopar M •A  6.0  F303 + Isopar M  5.5  c d)  5.0  co  4.5  to c a)  4.0 3.5 2a=30 ; L/D = 20; RR = 352:1; T = 35 C  3.0 r i i i 1000  *  2000  3000  4000  •  •  •  1  •  •  1  *  6000  5000  •  7000  Apparent Shear Rate (s' ) 1  Figure 8.4:  The tensile strength o f a blend o f two copolymers (F301 and F303) extruded at various apparent shear rates at 35°C.  8.3 The effect of lubricant mixture on PTFE paste extrusion In a different type o f experiment, Isopar® G and Isopar® V were mixed at different weight percentages to see whether or not it is possible to fix the physical properties o f the lubricant at desired values, i.e. viscosity and surface tension. A l l physical properties o f these mixtures were measured at 25°C. Figures 8.5, 8.6 and 8.7 summarize the results for the density, viscosity and surface tension o f the mixtures, respectivelly. W h i l e the density changes in a linear fashion, the other two physical properties change nonlinearly. A paste o f F104 H M W was prepared using a mixture o f 60 wt% o f Isopar® G with 40 wt% o f Isopar® V . Figures 8.8 and 8.9 show the results on the extrusion pressure. A s previously, the results for the blend are falling between those o f the individual pure components or the effect is very small and is within experimental error.  157  0.85 \-  0.70 ' 0  1  1  1  20  40  60  I 100  1_  80  wt% of Isopar V  Figure 8.5:  The density of mixtures of Isopar® G and Isopar® V as a function of composition at 25°C.  o  1  0  1  20  —  1  1  40  60  '  80  100  wt% of Isopar V  Figure 8.6:  The viscosity o f mixtures o f Isopar® G and Isopar® V as a function o f composition at 25°C.  158  40  20  100  60  wt.% of Isopar V  Figure 8.7  The surface tension o f mixtures o f Isopar® G and Isopar®. V as a function o f composition at 2 5 ° C . ,  80 —•—F104 HMW + Isopar G •—• F104 HMW + Isopar V  co  A  F104 HMW + Isopar G-V  •  70 •  0  1_  3 10  (A CO  -  60 ..A  c o  to 3 s_ X LU  50  •  *  A  -•  '  ^ ^ ^ ^ ^  40 1000  2000  • •  3000  2a=30°; L/D = 20; RR = 352:1; T = 35°C ' . . i . . . . i . . 7000 6000 5000 4000  A p p a r e n t S h e a r R a t e (s*)  Figure 8.8:  Steady state extrusion pressure o f a paste prepared with homopolymer F104 H M W and a mixture o f Isopar G & V at 35°C.  159  8.0  i  7.5 7.0 CL  i  I  i  i  i  i  I  i  i  i  i  •—F104 HMW + Isopar G • — F104 HMW + Isopar V F104 HMW + Isopar G - V  6.5 6.0  c  5.5  2 53  5.0  "55 c  4.5 i4.0 3.5 '3.0 1000  2a=30 ; L/D = 20; RR = 352:1; T = 35 C *  *  *  1  *  2000  •  * • • • • * • * • • • * * * 6000 3000 4000 5000 1  1  1  1  7000  Apparent S h e a r Rate (s')  Figure 8.9:  Tensile strength o f a paste prepared with homopolymer F104 H M W and amixture ofIsoparG&Vat35°C.  8.4 Effects of additives on paste extrusion Pastes o f F104 H M W were prepared with the addition o f boron nitride ( B N 2 , C T F 5 ) or nanoclays (Nanomer® I.44P, Nanomer® P G V ) and Isopar® M as a lubricant. The blends were prepared b y mixing first the two solid phases ( P T F E and either boron nitride or nanoclay) i n the desired proportions prior to the addition o f the lubricant. In all cases the lubricant concentration was kept at 18.7 wt%. The results are depicted in Figures 8.10 to 8.15. The results o f the corresponding virgin paste (no additives) are also included for the sake o f comparison. First, the addition o f 5wt% B N 2 increases the extrusion pressure significantly (Figure 8.10). Surprisingly, this increase in pressure has a positive and significant effect on the tensile strength o f the final dried extrudate. The addition o f 2 wt% o f B N 2 also increases the extrusion pressure at higher shear rates, but the tensile strength does not increase significantly. O n the other hand, the addition o f 2wt% o f C T F 5 increases the extrusion pressure, although the tensile strength does not increase accordingly. Finally, addition o f 5wt% o f C T F 5 decreases the extrusion pressure possibly by enhancing slip effects.  160  75  i  i  i  i  I  i  i  i  i  I  i  i  i  i  I  i  i  i  i  I  i  i  i  i  2ct=30°; L/D = 20; RR = 352:1; T = 35°C  I  i  i  •  70 CO  Q.  65  •—  3 CO CO  60  55 C  o  50  '(/> 3  X LU  h  — • — F104 HMW + Isopar M  45 t-  A F 1 0 4 HMW & 2% BN2 +Isopar M  j  — • — F104 HMW & 5% BN2 + Isopar M  J  — • — F104 HMW & 2% CTF5 + Isopar M I  40 1000  i  — T — F104 HMW & 5% CTF5 + Isopar M * • • • • * * * • * • • • 7000 5000 6000 4000 1  i  3000  2000  1  1  A p p a r e n t s h e a r rate (s~)  Figure 8.10:  Extrusion pressure o f F104 H M W + B N blends extruded at 35°C  8.0  I  I  I  I  I  |  I  I  I  I  |  •  I  •  I  2 a = 3 0 ° ; L/D = 20; RR = 352:1; T = 35°C  7.5  co CL  |  7.0 6.5  c  6.0  co  5.5  'to C  F104HMW+ Isopar M  5.0  F104 HMW & 2% BN2 + Isopar M F104 HMW & 5% BN2 + Isopar M  4.5  F104 HMW & 2% CTF5 + Isopar M  •  F104 HMW & 5% CTF5 + Isopar M  A  4.0 1000  •  * * * • • * • • • • 4000 3000 2000 1  1  1  5000  6000  7000  A p p a r e n t S h e a r R a t e (s")  Figure 8.11:  Tensile strength o f F104 H M W + B N blends extruded at 35°C.  161  The addition o f nanoclays, Figures 8.12 to 8.15, has a similar effect as that o f B N 2 . In fact, the blend at 5% o f nanoclay extrudes at significantly higher pressure and the extrudates exhibit higher tensile strengths compared to those o f the pure resin. It is worthwhile to notice thafeeven though the extrusion pressure increases with the inclusion o f the additives, there is no effect on the appearance o f the extrudates. In other words, that the extrudates do not exhibit any visible fracture or surface defects.  70  i  i  I  i  i  i  i  I  i  i  T  - F 1 0 4 HMW + Isopar M F104 HMW & 2% I.44P + Isopar M A CO  0_  <D v. 3 (A  in co i_  F104 HMW & 5% I.44P + Isopar M  65 \-  60  Q_ C  o "55 '  55  3 HI  50 1000  2000  3000  2a=30°; L/D = 20; RR = 352:1; T = 35°C • • • • • 6000 7000 5000 4000  A p p a r e n t S h e a r R a t e s (s* ) 1  Figure 8.12:  Extrusion pressure o f F104 H M W + Nanomer® I.44P blends extruded at 35°C.  162  8.0  i  i  i  I  i  i  |  I  I  I  I  |  I  I  I  I  |  I  I  I  I  F104 HMW + Isopar M 7.5  F104 HMW 8. 2% I.44P + Isopar M  h  F104 HMW & 5% I.44P + Isopar M  (0  CL  7.0 p  C  6.5  h  6.0  h  2> <-»  w  _o '35 c CD H  5.5 \2oc=30 ; L/D = 20; RR = 352:1; T = 35 C • 5.0 1000  *  •  •  1  •  •  •  •  2000  1  •  *  •  *  3000  1  •  •  *  *  4000  1  *  *  6000  5000  7000  A p p a r e n t S h e a r Rate ( s ) 1  F i g u r e 8.13:  Tensile strength o f F104 H M W + Nanomer® I.44P blends extruded at 35°C.  60 2a=30°; L/D = 20; RR = 352:1; T = 35°C (0  CL  1< Oi  55 h  /  v_ 3  CO CO  /  <D  CL C  o  50  h  A''  /  •  "35 3  X LU  — • — F104 HMW + Isopar M . — • — F104 HMW + 5% PGV A  1000  2000  3000  4000  F104 HMW +2% PGV 5000  6000  . . 7000  A p p a r e n t S h e a r R a t e (s' ) 1  F i g u r e 8.14:  Extrusion pressure o f F104 H M W + Nanomer® P G V blends extruded at 35°C.  163  7.5  i  i  i  •  I i  •  i  i  I  i  i  i  i  I  i  i  i  •  I  i  i  •  i  I  i  i  i  i  2ct=30°; L/D = 20; R R = 352:1; T = 35°C 7.0 CO QL  6.5  6.0  O)  c  CO  c  5.5  _co  5.0  -*-» CO  '55 c  CD  4.5  h  F104 HMW + Isopar M F104HMW + 5 % P G V F104HMW + 2 % P G V  4.0 1000  2000  3000  4000  6000  5000  7000  Apparent Shear Rate (s ) 1  Figure 8.15:  Extrusion pressure o f F104 H M W + Nanomer® P G V blends extruded at 35°C.  Figures 8.16 and 8.17 compare the effect o f all the additives at different concentrations. Boron nitride seems to have a stronger effect on paste extrusion than that o f the clays. 75  •  i  I i  I  '  1  '  [ 2 a = 3 0 ° ; L/D = 20; R R = 352:1; T = 35°C 70 LO-  C O 1_ 3 CO  V) CO  65  60 • — Isopar M  c  o  2% BN2  55 k  O—5%BN2 " A  'co 3 i— -*-» X LU  50  h  45 1000  2% CTF5  &  5% CTF5  • -  2% I.44P ,  V — 5% I.44P  2000  3000  4000  5000  •  2% PGV  O  5% PGV  6000  7000  A p p a r e n t S h e a r R a t e (s~)  Figure 8.16:  The effect o f B N and clay addition on the extrusion pressure o f F104 HMWpastes. 164  Figure 8.17:  The effect o f B N and clay addition on the tensile strength o f F104 H M W pastes.  Figures 8.18 and 8.19 depict the extrusion pressure and tensile strength o f extrudates obtained from blending o f F104 H M W with boron nitride or nanoclay and Isopar® G . The effect is essentially the same as was previously observed in Figures 8.10 to 8.13. The blend with Nanomer® I.44P extrudes at higher pressures and its tensile strengths increase accordingly. In all other cases the tensile strength decreases. Table 8.1 lists the mechanical properties o f these extrudates at the shear rate o f 5859 s ' . The tensile strength is generally lower than that o f the paste with no additive. The elastic modulus o f the blends with I.44P exhibit higher values, but their extensibility decreases. The addition o f C T F 5 at any concentration also decreases the elastic modulus o f the extrudate. However, addition o f 5 wt% C T F 5 slightly decreases the elastic modulus while its extensibility increases significantly. In general, it can be concluded that the incorporation o f additives such as clays and boron nitride into the P T F E matrix, causes the extrudate to behave as a more brittle material. O n the other hand, the addition o f a specific type o f boron nitride (BN2) was found to increase significantly both the extrusion pressure and the tensile strength o f the P T F E extrudates.  165  Table 8.1:  Mechanical properties o f PTFE-additive blends + Isopar conditions.  G extruded at different  Tensile Strength % Elongation at Elastic Modulus (MPa) Break (MPa) F104 HMW + Isopar® G, 2a = 30°, L/D = 20, RR = 352:1, y = 5859 s 97.0(5.5) 90.5 (5.4) 5.6 (0.2) No Additive 71.9(5.7) 5.4 (0.2) 90.8 (7.9) CTF5 @ 2 wt% 122.5 (10.8) 5.1 (0.1) 81.8(6.5) CTF5 @ 5 wt% 74.6 (2.3) 103.1 (4.1) 5.5 (0.1) I.44P @ 2 wt% 104.9(2.1) 59.3 (5.8) 5.8 (0.2) I.44P @ 5 wt% Additive  A  65  i  i  i  i  I  i  i  i  i  I  i  i  i  i  I  i  i  i  i  i  I  i  i  I  i  i  i  2 a = 3 0 ° ; L/D = 20; RR = 352:1; T = 35°C  CO Q.  £  55  3 CO  to o 0.  c  50  o "(0  3  P.  45  40 1000  • •  I •  2000  3000  4000  F104 HMW + Isopar G F104 HMW & 2% I44P + Isopar G F104 HMW & 5% I44P + Isopar G F104 HMW & 2% CTF5 + Isopar G F104 HMW & 5% CTF5 + Isopar G • * * * * * * * * * * * *  5000  6000  7000  A p p a r e n t S h e a r Rate ( s ) 1  Figure 8.18:  Extrusion pressure o f blends of F104 H M W + Isopar® G and Boron Nitride.  166  8.0  •  I  '  •  •  '  I  I  1  i  i  i  i  I  i  i  i  i  I  i  i  i  i  — • — F104 HMW + Isopar G  7.5 CO Q.  •  F104 HMW & 2% I44P + Isopar G  - A  F104 HMW & 5% I44P + Isopar G  — T — F104 HMW & 2% CTF5 + Isopar G  7.0  —•-  F104 HMW & 5% CTF5 + Isopar G  6.5 C  6.0  :  co 5.5 "55 c  5.0  CO  4.5  2a=30 ; L/D = 20; RR = 352:1; T = 35 C  • • • • • • • • • * • • • • • 4.0 4000 2000 3000 1000 1  1  1  1  *  5000  *  *  6000  7000  A p p a r e n t S h e a r R a t e (s~)  Figure 8.19:  Tensile strength o f blends o f F104 H M W + Isopar® G and Boron Nitride.  8.5 Summary A series o f processing aids different from the isoparaflnnic lubricants commonly used have been tested in order to identify alternative enhanced processing aids. In addition, blends o f different P T F E fine powder resins have been prepared and extruded in an attempt to improve the P T F E paste extrusion process. The pastes and blends prepared in the regular way were extruded under conditions that have been found to be optimum for the P T F E paste with isoparaflnnic lubricants and their mechanical properties were measured. First, 1:1 mixtures o f P T F E resins (homopolymer + copolymer and copolymer + copolymer) were made and this was mixed with the lubricant to prepare the paste. In a second method, the respective pastes were first prepared prior to blending. Regardless o f the way the blend was prepared, it was found that there is no significant effect on the extrusion conditions and the mechanical properties o f the extrudates. In fact, these properties fall between those exhibited by the pure components. One possible reason for this behaviour is that all the resins used in this work have very similar particle size distributions. In a different experiment, a mixture o f two isoparaflnnic lubricants (having different physical properties) were prepared and then incorporated with the resin to make up the paste. 167  Once again, the extrusion conditions and mechanical properties o f the resulting extrudates fell between those o f the pure components. The incorporation o f additives into the resin-lubricant matrix yielded promising results. These additives are commonly used as polymer processing aids to reduce or eliminate surface fracture. The additives were added in different proportions, but the lubricant concentration was kept always constant, namely at about 18.7 wt% o f lubricant. It was found that in the case o f boron nitride, the presence o f the additive increased the extrusion pressure and significantly improved the mechanical properties o f the extrudates. However, not all the additives exhibited a beneficial effect. The addition o f clay was found to decrease the elastic modulus o f the extrudates and to increase the extensibility significantly.  168  CHAPTER 9 Conclusions, Recommendations and Contribution to knowledge 9.1 Conclusions The rheological properties o f a number o f polytetrafluoroethylene (PTFE) pastes were studied relevant to paste extrusion process. The effects o f the physical properties o f lubricants and the geometrical characteristics o f the extrusion die on the extrusion pressure and mechanical properties  o f the  final  extrudates were  also examined. L i q u i d migration and  preform  inhomogeneities were studied as functions o f the physical properties o f the lubricants, and the level and duration o f pressure applied. First, a number o f lubricants were identified as suitable processing aids for the paste extrusion o f polytetrafluoroethylene (PTFE). They were characterized in terms o f both flow and surface properties. It was found that it is possible to alter the flow and surface properties o f these lubricants independently. Thus, it became possible to study their relative effects on preforming and P T F E paste extrusion. Based on this study, it was concluded that preforming quality increases with increase o f lubricant viscosity and with improvement in the wettability characteristics o f the lubricant with P T F E . However, lubricant migration becomes important at longer times. Therefore, the applied pressure and its duration need to be optimized depending on the physical properties o f lubricant. It was also found that the physical properties o f the lubricants (viscosity and surface tension) play a significant role. First, the use o f a lubricant having a higher viscosity can produce a more uniform preform as liquid migration is minimal. In addition, increasing the wettability o f lubricant with P T F E produces better mixture/pastes. Furthermore, this has an effect on the preform preparation. Excellent wetting would produce a uniform preform even under extreme conditions. The physical properties o f lubricants were found to play a significant role in the extrusion o f P T F E pastes. Increasing the wettability o f lubricant with P T F E and decreasing the lubricant viscosity causes a reduction in the extrusion pressure and an increase in the tensile strength o f the extrudates. It was also found that the geometrical characteristics o f dies also play a significant role in the process o f paste extrusion. More specifically, the extrusion pressure increases with L/D ratio and reduction ratio, while it exhibits a minimum with the contraction angle at about 60°. These effects are also reflected in the mechanical properties of the  169  extrudates. The tensile strengths o f dried and sintered extrudates exhibit optimum values at an L/D equal to about 20, and a reduction ratio equal to about 150, whereas they exhibit a minimum value at about 60°. Finally, it was found that a higher molecular weight resin produces stronger fibrils that account for the mechanical superiority o f the final extrudate. Therefore, i f the objective is to prepare a product having a high tensile strength, a homopolymer o f high molecular weight should be used in combination with a lubricant having a low surface tension such as H F E - 7 5 0 0 and minimum viscosity up to a certain value. T o low a viscosity would produce a nonuniform preform. A s far as die selection concerns, this should have an L / D ratio o f about 20, a reduction ratio o f around 150 and a contraction angle o f no more than 60°. The rheology o f P T F E paste was found to be complex in relation to the paste extrusion process. The paste starts as a two phase fluid and behaves as a semisolid at the exit o f the die due to the formation o f fibrils. The rheology o f the initial state and final states were studied in detail. The initial oversaturated suspension was found to exhibit yield stress which although small was measureable. Once this yield stress is exceeded, the structure breaks and flow is initiated. The viscosity then exhibits a shear thinning behaviour. The rheology o f the final state can be approximated by an Ogden elastic body at least based on the available data. Plasticity effects are present, although it is believed that during extrusion due to the presence o f lubricant large deformations o f fibrilated zones are prevented i.e. perhaps due to significant slippage. Overall it was concluded that the rheology can be represented by a constitutive equation having the following form: * where T is the overall stress, ^  p  = .0 " £ ) * p  +  [7.1]  is the contribution to the stress tensor from the presence o f  unfibrillated particles (initial state o f the oversaturated suspension), T  f  the contribution to the  stress tensor from the presence o f the fibrillated particles and £, represents the fraction o f the fibrillated particles with £ = 0 representing an unfibrillated paste i.e. just preformed before subjected to flow and £ = 1 representing a fully fibrillated paste. This constitutive equation can be used for flow simulations provided that an equation for the dynamics o f £ is also developed. Regarding the use o f additives, different processing aids used with molten polymers were tested. Promising results are reported in this work. Due to the presence o f these additives the steady-state extrusion pressures experience an incremease. However, the mechanical  170  properties o f the extrudates are also affected in various manners. The presence o f boron nitride increases the tensile strength o f the extrudate wheeras the presence o f clay decrease their elastic modulus while increasing their extensibility.  9.2 Contributions to Knowledge Several novel contributions to knowledge have resulted from this research work. These are identified as follows. 1. The current commercial procedure for P T F E paste extrusion has been analyzed in detail, and the physical significance o f each experimental aspect o f the process has been investigated. Specific variables have been explored in an attempt to improve process efficiency. These are more related to the type o f lubricant recommended for use. 2.  The preforming behavior o f P T F E pastes has been studied as a function o f the physical  properties o f the lubricants. The results provide an understanding on how various operating variables affect the quality o f P T F E preforms. 3.  The rheological behavior o f P T F E pastes has been studied experimentally at both states  before and after the extrusion. Effects o f various operating variables have been determined and discussed. 4. The effects o f various operating variables on the quality o f P T F E paste extrudates have been analyzed. The results provide an understanding o f the role o f fibrillation in defining the final product properties. W i t h such understanding, it is possible to optimize the' extrusion operating variables in order to produce extrudates that are commercially acceptable, with the ultimate objective o f reducing the amount o f process rejects. The results can also be used as a basis for resin selection, for a desired particular end use. Overall this work has contributed to the fundamental  knowledge o f P T F E paste  extrusion, which is still in its infancy as far as research is concerned. Undoubtedly, more indepth studies need to be performed in the future to completely unravel the science behind the process. However, many o f the findings in this work have provided significant initial steps towards a better macroscopic and microscopic understanding o f the process and, therefore, allowed the commercial implementation o f P T F E paste extrusion to be carried out with greater confidence. The introduction o f the slit die to produce rectangular-shape  extrudates has  facilitated the extensional rheology o f the samples allowing a different approach in this area.  171  9.3 Recommendations for Future Work Several important aspects o f P T F E paste extrusion are yet to be studied. These are recommended below, as possible objectives for future research work. 1.  The one-dimensional mathematical model developed by Ariawan (2000) to predict the  flow o f P T F E pastes is dependent on several material constants which can be studied by using rheological experiments. This way the model can be validated and become more useful. Inconsistencies resulting from this exercise can be corrected by modifying the proposed model. 2.  In order to make the rheological study o f P T F E pastes more complete, the effects o f  other variables, such as resin particle size and its distribution ( i f it is indeed possible to vary in practice) should be investigated in the future. The flow o f P T F E pastes through a hyperbolic die and a crosshead die for a wire coating process are also interesting and commercially useful. 3.  F l o w simulations should be performed by utilizing the proposed constitutive equation.  This way the usefulness  o f this model can be validated and perhaps this w i l l lead to  development o f other advanced models. 4.  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The parameters are found by u s i n g simplex % method c a l l e d by N_M f u n c t i o n format s h o r t g D = csvread('hel_13.txt'); % I n i t i a l values x = D(:,l); y = D(:,2); np = 2; q= (sqrt(np+1)-1)/(np*sqrt(2) p = q + 1/sqrt(2) ; K = [0 0;p q;q p ] ; [kopt,stage,k_plot,y_plot] =  % % % )  Shear s t r e s s (MPa) Extensional r a t i o , alpha Number of parameters ;  N_M(K,x,y);  % C a l c u l a t i n g the f i t t e d f u n c t i o n v a l u e y _ c a l = f e v a l ( go' ,kopt,x) ; 1  % P r i n t i n g out the r e s u l t s fprintf('\n Nelder-Mead Simplex Method \ ) fprintf('\n _ \n' fprintf('\n Parameter Values to Minimize \n' I fprintf('\n the O b j e c t i v e F u n c t i o n \ ) fprintf( \n' ) fprintf( \n fprintf('\n \n ,kopt (1)) %7 .4f m fprintf( \n \n ,kopt (2)) %7 .4f n \n' ) fprintf('\n fprintf('\n %3 . Of \n' , stage) Number of I t e r a t i o n s \n' fprintf('\n fprintf('\n Minimum v a l u e of the V) fprintf('\n %6 . 3f Objective Function \ ,y_plot(length(y_plot))) \n' fprintf('\n fprintf('\n Optimum Value of the \') %6 . 3f fprintf( \n Function \',y_cal(length(y))) \n' ) fprintf('\n 1  1  1  1  1  1  1  % P l o t i n g the r e s u l t s n = 1:stage; subplot(2,2,1); plot(n,k_plot(n,1),'b ,n,k_plot(n,2),•r ) t i t l e ( ' V a l u e of parameters at each i t e r a t i o n ' ) legend('m','n') xlabel('Number of i t e r a t i o n s ' ) ylabel('parameters ' ) 1  1  subplot(2,2,2) ; plot(n,yjplot) t i t l e ( ' V a l u e of the O b j e c t i v e F u n c t i o n at Each I t e r a t i o n ' ) xlabel('Number of I t e r a t i o n s ' ) y l a b e l ( ' V a l u e of the O b j e c t i v e Function') subplot(2,2,3) ; p l o t ( x , y , ' r ' , x , y _ c a l , 'b ' )  179  legend('Observed D a t a ' , ' F i t t e d Data') t i t l e ( ' C o m p a r i s o n Between Observed and F i t t e d x l a b e l ( ' a l p h a = L/LO') y l a b e l ( Sigma (MPa) ) 1  Data ) 1  1  % N_M.m i s program t o use Nelder-Mead (simplex) method % t o f i n d the parameters t h a t b e s t f i t e x p e r i m e n t a l d a t a function  [kopt, stage, k j p l o t , y_jplot] = N_M(K,x,y)  % D e f i n i n g the Nelder-Mead c o e f f i c i e n t s f o r the b a s i c o p e r a t i o n s alpha = 1 ; % Reflection coefficient betha =0.5; % Contraction coefficient gamma = 2 ; % Expansion c o e f f i c i e n t  n = size(K,2); m = n + 1;  % Number of independent v a r i a b l e s i n the o b j e c t i v e % Number of v e r t i c e s of the p o l y h e d r o n  % Nelder-Mead method stage = criteria epsilon nstages  function  starts  0; = 1; = le-6; = 3000;  while (stage < nstages) & ( c r i t e r i a > e p s i l o n ) % E v a l u a t i n g the f u n c t i o n i n each v e r t i x of the f o r i = l:m S ( i ) = f e v a K ' f o ' , K ( i , :) ,x,y) ; end % S e t t i n g the sh = S(1);h = ss = S(1);s = s i = S(1);1 =  initial 1; kh = 1; ks = 1; k l =  simplex  optimum v a l u e s t o f i n d those t h a t o p t i m i z e the f u n c t i o n % Temporary maximum K(h, % Temporary second maximum K(s, K(l, kopt = k l ; % Temporary minimum  % F i n d i n g the maximum and minimum v a l u e of the f o r i = 2:m i f S ( i ) > ss i f S ( i ) > sh S S = sh; S = h; ks = K ( s , : ) ; sh = S ( i ) ; h = i ; kh = K ( h , : ) ; else ss = S (i) ; S = i ; ks = K(s, :) ; end e l s e i f S(i) < s i s i = S (i) ; 1 =i ; k l = K ( l , :),; kopt = k l ; end end k_plot(stage+1,:) = kopt; y_plot(stage+1) = s i ;  180  function  % C a l c u l a t i n g the c o o r d i n a t e s of the c e n t r o i d o f the simplex for i = l : n sumk = 0; f o r j = l:m sumk = sumk + K ( j , i ) ; end k o ( i ) = (sumk - k h ( i ) ) / n ; end % R e f l e c t i n g the xh p o i n t  kr = ko + a l p h a * (ko - k h ) ; through the c e n t r o i d sr = f e v a l ( ' f o ' , k r , x , y ) ; at the r e f l e c t i n g p o i n t s  % and e v a l u a t i n g  if  sr < s i ke = gamma * k r + (1-gamma)*ko; se = f e v a l ( ' f o ' , k e , x , y ) ; the expanding p o i n t s i f se < s i kh = ke; K(h,:) = kh; sh = se; else kh = k r ; K(h,:) = kh; sh = s r ; end e l s e i f s r < ss kh = k r ; K(h,:) = kh; sh = s r ; else i f s r < sh kh = k r ; K(h,:) = kh; sh = s r ; kc = betha * kh + ( l - b e t h a ) * k o ; sc = f e v a l ( ' f o ' , k c , x , y ) ; the c o n t r a c t i n g p o i n t s i f sh > sc kh = kc; K(h,:) = kh; sh = S C ; else f o r i = l:m K(i,:) = (K(i,:)+kl)/2; end end else f o r i = l:m K(i,:) = (K(i,:)+kl)/2; end end end  the f u n c t i o n  % Expanding the p o l y t o p e % E v a l u a t i n g the f u n c t i o n a t  % Contracting % Evaluating  % Shrinking  the p o l y t o p e the f u n c t i o n a t  the p o l y t o p e  % C a l c u l a t i n g the c r i t e r i a of converge sumconv = 0; for  i = l:m sumconv = sumconv + ( f e v a l ( ' f o ' , K ( i , : ) , x , y ) - f e v a l ( f o , k o , x , y ) ) " 2 ; 1  end c r i t e r i a = (sumconv/ra) 0.5; stage = stage + 1; kopt = k l ; A  181  1  end  function z = f(k,x,y) n = length(x); sum = 0 ;  for  i = l:n T i l = k ( l ) * ( x ( i ) ( k ( 2 ) -1) - x ( i ) * - ( l + k ( 2 ) / 2 ) ) ; sum = sum + ( y ( i ) - T i l ) * 2 ; A  end z =  sum;  f u n c t i o n z = g(k,x) n = length(x); for  i = l:n z(i) = k ( l ) * ( x ( i ) ( k ( 2 ) - 1 ) - x ( i ) - ( l + k ( 2 ) / 2 ) ) ; A  A  end  182  

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