The Complete Part Design Handbook: For Injection Molding Of Thermoplastics

E. Alfredo Campo The Complete Part Design Handbook For Injection Molding of Thermoplastics Hanser Publishers, Munich ...

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E. Alfredo Campo

The Complete Part Design Handbook For Injection Molding of Thermoplastics

Hanser Publishers, Munich • Hanser Gardner Publications, Cincinnati

The Author: E. Alfredo!Campo,!1213!Cerrito!Perdido!Lane!,!El!Paso,!TX!79912,USA Distributed!in!the!USA!and!in!Canada!by Hanser!Gardner!Publications,!Inc. 6915!Valley!Avenue,!Cincinnati,!Ohio!45244-3029,!USA Fax:!(513)!527-8801 Phone:!(513)!527-8977!or!1-800-950-8977 www.hansergardner.com Distributed!in!all!other!countries!by Carl!Hanser!Verlag Postfach!86!04!20,!81631!München,!Germany Fax:!+49!(89)!98!48!09 www.hanser.de The!use!of general!descriptive!names, trademarks, etc., in!this!publication, even!if the! former! are! not! especially! identiied, is! not! to! be! taken! as! a! sign! that! such! names, as understood!by!the!Trade!Marks!and!Merchandise!Marks!Act, may!accordingly!be!used! freely!by!anyone. While!the!advice!and!information!in!this!book!are!believed!to!be!true!and!accurate!at!the date!of going!to!press, neither!the!authors!nor!the!editors!nor!the!publisher!can!accept any! legal! responsibility! for! any! errors! or! omissions! that! may! be! made. The! publisher! makes!no!warranty,!express!or!implied,!with!respect!to!the!material!contained!herein. Library!of!Congress!Cataloging-in-Publication!Data Campo,!E. Alfredo. !!The!complete!part!design!handbook!:!for!injection!molding!of!thermoplastics /!E. Alfredo!Campo. !!!!!!!p.!cm. !!Includes!index. !!ISBN-13:!978-1-56990-375-9 !!ISBN-10:!1-56990-375-1 !1.!!Injection!molding!of!plastics--Handbooks,!manuals,!etc.!2. Thermoplastics--Design--Handbooks,!manuals,!etc.!3.!!Plastics--Handbooks, manuals,!etc.!!I.!Title. !!TP1150.C36!2006 !!668.4‘23--dc22 !!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!2006010219 Bibliograische!Information!Der!Deutschen!Bibliothek Die!Deutsche!Bibliothek!verzeichnet!diese!Publikation!in!der!Deutschen! Nationalbibliograie;!detaillierte!biblio-graische!Daten!sind!im!Internet!über!!abrufbar. ISBN-10:!3-446-40309-4 ISBN-13:!978-3-446-40309-3 All!rights!reserved.No!part!of this!book!may!be!reproduced!or!transmitted!in!any!form!or! by!any!means, electronic!or!mechanical, including!photocopying!or!by!any!information storage!and!retrieval!system,!without!permission!in!writing!from!the!publisher. ©!Carl!Hanser!Verlag,!Munich!2006 Production!Management:!Oswald!Immel Typeset!by!Manuela!Treindl,!Laaber,!Germany Coverconcept:!Marc!Müller-Bremer,!Rebranding,!München,!Germany Coverdesign:!MCP!•!Susanne!Kraus!GbR,!Holzkirchen,!Germany Printed!and!bound!by!Druckhaus!“Thomas!Müntzer”!GmbH,!B.!Langensalza,!Germany

V

Dedication To my wife Sandy, my son Jaime, and my daughter Michelle for the love and understanding they gave me during the six long years that it took me to produce all the original drawings, calculate examples and write this comprehensive handbook. For my uncle, Captain Jaime Merchán and my aunt Ruth Reichman for the assistance provided while attending the University of Miami. For my great grand uncle Rafael Ramirez who paid for my elementary and high school education after I lost my father at the age of ﬁve. In honor of my beloved departed brother, Leoncio, and sister in-law Becky. For my grand nephew, Alfredo Campo V. for winning the BMX world championship and the sportsman of the year in my native country Ecuador. For Richard Scott, Bob Rackley and Bill Hawkins from the Du Pont Film Department for their leadership, teaching and guidance in how to be successful in the engineering technical research and development ﬁelds. For Larry Gillespie, Director of Du Pont Engineering Polymers for providing me the opportunity to work in most of the technical positions in plastics, such as development of new compounds, economic project analysis, designs of international compounding facilities and laboratories in Japan, Mexico and Brazil. For the Du Pont Film Department, Textile Fibers Department, Engineering Polymers Department, and Du Pont do Brazil that provided me several technical assignments in Plastics during thirty years of service to the company. For all the Du Pont Plastics Customers that gave me the opportunity to improve their plastic products and optimize their injection molding manufacturing processes. E. Alfredo Campo

VII

Preface This handbook was written for the injection molding product designer who has a limited knowledge of engineering polymers. It is a guide for the designer to decide which resin and design geometries to use for the design of plastic parts. It can also offer knowledgeable advice for resin and machine selection and processing parameters. Manufacturer and end user satisfaction is the ultimate goal. This book is an indispensable, all inclusive, reference guide that can be used by any plastic product designer. There is no need to search through many books and catalogs for needed information. New illustrations, graphs and equations have been included to provide additional clarity for complex ideas. The equations have been veriﬁed to ensure correctness and not just copied from another source. Thousands of hours of research and cross referencing have gone into the completion of this work. In addition, more than 35 years of the “hands-on” experience of a plastics expert have been incorporated in this handbook. The following topics are covered: Chapter 1

Plastic Materials Selection Guide: Includes an introduction to plastic materials, the beginning of plastics, classiﬁcation of polymer families. Each resin is discussed by its basic chemistry, properties, processing characteristics, advantages, disadvantages and limitations, typical applications and several product illustrations. Thermoplastic materials (35 generic families), thermoplastic elastomer materials (8 generic families), liquid injection molding of silicone, thermoset materials (16 generic families).

Chapter 2

Engineering Product Design: Starts with the introduction to structural product design principles, mechanical strength properties of thermoplastics. Centroid, section area, moment of inertia equations and tables. Beam deﬂection analysis methods. Structure analysis of beams, columns, ﬂat circular plates, and torsion.

Chapter 3

Structural Design for Thermoplastics: Discusses the product wall thickness, structural rib design, sharp corners, bosses, threads, undercuts, integral life hinges, pin hinges. Encapsulation of inserts, types of metal inserts and anchorage, and electrical lead inserts.

Chapter 4

Thermoplastic Gearing Design: An introduction to and classiﬁcation of gears. Standard spur, helical, bevel, and worm gears; properties required for thermoplastic molded spur gears, mounting gears on metal shafts, tolerances and mold shrinkage of gears. Plastic spur and helical gearing technology design, strength, horsepower rating, equations, tables, analysis examples and gear speciﬁcation illustrations.

Chapter 5

Plastic Journal Bearing Design: An introduction to types of materials for journal bearings. Theory and design for lubrication. Design principles, performances, dimensions, clearances, molding effects, PV limits and surface ﬁnishing. Self-lubricated thermoplastic bearings. Equations, tables, and analysis examples.

VIII

Preface Chapter 6

Thermoplastic Spring Design: Introduces cantilever beam spring design, applications, and analysis examples. Locating, ﬁxing clip, ﬂexible hinges, and torsional spring applications. Belleville spring washers’ equations, tables, and analysis examples.

Chapter 7

Thermoplastic Pressure Vessel Design: Discusses thin- and thick-walled pressure vessels’ basic principles, equations, tables, analysis examples, design guidelines, applications, and pressure vessel regulations.

Chapter 8

Thermoplastic Assembly Methods: Joining two or more components together: assembly method is selected based on product design geometry, size, end use requirements, thermoplastic material characteristics, automatic or manual assembly operation, and manufacturing costs. Each assembly method provides a description, process sequence, advantages and limitations, typical applications, equipment, product joint design, and its variations.

Chapter 9

Thermoplastic Effects on Design: Starts discussing the polymer melt behavior, reinforcement, degradation, moisture characteristics, mold shrinkage and critical properties. The molding process effects caused by molding cycle, melt/mold temperature, injection pressure and speed, etc. on product design dimensions, surface ﬁnishing, weld line strength and impact resistance and other molding problems.

Chapter 10

Thermoplastic Injection Mold Design: Provides an introduction of injection molds, classiﬁcation and effects on product design. Types of steels, chemical composition, effects of alloying, heat treatment, properties and characteristics. Types of steels used for mold bases and mold components. Cavity surfaces ﬁnish procedures and speciﬁcations. Types of injection mold designs. Cold runners (two- and three-plate molds, interchangeable mold inserts and vertical insert encapsulation mold). Hot runner molds (internally and externally heated, insulated). Mold design system and other considerations, such as number of cavities, parting line, ejection, cooling, cold runner, gating, venting, cavity inserts sidewall strength, support pillars, molded parts tolerances, mold designer check list, general speciﬁcations for mold construction are covered.

Chapter 11

Performance Testing of Thermoplastics: It introduces various tests to which thermoplastic polymers are subjected, describes their properties (statistical analysis), such as mechanical, thermal, chemical resistance, rheometer melt viscosity, soldering heat resistance, electrical, ﬂammability, smoke generation, weathering and micro-organism resistance. Test description, procedures, apparatus, test specimen and conditioning, and their signiﬁcance are discussed here.

Chapter 12

Thermoplastic Product Cost Analysis: It discusses molding process variables and capital equipment cost. Three cost analysis methods are used to estimate the molded product user’s price.

IX

Contents Preface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . VII 1 Polymeric Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 1.1 Introduction to Plastic Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 1.1.1 Beginning of Plastics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 1.1.2 Polymer Families . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 1.2 Thermoplastic Polymers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4 1.2.1 Classiﬁcation of Polymers by Performance . . . . . . . . . . . . . 4 1.2.2 Molecular Structure of Plastic Materials . . . . . . . . . . . . . . . 6 1.2.3 Acrylonitrile-Butadiene-Styrene (ABS) . . . . . . . . . . . . . . . . 6 1.2.4 Acetal (POM, Polyacetal) . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9 1.2.5 Polymethyl Metacrylate (Acrylic, PMMA) . . . . . . . . . . . . 12 1.2.6 High Temperature Nylon (HTN) . . . . . . . . . . . . . . . . . . . . 14 1.2.7 Ionomer Polymers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 1.2.8 Liquid Crystal Polymer (LCP) . . . . . . . . . . . . . . . . . . . . . . . 18 1.2.9 Polyamide (PA, Nylon) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 1.2.10 Polyetherimide (PEI) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23 1.2.11 Polyarylate (PAR) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25 1.2.12 Polyetherether Ketone (PEEK). . . . . . . . . . . . . . . . . . . . . . . 27 1.2.13 Polycarbonate (PC) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28 1.2.14 Modiﬁed Polyphenylene Oxide (PPO). . . . . . . . . . . . . . . . 31 1.2.15 Polybutylene Terephthalate (PBT) . . . . . . . . . . . . . . . . . . . 33 1.2.16 Polyethylene Terephthalate (PET) . . . . . . . . . . . . . . . . . . . . 34 1.2.17 Polyethylene (PE) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36 1.2.18 Polytetraﬂuoroethylene (PTFE) . . . . . . . . . . . . . . . . . . . . . 39 1.2.19 Polyphenylene Sulﬁde (PPS) . . . . . . . . . . . . . . . . . . . . . . . . 44 1.2.20 Polypropylene (PP) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46 1.2.21 Polystyrene (PS) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48 1.2.22 Polysulfone (PSU) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49 1.2.23 Polyvinyl Chloride (PVC) . . . . . . . . . . . . . . . . . . . . . . . . . . . 51 1.2.24 Styrene Acrylonitrile (SAN) . . . . . . . . . . . . . . . . . . . . . . . . . 53 1.3 Thermoplastic Elastomers (TPE) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55 1.3.1 Thermoplastic Elastomer Families . . . . . . . . . . . . . . . . . . . 56 1.3.2 Thermoplastic Polyurethane Elastomer (TPU) . . . . . . . . 57 1.3.3 Styrenic Block Copolymer (SBS). . . . . . . . . . . . . . . . . . . . . 60 1.3.4 Polyoleﬁn Thermoplastic Elastomer (TPO) . . . . . . . . . . . 62 1.3.5 Elastomeric Alloy Thermoplastic Vulcanized (TPV). . . . 65 1.3.6 Melt Processible Rubber (MPR) . . . . . . . . . . . . . . . . . . . . . 69 1.3.7 Copolyester Thermoplastic Elastomer . . . . . . . . . . . . . . . 71 1.3.8 Polyamide Thermoplastic Elastomer . . . . . . . . . . . . . . . . . 75 1.4 Liquid Injection Molding Silicone (LIM®) . . . . . . . . . . . . . . . . . . . . 77 1.4.1 LIM® Silicone Processing . . . . . . . . . . . . . . . . . . . . . . . . . . . 79 1.5 Thermoset Polymers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 82 1.5.1 Polyester Alkyd (PAK) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83 1.5.2 Diallyl Phthalate/Isophthalate (DAP, DAIP) . . . . . . . . . . . 85 1.5.3 Melamine Formaldehyde (MF) . . . . . . . . . . . . . . . . . . . . . . 87 1.5.4 Cellulosic Ester . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88 1.5.5 Cyanate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89 1.5.6 Epoxy (EP) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 92 1.5.7 Phenol Formaldehyde (Phenolic, PF). . . . . . . . . . . . . . . . . 94

X

Contents 1.5.8 1.5.9 1.5.10 1.5.11 1.5.12 1.5.13 1.5.14 1.5.15 1.5.16

Polybutadiene (PB) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97 Bismaleimide (BMI) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97 Unsaturated Polyester (UP) . . . . . . . . . . . . . . . . . . . . . . . . . 98 Polyimide (PI) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 101 Polyxylene . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103 Polyurethane (PUR). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 104 Silicone (SI) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 107 Urethane Hybrid . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109 Vinyl Ester (BPA) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111

2 Engineering Product Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1 Understanding the Properties of Materials . . . . . . . . . . . . . . . . . . . 2.1.1 Plastics Selection Guidelines . . . . . . . . . . . . . . . . . . . . . . . 2.2 Structural Design of Thermoplastic Components . . . . . . . . . . . . . 2.2.1 Stress-Strain Behavior . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.2 Tensile Testing of Viscoelastic Materials. . . . . . . . . . . . . . 2.3 Mechanical Properties of Materials . . . . . . . . . . . . . . . . . . . . . . . . . . 2.4 Tension and Compression Curves . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.5 Modulus of Elasticity (E) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.6 Stress and Strain Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.7 Thermoplastics Elastic Design Method . . . . . . . . . . . . . . . . . . . . . . 2.7.1 Working Stress. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.7.2 Compressive Stress . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.7.3 Flexural Stress . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.7.4 Coefﬁcient of Linear Thermal Expansion (α) . . . . . . . . 2.7.5 Poisson’s Ratio (υ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.7.6 Moisture Effects on Nylon . . . . . . . . . . . . . . . . . . . . . . . . . 2.7.7 Effects of Temperature on the Behavior of Thermoplastics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.8 Stress-Strain Recovery (Hysteresis) . . . . . . . . . . . . . . . . . . . . . . . . . . 2.8.1 Creep Behavior of Thermoplastics . . . . . . . . . . . . . . . . . . 2.8.2 Creep and Rupture Under Long-Term Load . . . . . . . . . . 2.8.3 Creep and Relaxation of Thermoplastics. . . . . . . . . . . . . 2.9 Flexural Beam Stress Distribution . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.10 Viscoelastic Modulus Design Method . . . . . . . . . . . . . . . . . . . . . . . . 2.11 Centroid, Section Area, and Moment of Inertia . . . . . . . . . . . . . . . 2.12 Radius of Gyration . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.13 Stress Analysis of Beams . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.13.1 Types of Loads. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.13.2 Normal Stresses in Beams . . . . . . . . . . . . . . . . . . . . . . . . . . 2.13.3 Shearing Force . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.14 Beam Deﬂection Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.14.1 Beam Deﬂection by Double Integration Method . . . . . 2.14.2 Beam Deﬂection Moment Area Method . . . . . . . . . . . . . 2.14.3 Applications of Moment Area and Double Integration Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.14.4 Beam Deﬂection Superposition Method . . . . . . . . . . . . . 2.15 Column Structural Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.15.1 Long Slender Column Critical Load (PCr) . . . . . . . . . . . . 2.15.2 Column Slenderness Ratio (L / r) . . . . . . . . . . . . . . . . . . . 2.15.3 Eccentrically Loaded Columns . . . . . . . . . . . . . . . . . . . . . 2.16 Flat Circular Plates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.16.1 Classiﬁcation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.16.2 Stress Analysis Methods . . . . . . . . . . . . . . . . . . . . . . . . . . .

115 115 117 120 121 122 126 129 129 130 131 132 133 134 135 136 136 137 138 138 139 139 145 147 150 158 158 158 159 164 168 169 178 179 183 188 188 188 188 194 195 195

XI

Contents 2.16.3 2.16.4 2.16.5 2.16.6

Flat Circular Plate Equations . . . . . . . . . . . . . . . . . . . . . . . Flat Circular Plate Stresses . . . . . . . . . . . . . . . . . . . . . . . . . Theory of Flexure Comparison . . . . . . . . . . . . . . . . . . . . . Circular Plates Simply Supported, Concentrated Center Load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.16.7 Flat Circular Plate under Concentrated Center Load . . 2.16.8 Flat Circular Plate with Fixed Edge . . . . . . . . . . . . . . . . . . 2.16.9 Flat Circular Plate Compensation Factor for Deﬂection. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.16.10 Flat Circular Plate Bending under Edge Boundaries . . . 2.17 Torsion Structural Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 Structural Designs for Thermoplastics . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1 Uniform and Symmetrical Wall Thickness . . . . . . . . . . . . . . . . . . . 3.1.1 Part Geometries Difﬁcult to Mold . . . . . . . . . . . . . . . . . . 3.1.2 Wall Draft Angle per Side . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2 Structural Rib Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1 Rib Strength Analysis Method . . . . . . . . . . . . . . . . . . . . . . 3.3 Internal Sharp Corners and Notches . . . . . . . . . . . . . . . . . . . . . . . . . 3.4 Injection Molded Thermoplastic Bosses. . . . . . . . . . . . . . . . . . . . . . 3.5 Injection Molded Thermoplastic Threads . . . . . . . . . . . . . . . . . . . . 3.6 Collapsible Core for Molding Internal Threads . . . . . . . . . . . . . . . 3.7 Preferred Standard Thread Forms for Thermoplastics . . . . . . . . . 3.7.1 Thermoplastic Threads Creep Effects. . . . . . . . . . . . . . . . 3.8 Injection Molded Products with Undercuts. . . . . . . . . . . . . . . . . . . 3.9 Injection Molded Integral Life Hinges . . . . . . . . . . . . . . . . . . . . . . . 3.9.1 Injection Molded Integral Life Hinge Design . . . . . . . . . 3.9.2 Mold Design Considerations for Hinges . . . . . . . . . . . . . 3.9.3 Proper Gate Design for Life Hinges . . . . . . . . . . . . . . . . . 3.10 Conventional Types of Pin Hinges. . . . . . . . . . . . . . . . . . . . . . . . . . . 3.11 Metal Inserts for Thermoplastic Encapsulation . . . . . . . . . . . . . . . 3.11.1 Machined Metal Threaded Insert Tolerances . . . . . . . . . 3.11.2 Thermoplastic Boss Wall Thickness for Metal Inserts. . 3.11.3 Press/Lock Slotted Metal Insert Installation After Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.11.4 Cold Forged Metal Inserts for Encapsulation . . . . . . . . . 3.11.5 Threaded Female Metal Inserts . . . . . . . . . . . . . . . . . . . . . 3.11.6 Metal Inserts Anchorage for Thermoplastic Encapsulation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.11.7 Metal Insert Encapsulating Process Problems . . . . . . . . 3.11.8 Special Metal Inserts Anchorage for Encapsulation . . . 3.11.9 Electrical Lead Inserts for Encapsulation. . . . . . . . . . . . . 3.11.10 Inserts Preparation for Molding Encapsulation . . . . . . . 4 Thermoplastic Gearing Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1 Classiﬁcation of Gears . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.1 Gears Parallel to the Shaft Axis . . . . . . . . . . . . . . . . . . . . . 4.1.2 Bevel Gears, Nonparallel and Intersecting Shafts . . . . . . 4.1.3 Hypoid Gears, Nonparallel and Nonintersecting Shafts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.4 Gears for Straight Linear Motion . . . . . . . . . . . . . . . . . . . 4.2 Standard Injection Molded Thermoplastic Gears . . . . . . . . . . . . . 4.2.1 Selection of Thermoplastic Resins for Gears . . . . . . . . . 4.2.2 Horsepower Equations for Gears . . . . . . . . . . . . . . . . . . .

196 197 198 198 199 199 200 200 207 211 211 212 213 213 215 222 222 224 224 225 227 227 232 233 235 236 237 239 240 240 242 243 244 246 249 250 253 255 257 258 258 259 261 262 263 264 266

XII

Contents 4.2.3 Spur Gear Terminology and Deﬁnitions . . . . . . . . . . . . . Properties Required for Injection Molded Thermoplastic Gears . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Thermoplastic Spur Gear Design Requirements . . . . . . . . . . . . . . 4.4.1 Gating Effects on Thermoplastic Gear Roundness Dimensions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.2 Multifunction Designs with Thermoplastic Gears. . . . . 4.4.3 Mounting Thermoplastic Gears on Metal Shafts . . . . . . 4.4.4 Standard Spur Gears, Equations, and Calculations . . . . 4.4.5 Spur Gear Pitch Backlash . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.6 Standard Spur Gear Tooth Size Selection . . . . . . . . . . . . 4.4.7 Standard Gear Total Composite Tolerances . . . . . . . . . . Tolerances and Mold Shrinkage of Thermoplastic Gears . . . . . . . Standard Helical Gears . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Standard Straight Bevel Gears . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Standard Worm Gears . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.8.1 Standard Worm Gear Analysis . . . . . . . . . . . . . . . . . . . . . . Plastic Gearing Technology Designs . . . . . . . . . . . . . . . . . . . . . . . . . 4.10.1 Spur and Helical Gears PGT-1 Tooth Design . . . . . . . . . 4.10.2 Spur and Helical Gears PGT-2 Tooth Design . . . . . . . . . 4.10.3 Spur and Helical Gears PGT-3 Tooth Design . . . . . . . . . 4.10.4 Spur and Helical Gears PGT-4 Tooth Design . . . . . . . . . 4.10.5 Plastic Gearing Technology Tooth Form Design Variables . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.10.6 Maximum Allowable Outside Diameter DO (Max.) . . . 4.10.7 Spur Gear Tooth Form Comparison. . . . . . . . . . . . . . . . . 4.10.8 Mating Spur Gears Tooth Form Comparison . . . . . . . . . 4.10.9 PGT Spur Mating Gears Strength Balance . . . . . . . . . . . 4.10.10 PGT Close Mesh Center Distance Between Spur Gears . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.10.11 Maximum Close Mesh Center Distance . . . . . . . . . . . . . PGT Helical Thermoplastic Gearing . . . . . . . . . . . . . . . . . . . . . . . . . 4.11.1 PGT-1 Helical Mating Gears Strength Balance . . . . . . . . 4.11.2 PGT-1 Helical Mating Gears Center Distance . . . . . . . . PGT Spur and Helical Gears Horsepower Rating . . . . . . . . . . . . . . 4.12.1 PGT Gear Horsepower Equation Basic Parameters . . . . PGT Spur and Helical Gear Speciﬁcations. . . . . . . . . . . . . . . . . . . .

308 309 314 319 322 323 324 328

5 Plastic Journal Bearing Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2 Materials Used for Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.1 Babbitt Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.2 Bronze Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.3 Sintered Porous Metal Journal Bearings . . . . . . . . . . . . . 5.2.4 Plugged Bronze Journal Bearings . . . . . . . . . . . . . . . . . . . 5.2.5 Carbon-Graphite Journal Bearings . . . . . . . . . . . . . . . . . . 5.2.6 Cast-iron Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.7 Wooden Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.8 Rubber Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.9 Self-Lubricated Thermoplastic Journal Bearings . . . . . . 5.3 Hydrodynamics of Lubrication . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4 Journal Bearings Design for Lubrication . . . . . . . . . . . . . . . . . . . . . 5.5 Journal Bearing Design Principles . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.5.1 Journal Bearing Nomenclature and Equations . . . . . . . .

335 335 335 336 336 336 336 337 337 337 337 338 339 342 345 345

4.3 4.4

4.5 4.6 4.7 4.8 4.10

4.11

4.12 4.13

268 272 273 275 277 279 279 281 282 283 287 289 290 292 293 294 295 297 298 299 300 302 303 304 305

XIII

Contents

5.6 5.7 5.8 5.9 5.10 5.11 5.12

5.13

5.14 5.15

5.16

5.17

5.5.2 Thermoplastic Journal Bearing Axial Wall Thickness . . 5.5.3 Mounting Thermoplastic Journal Bearings . . . . . . . . . . Split Bushing Thermoplastic Journal Bearings . . . . . . . . . . . . . . . . Self-Centering Thermoplastic Journal Bearings . . . . . . . . . . . . . . . Journal Bearing Load Carrying Contact Surface (C) . . . . . . . . . . . Load Reaction Across the Length of Thermoplastic Bearing . . . . Injection Molded Journal Bearings Process Defects . . . . . . . . . . . . Factors Affecting Journal Bearing Performance . . . . . . . . . . . . . . . Factors Affecting Journal Bearing Dimensions . . . . . . . . . . . . . . . . 5.12.1 Length-to-Inside Diameter Ratio of Journal Bearings . 5.12.2 Types of Service and Motion of Journal Bearings . . . . . 5.12.3 Thermoplastic Journal Bearing Annealing Effects . . . . . 5.12.4 Acetal Homopolymer Moisture Absorption Effects . . . 5.12.5 TFE and Nylon 6/6 Moisture Absorption Effects . . . . . . 5.12.6 Temperature Effects on Thermoplastic Journal Bearings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.12.7 Thermal Effects on Thermoplastic Journal Bearing Clearances . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Journal Bearing Pressure-Velocity (PV) Limits . . . . . . . . . . . . . . . . 5.13.1 Methods to Determine the PV Limits of Plastics . . . . . . 5.13.2 Journal Bearing Coefﬁcient of Friction . . . . . . . . . . . . . . 5.13.3 Journal Bearing Failures Due to Small Clearances. . . . . 5.13.4 Deﬁnition of Different Types of Wear . . . . . . . . . . . . . . . Mating Material Hardness and Surface Finishing. . . . . . . . . . . . . . Self-Lubricated Thermoplastic Journal Bearings . . . . . . . . . . . . . . 5.15.1 Vespel® Polyimide Bearings . . . . . . . . . . . . . . . . . . . . . . . . 5.15.2 Journal Bearing Pressure Equation . . . . . . . . . . . . . . . . . . 5.15.3 Vespel® Wear Factor Effects Caused by Temperature . . 5.15.4 Vespel® Wear Transition Temperature . . . . . . . . . . . . . . . 5.15.5 Frictional Behavior of Vespel® . . . . . . . . . . . . . . . . . . . . . . 5.15.6 Vespel® Journal Bearings Length to Inside Diameter Ratio . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.15.7 Vespel® Thrust Bearing Ratio Between Diameters . . . . 5.15.8 Vespel® Journal Bearing Initial Clearance (cI) . . . . . . . . 5.15.9 Vespel® Journal Bearing Inside Diameter (dB) . . . . . . . . Teﬂon® (TFE) Fabric Composite Bearings . . . . . . . . . . . . . . . . . . . 5.16.1 Bearing Physical Properties . . . . . . . . . . . . . . . . . . . . . . . . 5.16.2 Bearing PV Limit Rating . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.16.3 Journal Bearing Clearances (c) . . . . . . . . . . . . . . . . . . . . . Thermoplastic Kevlar® Reinforced Bearings . . . . . . . . . . . . . . . . . .

6 Thermoplastic Molded Spring Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Thermoplastic Molded Spring Design Considerations . . . . . . . . . 6.3 Thermoplastic Helical Compression Springs . . . . . . . . . . . . . . . . . 6.4 Thermoplastic Molded Cantilever Beam Springs . . . . . . . . . . . . . . 6.5 Cantilever Beam Spring Design Analysis . . . . . . . . . . . . . . . . . . . . . 6.5.1 Initial Modulus of Elasticity Cantilever Beam Analysis Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.2 Stress-Strain Curve Cantilever Beam Analysis Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.3 Empirical Data Cantilever Spring Analysis Method . . . 6.6 Thermoplastic Cantilever Spring Applications . . . . . . . . . . . . . . . . 6.7 Thermoplastic Belleville Spring Washers . . . . . . . . . . . . . . . . . . . . .

347 347 348 348 350 350 351 352 353 354 354 354 355 355 356 357 358 359 359 360 361 362 363 366 367 368 369 369 370 370 370 371 373 374 374 375 375 377 377 378 378 379 381 381 381 382 385 388

XIV

Contents 6.7.1 6.7.2 6.7.3

Acetal Homopolymer Belleville Spring Washer Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 389 Belleville Spring Washer Loading Rate. . . . . . . . . . . . . . . 392 Belleville Spring Washer Long-Term Loading Characteristics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 392

7 Thermoplastic Pressure Vessel Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.1 Thermoplastic Thin-Walled Pressure Vessels . . . . . . . . . . . . . . . . . 7.2 Thin-Walled Cylinder Basic Principles . . . . . . . . . . . . . . . . . . . . . . . 7.3 Thick-Walled Pressure Vessels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.1 Lame’s Equation for Thick-Walled Cylinders . . . . . . . . . 7.3.2 Maximum Stresses with Internal and External Pressures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.3 Maximum Stresses for Internal Pressure Only . . . . . . . . 7.4 Designing Cylinders for Cost Reduction . . . . . . . . . . . . . . . . . . . . . 7.5 Thermoplastic Pressure Vessels Design Guidelines . . . . . . . . . . . . 7.5.1 Preliminary Pressure Vessel Design. . . . . . . . . . . . . . . . . . 7.6 Testing Prototype Thermoplastic Pressure Vessels . . . . . . . . . . . . . 7.6.1 Redesign and Retesting the Pressure Vessels . . . . . . . . . . 7.7 Pressure Vessel Regulations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.7.1 ASME Pressure Vessel Code . . . . . . . . . . . . . . . . . . . . . . . .

393 393 394 396 396

8 Thermoplastic Assembly Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2 Cold Heading Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2.1 Cold Heading Procedure and Equipment . . . . . . . . . . . . 8.3 Electro Fusion Fitting System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.1 The SEF-System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.4 Hot Plate Welding Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.4.1 Hot Plate Welding Joint Design . . . . . . . . . . . . . . . . . . . . . 8.4.2 Flash or Weld Bead . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.5 Solvent and Adhesive Bonding Methods . . . . . . . . . . . . . . . . . . . . . 8.5.1 Solvents Used to Bond Thermoplastic Polymers . . . . . . 8.6 Adhesive Bonding Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.1 Adhesive Families . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.2 Adhesive Concerns . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.3 Adhesives Bonding Selection . . . . . . . . . . . . . . . . . . . . . . . 8.6.4 Ultra Violet Curable Adhesives . . . . . . . . . . . . . . . . . . . . . 8.6.5 Adhesive Surface Preparation. . . . . . . . . . . . . . . . . . . . . . . 8.6.6 Adhesive Application and Curing Methods . . . . . . . . . . 8.6.7 Joint Design for Adhesive Bonding . . . . . . . . . . . . . . . . . . 8.7 Metal Fasteners Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.7.1 Thermoplastic Bosses and Self-Tapping Screws . . . . . . . 8.7.2 Thread Forming and Thread Cutting Screws . . . . . . . . . 8.8 Press Fitting Method. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.8.1 Press Fitting Interference . . . . . . . . . . . . . . . . . . . . . . . . . . 8.8.2 Circular Press Fitting Assembly Method . . . . . . . . . . . . . 8.9 Snap Fitting Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.9.1 Circular Undercut Snap Fitting Joints . . . . . . . . . . . . . . . 8.9.2 Suggestions for Stripping Circular Undercut Snap Fitting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.9.3 Cantilevered Latch Snap Fitting Joint. . . . . . . . . . . . . . . . 8.9.4 Cantilever Snap Fit Latch Design Guidelines . . . . . . . . . 8.9.5 Cantilever Latch Snap Fit Mathematical Model . . . . . . .

405 405 405 406 408 409 410 412 413 413 414 416 416 419 420 421 424 425 425 427 429 430 437 439 441 444 445

398 398 400 400 400 402 402 402 403

446 447 449 450

XV

Contents 8.9.6

8.10

8.11

8.12

8.13

8.14

8.15 8.16 8.17

Cantilever Snap Latch Beam Permissible Deﬂection (δ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.9.7 Cantilever Latch Beam Assembly Force (W) . . . . . . . . . . 8.9.8 Design and Material Considerations . . . . . . . . . . . . . . . . 8.9.9 Uniform Cross Section Cantilever Beam . . . . . . . . . . . . . 8.9.10 Tapered Cross Section Cantilever Beam . . . . . . . . . . . . . Electromagnetic Welding Method . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.10.1 Electromagnetic Welding Process . . . . . . . . . . . . . . . . . . . 8.10.2 Electromagnetic Welding Coil Design . . . . . . . . . . . . . . . 8.10.3 Electromagnetic Welding Joint Design . . . . . . . . . . . . . . 8.10.4 Available Welding Gasket Shapes and Forms . . . . . . . . . Vibration Welding Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.11.1 High Frequency Vibration Welding . . . . . . . . . . . . . . . . . 8.11.2 Vibration Welding Modes . . . . . . . . . . . . . . . . . . . . . . . . . . 8.11.3 Comparing Vibration Welding to Other Assembly Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.11.4 Vibration Welding Equipment . . . . . . . . . . . . . . . . . . . . . . 8.11.5 Vibration Welding Joint Design. . . . . . . . . . . . . . . . . . . . . 8.11.6 Vibration Welding Aligning and Fixturing . . . . . . . . . . . 8.11.7 Vibration Welding Tolerances . . . . . . . . . . . . . . . . . . . . . . 8.11.8 Vibration Welding Equipment . . . . . . . . . . . . . . . . . . . . . . Spin Welding Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.12.1 Applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.12.2 Basic Spin Welding Equipment . . . . . . . . . . . . . . . . . . . . . 8.12.3 Spin Welding Variables . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.12.4 Types of Spin Welding Processes . . . . . . . . . . . . . . . . . . . . 8.12.5 Spin Welding Joint Designs . . . . . . . . . . . . . . . . . . . . . . . . 8.12.6 Spin Welding Process Suggestions. . . . . . . . . . . . . . . . . . . Ultrasonic Welding Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.13.1 Ultrasonic Welding Basic Principles . . . . . . . . . . . . . . . . . 8.13.2 Ultrasonic Welding Basic Components . . . . . . . . . . . . . . 8.13.3 Ultrasonic Welding Equipment . . . . . . . . . . . . . . . . . . . . . 8.13.4 Ultrasonic Welding Process Variables . . . . . . . . . . . . . . . . 8.13.5 Ultrasonic Welding Joint Designs . . . . . . . . . . . . . . . . . . . 8.13.6 Ultrasonic Welding Energy Director Butt Joint . . . . . . . 8.13.7 Ultrasonic Welding Method Design Limitations . . . . . . 8.13.8 Weldability of Thermoplastic Materials . . . . . . . . . . . . . . 8.13.9 Effects Caused by Thermoplastic Additives on Ultrasonic Welding. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ultrasonic Insertion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.14.1 Applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.14.2 Ultrasonic Insertion Conﬁgurations . . . . . . . . . . . . . . . . 8.14.3 Ultrasonic Insertion Product Design . . . . . . . . . . . . . . . . 8.14.4 Ultrasonic Insertion Equipment Requirements . . . . . . . 8.14.5 Ultrasonic Insertion Process Guidelines . . . . . . . . . . . . . Ultrasonic Stud Staking Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.15.1 Ultrasonic Stud Staking Joint Design . . . . . . . . . . . . . . . . Ultrasonic Stud Heading Method . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.16.1 Thermoplastic Stud Proﬁles for Ultrasonic Heading . . Ultrasonic Spot Welding Method. . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.17.1 Hand-Held Ultrasonic Spot Welder . . . . . . . . . . . . . . . . .

452 453 454 454 455 458 459 460 463 464 465 465 466 469 471 472 473 474 474 476 476 476 477 477 480 480 482 482 483 483 487 489 492 494 496 497 500 500 501 502 502 503 503 503 506 506 509 510

9 Thermoplastic Effects on Product Design . . . . . . . . . . . . . . . . . . . . . . . . . 511 9.1 Polymer Melt Behavior. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 511

XVI

Contents 9.1.1 Thermoplastics Glass Transition Temperature . . . . . . . . General Characteristics of Polymers . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.1 Critical Properties of Thermoplastics. . . . . . . . . . . . . . . . Polymer Reinforcements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.1 Types of Fiber Reinforcements . . . . . . . . . . . . . . . . . . . . . 9.3.2 Isotropic Warpage of Fiber Reinforced Resins . . . . . . . . 9.3.3 Fiber Glass Reinforcement Limitations . . . . . . . . . . . . . . 9.3.4 Injection Molding Process Effects on Fiber Glass Orientation . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5 Tensile Stress Effects Caused by Fiber Glass Orientation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.6 Flexural Modulus Effects Caused by Fiber Glass Orientation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Chemical and Environmental Resistance . . . . . . . . . . . . . . . . . . . . . 9.4.1 Effects of the Environment . . . . . . . . . . . . . . . . . . . . . . . . . Types of Degradations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.5.1 Oxidative Degradation . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.5.2 Radiation Degradation . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.5.3 Photo Oxidation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.5.4 Mechanical Degradation . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.5.5 Microbial Degradation . . . . . . . . . . . . . . . . . . . . . . . . . . . . Moisture Effects on Nylon Molded Parts . . . . . . . . . . . . . . . . . . . . . Aqueous Potassium Acetate for Moisture Conditioning Nylon . . Injection Molding Cycles . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Cavity Surface Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Cavity Temperature Control . . . . . . . . . . . . . . . . . . . . . . . . . . 9.10.1 Mold and Post-Mold Shrinkage. . . . . . . . . . . . . . . . . . . . . Process Condition Effects on Mold Shrinkage . . . . . . . . . . . . . . . . Post-Mold Shrinkage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Weld Lines. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

513 513 514 515 516 517 517

10 Injection Mold Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1 Classiﬁcation of Injection Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2 Effects of Product Design on the Injection Molding Process . . . . 10.2.1 Uniform Wall Thickness . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.2 Balance Geometrical Conﬁguration . . . . . . . . . . . . . . . . . 10.2.3 Smooth Internal Sharp Corners . . . . . . . . . . . . . . . . . . . . 10.2.4 Draft Walls . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.5 Feather Edges. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.6 Proportional Boss Geometries . . . . . . . . . . . . . . . . . . . . . . 10.2.7 Gate Type and Location . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.8 Molded Product Ejection Surface Area . . . . . . . . . . . . . . 10.2.9 Molded Product Tolerances . . . . . . . . . . . . . . . . . . . . . . . . 10.2.10 Surface Finish of Molded Product . . . . . . . . . . . . . . . . . . 10.3 Effects of Mold Design on the Injection Molding Process . . . . . . 10.3.1 Runner System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.2 Mold Cooling System. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.3 Ejector System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.4 Mold Venting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.5 Other Mold Devices . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.4 Design Considerations for Injection Molds . . . . . . . . . . . . . . . . . . . 10.4.1 Preliminary Mold Design . . . . . . . . . . . . . . . . . . . . . . . . . . 10.4.2 Detailed Mold Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.5 Types of Steels Required for Injection Molds . . . . . . . . . . . . . . . . .

545 545 546 547 547 547 547 547 548 548 548 548 549 549 549 549 550 550 550 550 551 552 553

9.2 9.3

9.4 9.5

9.6 9.7 9.8 9.9 9.10 9.11 9.12 9.13

517 518 519 520 521 522 522 522 522 522 523 523 527 528 529 530 531 533 538 541

XVII

Contents 10.5.1 Major Steel Families . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.6 Steels for Thermoplastic Injection Molds . . . . . . . . . . . . . . . . . . . . 10.6.1 General Steel Selection Procedures . . . . . . . . . . . . . . . . . . 10.6.2 Properties and Characteristics of Tool Steels . . . . . . . . . 10.6.3 Effects of Alloying Elements on Tool Steel Properties. . 10.6.4 Chemical Composition of Steels Used for Molds . . . . . 10.6.5 Effects of Alloying on Tool Steels . . . . . . . . . . . . . . . . . . . 10.6.6 Effects of Heat Treatment on Tool Steel Properties . . . . 10.6.7 Prehardened Tool Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.6.8 Carburizing Tool Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.6.9 Oil and Air Hardening Tool Steels. . . . . . . . . . . . . . . . . . . 10.6.10 Stainless Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.6.11 Steels Used in Thermoplastic Injection Mold Components . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.7 Mold Cavity Surface Finishing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.7.1 Mold Surface Finishing Process Procedures . . . . . . . . . . 10.8 Thermoplastic Injection Mold Bases . . . . . . . . . . . . . . . . . . . . . . . . . 10.8.1 Standard Mold Base Components. . . . . . . . . . . . . . . . . . . 10.8.2 Functions of the Mold Base Components . . . . . . . . . . . . 10.8.3 Types of Standard Mold Bases . . . . . . . . . . . . . . . . . . . . . . 10.9 Types of Thermoplastic Injection Molds . . . . . . . . . . . . . . . . . . . . . 10.9.1 Two-Plate Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.9.2 Round Mate® Interchangeable Insert Molds . . . . . . . . . 10.9.3 Master Unit Die Interchangeable Insert Molds . . . . . . . 10.9.4 Three-Plate Mold Cold Runner System . . . . . . . . . . . . . . 10.9.5 Vertical Insert Mold for Thermoplastic Encapsulations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.9.6 Hot Runner Molding Systems . . . . . . . . . . . . . . . . . . . . . . 10.9.7 Hot Runner Mold Temperature Control Systems . . . . . 10.9.8 Hot Runner Mold Gates (Drops) . . . . . . . . . . . . . . . . . . . 10.9.9 Types of Hot Runner Molding Systems . . . . . . . . . . . . . . 10.9.10 Thermoplastic Stack Injection Molds. . . . . . . . . . . . . . . . 10.9.11 Lost Core Thermoplastic Injection Molds. . . . . . . . . . . . 10.10 Number of Mold Cavities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.10.1 Cavity Number Limitations . . . . . . . . . . . . . . . . . . . . . . . . 10.10.2 Number of Mold Cavities Equation . . . . . . . . . . . . . . . . . 10.11 Mold Parting Line . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.11.1 Flat Mold Parting Line. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.11.2 Non-Flat Mold Parting Line . . . . . . . . . . . . . . . . . . . . . . . . 10.11.3 Balancing of Mold Parting Line Surfaces . . . . . . . . . . . . . 10.12 Mold Ejection Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.12.1 Ejector Plate Assembly. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.12.2 Ejector Plate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.12.3 Retaining Plate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.12.4 Ejector Sleeves . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.12.5 Types of Mold Ejection Systems . . . . . . . . . . . . . . . . . . . . 10.13 Injection Mold Cooling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.13.1 Mold Temperature Control . . . . . . . . . . . . . . . . . . . . . . . . 10.13.2 Factors Affecting Mold Cooling. . . . . . . . . . . . . . . . . . . . . 10.13.3 Effects Caused by Elevated Mold Temperature . . . . . . . 10.13.4 Effects Caused by Too Low a Mold Temperature . . . . . . 10.13.5 Mold Heat Transfer Methods . . . . . . . . . . . . . . . . . . . . . . . 10.13.6 Mold Cavity Insert Cooling . . . . . . . . . . . . . . . . . . . . . . . . 10.14 Injection Molding Machine Nozzle . . . . . . . . . . . . . . . . . . . . . . . . . .

553 557 558 559 559 559 560 562 564 566 567 568 569 571 573 578 578 579 582 583 584 585 585 586 587 588 589 590 593 601 602 606 606 606 607 607 608 610 610 611 611 611 611 612 615 616 617 617 618 618 631 639

XVIII

Contents 10.14.1 Mold Cold Runner System . . . . . . . . . . . . . . . . . . . . . . . . . 10.14.2 Determining the Injection Pressure Needed . . . . . . . . . . 10.14.3 Cold Runner Flow Tab. . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Cavity Gating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.15.1 Types of Mold Cavity Gates . . . . . . . . . . . . . . . . . . . . . . . . 10.15.2 Different Types of Hot Runner Gates . . . . . . . . . . . . . . . . Gate Molding Effects . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Venting Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.17.1 Product Design for Venting . . . . . . . . . . . . . . . . . . . . . . . . 10.17.2 Venting Characteristics of Thermoplastic Polymers . . . 10.17.3 Mold Deposit Problems . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.17.4 How to Avoid Venting Problems . . . . . . . . . . . . . . . . . . . . 10.17.5 Planning Mold Venting . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.17.6 Mold Venting Process Problems . . . . . . . . . . . . . . . . . . . . 10.17.7 Mold Venting Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.17.8 Mold Venting Using Sintered Porous Insert Plugs . . . . . 10.17.9 Logic Seal (Negative Coolant Pressure) Mold Venting . 10.17.10 Mold Cavity Vacuum Venting System . . . . . . . . . . . . . . . Mold Cavity Insert Contact Area Strength . . . . . . . . . . . . . . . . . . . . 10.18.1 Cavity Insert Sidewall Strength . . . . . . . . . . . . . . . . . . . . . 10.18.2 Methods to Calculate the Strength of Cavity Insert Sidewall . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Layout Case Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Support Pillars . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Tolerances for Thermoplastic Molded Parts . . . . . . . . . . . . . . . . . . 10.21.1 Factors Affecting Dimensional Control Tolerances . . . . General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins . . . . . . . . . . . . . . . . . . . . . 10.22.1 Mold Design Requirements . . . . . . . . . . . . . . . . . . . . . . . . 10.22.2 Mold Drawing Standards . . . . . . . . . . . . . . . . . . . . . . . . . . 10.22.3 Required Types of Tool Steels for Mold Construction . 10.22.4 Mold Construction Requirements . . . . . . . . . . . . . . . . . . Mold Tryout – Debug – Approvals – “MQ1” Requirements. . . . . 10.23.1 Mold Tryout or Evaluation . . . . . . . . . . . . . . . . . . . . . . . . . 10.23.2 Mold Debug Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.23.3 Approval of Molded Parts and Pre-Production Molding Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.23.4 Mold Cavity and Core Surface Temperatures . . . . . . . . . 10.23.5 “MQ1” Requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

639 653 654 655 656 663 664 666 667 669 669 670 671 672 674 690 691 693 698 699

11 Performance Testing of Thermoplastics . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1 Property Data Sheet for Thermoplastics . . . . . . . . . . . . . . . . . . . . . 11.2 Tensile Testing (ASTM D-638) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.1 Tensile Testing Equipment . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.2 Tensile Test Specimen . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.3 Specimen Conditioning . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.4 Tensile Strength Test Procedures . . . . . . . . . . . . . . . . . . . . 11.2.5 Tensile Modulus and Elongation . . . . . . . . . . . . . . . . . . . . 11.2.6 Molecular Orientation Effects . . . . . . . . . . . . . . . . . . . . . . 11.2.7 Crosshead Speed Effects . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.8 Temperature Effects . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.9 Moisture Absorption Effects. . . . . . . . . . . . . . . . . . . . . . . . 11.2.10 Stress-Strain Effects Caused by Creep . . . . . . . . . . . . . . . 11.3 Flexural Testing (ASTM D-790) . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

723 724 725 725 726 726 726 727 728 729 729 729 730 730

10.15

10.16 10.17

10.18

10.19 10.20 10.21 10.22

10.23

700 704 705 705 707 709 709 709 711 713 720 720 720 720 720 721

XIX

Contents

11.4

11.5

11.6

11.7

11.8

11.9

11.10 11.11 11.12

11.13

11.14

11.15 11.16 11.17 11.18

11.3.1 Apparatus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.3.2 Test Procedures and Equations . . . . . . . . . . . . . . . . . . . . . 11.3.3 Modulus of Elasticity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Compressive Strength Testing (ASTM D-695) . . . . . . . . . . . . . . . . 11.4.1 Compressive Testing Apparatus . . . . . . . . . . . . . . . . . . . . . 11.4.2 Test Specimens and Conditioning. . . . . . . . . . . . . . . . . . . 11.4.3 Test Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.4.4 Stress-Strain Tension and Compression Curves. . . . . . . Shear Strength Testing (ASTM D-732) . . . . . . . . . . . . . . . . . . . . . . . 11.5.1 Test Specimen and Apparatus . . . . . . . . . . . . . . . . . . . . . . 11.5.2 Test Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.5.3 Signiﬁcance and Limitations . . . . . . . . . . . . . . . . . . . . . . . Surface Hardness Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.6.1 Rockwell Hardness Testing (ASTM D-785-60T) . . . . . . 11.6.2 Barcol Hardness Testing (ASTM D-2583) . . . . . . . . . . . . 11.6.3 Factors Affecting the Test Results . . . . . . . . . . . . . . . . . . . Abrasion Resistance Testing (ASTM D-1044) . . . . . . . . . . . . . . . . . 11.7.1 Taber Abrasion Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.7.2 Theoretical Analysis of Wear . . . . . . . . . . . . . . . . . . . . . . . Coefﬁcient of Friction (ASTM D-1894) . . . . . . . . . . . . . . . . . . . . . . 11.8.1 Coefﬁcient of Friction of Thermoplastic Materials . . . . 11.8.3 Effects of Lubricants . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mold Shrinkage Test (ASTM D-955) . . . . . . . . . . . . . . . . . . . . . . . . 11.9.1 Purpose of the Mold Shrinkage Test . . . . . . . . . . . . . . . . . 11.9.2 Factors Affecting Mold Shrinkage . . . . . . . . . . . . . . . . . . . 11.9.3 Injection Molding Effects on Shrinkage. . . . . . . . . . . . . . 11.9.4 Requirements for Sampling . . . . . . . . . . . . . . . . . . . . . . . . 11.9.5 Test Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Speciﬁc Gravity Testing (ASTM D-792) . . . . . . . . . . . . . . . . . . . . . . 11.10.1 Test Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Density Gradient Testing (ASTM D-1505) . . . . . . . . . . . . . . . . . . . Water Absorption Testing (ASTM D-570) . . . . . . . . . . . . . . . . . . . . 11.12.1 Test Specimen . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.12.2 Test Procedure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Impact Resistance Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.13.1 Pendulum Impact Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.13.2 Charpy Impact Testing (ASTM D-256) . . . . . . . . . . . . . . 11.13.3 Chip Impact Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.13.4 Tensile Impact Testing (ASTM D-1822) . . . . . . . . . . . . . 11.13.5 Drop Weight Impact Testing (ASTM D-3029) . . . . . . . . 11.13.6 Falling Weight Impact Testing . . . . . . . . . . . . . . . . . . . . . . 11.13.7 Instrumented Impact Testing . . . . . . . . . . . . . . . . . . . . . . . Creep, Rupture, Relaxation, and Fatigue . . . . . . . . . . . . . . . . . . . . . 11.14.1 Tensile Creep Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.14.2 Flexural Creep Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.14.3 Procedure for Applying Creep Modulus . . . . . . . . . . . . . Melting Point Test (ASTM D-795) . . . . . . . . . . . . . . . . . . . . . . . . . . Vicat Softening Point (ASTM D-1525) . . . . . . . . . . . . . . . . . . . . . . . 11.16.1 Melting Point, Glass Transition Temperature . . . . . . . . . Brittleness Temperature (ASTM D-746) . . . . . . . . . . . . . . . . . . . . . 11.17.1 Test Apparatus and Procedures . . . . . . . . . . . . . . . . . . . . . UL – Temperature Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.18.1 Relative Thermal Indices . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.18.2 Long Term Thermal Aging Index . . . . . . . . . . . . . . . . . . .

731 732 733 733 734 734 734 735 735 735 736 736 736 737 739 740 740 741 741 742 743 744 744 744 745 745 745 746 748 749 750 750 751 751 751 753 755 755 755 756 757 758 761 761 762 764 767 767 768 768 768 770 770 772

XX

Contents 11.18.3 Creep Modulus/Creep Rupture Tests . . . . . . . . . . . . . . . . 11.19 Heat Deﬂection Temperature (ASTM D-648). . . . . . . . . . . . . . . . . 11.19.1 Apparatus and Test Specimens . . . . . . . . . . . . . . . . . . . . . 11.19.2 Test Procedure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.19.3 Test Variables and Limitations . . . . . . . . . . . . . . . . . . . . . . 11.20 Soldering Heat Resistance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.21 Coefﬁcient of Linear Thermal Expansion Testing . . . . . . . . . . . . . 11.21.1 Test Procedure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.22 Thermal Conductivity Testing (ASTM C-177) . . . . . . . . . . . . . . . . 11.23 Melt Flow Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.23.1 Moisture Content . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.24 Melt Index Testing (ASTM D-1238) . . . . . . . . . . . . . . . . . . . . . . . . . 11.24.1 Melt Flow Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.25 Capillary Rheometer Melt Viscosity Testing (ASTM D-1703) . . . 11.25.1 Melt Viscosity vs. Shear Rate Curves. . . . . . . . . . . . . . . . . 11.26 Electrical Properties Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.26.1 Underwriter’s Laboratories (UL) Yellow Cards . . . . . . . 11.26.2 How to Read and Interpret the “UL Yellow Card” . . . . . 11.26.3 “UL Insulation Systems Recognition” . . . . . . . . . . . . . . . 11.27 Electrical Insulation Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.28 Electrical Resistance Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.28.1 Volume Resistivity Testing (ASTM D-257) . . . . . . . . . . . 11.28.2 Surface Resistivity Testing (ASTM D-257) . . . . . . . . . . . 11.28.3 Dielectric Strength Testing (ASTM D-149) . . . . . . . . . . . 11.28.4 Dielectric Constant Testing (ASTM D-150) . . . . . . . . . . 11.28.5 Dissipation Factor Testing (ASTM D-150) . . . . . . . . . . . 11.28.6 Arc Resistance Testing (ASTM D-495) . . . . . . . . . . . . . . . 11.28.7 High Voltage Arc Tracking Rate (UL-746 A) . . . . . . . . . . 11.28.8 Comparative Track Index Testing (ASTM D-3638/ UL 746 A) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.29 Self and Flash Ignition Temperature Testing (ASTM D-1929) . . 11.29.1 Test Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.29.2 High Current Arc Ignition Testing (UL 746A) . . . . . . . . 11.29.3 Hot Wire Coil Ignition Testing (UL 746A/ASTM D-3874) . . . . . . . . . . . . . . . . . . . . . . . . . 11.29.4 Hot Mandrel Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.29.5 Glow Wire Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.30 Flammability Characteristics of Polymers . . . . . . . . . . . . . . . . . . . . 11.30.1 Inherently Flame Retardant Polymers . . . . . . . . . . . . . . . 11.30.2 Less Flame Retardant Polymers . . . . . . . . . . . . . . . . . . . . . 11.30.3 Flammable Polymers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.31 UL 94 Flammability Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.31.1 Horizontal Burning Testing, UL 94HB. . . . . . . . . . . . . . . 11.31.2 Vertical Burning Testing, UL 94-V0, UL 94-V1, UL 94-V2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.31.3 Vertical Burning Testing, UL 94-5V, UL 94-5VA, UL 94-5VB . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.31.4 Factors Affecting UL 94 Flammability Testing . . . . . . . . 11.32 Limited Oxygen Index Testing (ASTM D-2863) . . . . . . . . . . . . . . . 11.32.1 Test Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.32.2 Factors Affecting the Test Results . . . . . . . . . . . . . . . . . . . 11.33 Smoke Generation Testing. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.33.1 Smoke Density Testing (ASTM D-2843) . . . . . . . . . . . . . 11.34 Weathering Tests for Thermoplastic Materials . . . . . . . . . . . . . . . .

773 774 774 775 775 775 776 777 777 779 780 780 781 782 783 784 785 786 791 792 792 793 794 795 797 800 801 803 804 805 805 806 807 807 807 809 810 810 810 811 811 812 813 815 815 816 816 817 817 818

XXI

Contents 11.34.1 Weathering Creep Factors (Degradation) . . . . . . . . . . . . 11.34.2 Ultraviolet (UV) Radiation . . . . . . . . . . . . . . . . . . . . . . . . 11.34.3 Temperature. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.34.4 Moisture . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.34.5 Oxidation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.34.6 Micro-Organisms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Accelerated Weathering Testing (ASTM G 23) . . . . . . . . . . . . . . . . 11.35.1 Exposure to Fluorescent UV Lamp, Condensation (ASTM G 53) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.35.2 Accelerated Weather Testing, Weather-Ometer® . . . . . . 11.35.3 Exposure to Carbon Arc Light and Water Testing (ASTM D-1499) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.35.4 Exposure to Xenon Arc Light and Water Testing (ASTM D-2565) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.35.5 Outdoor Weathering Testing of Thermoplastics (ASTM D-1435) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fungi Resistance Testing of Thermoplastics (ASTM G 21) . . . . . Bacteria Resistance Testing of Thermoplastics (ASTM G 22) . . . Fungi and Bacteria Outdoor Exposure Resistance Limitations . .

827 828 829 829

12 Thermoplastic Product Cost Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.1 Injection Molding Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2 Molding Cycle Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.3 Material Handling (Regrinds) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.4 Capital Equipment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.5 Injection Molding Machine Size . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.6 Injection Molding Machine Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.7 Machine Installation and Safety Considerations . . . . . . . . . . . . . . 12.8 Auxiliary Equipment and Automation . . . . . . . . . . . . . . . . . . . . . . . 12.9 Mold Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.10 Molded Products Cost Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.10.1 Cost Analysis Basic Method . . . . . . . . . . . . . . . . . . . . . . . . 12.10.2 Cost Analysis Graph Method . . . . . . . . . . . . . . . . . . . . . . . 12.10.3 Advanced Cost Analysis Method . . . . . . . . . . . . . . . . . . . . 12.11 Secondary Molding Operations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.12 Additional Manufacturing Costs . . . . . . . . . . . . . . . . . . . . . . . . . . . .

831 832 832 833 833 833 836 837 837 838 841 841 842 843 848 848

Appendix . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Acronyms for Polymeric Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Common Acronyms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Process Acronyms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Reinforcement and Filler Acronyms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Nomenclature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . English and Metric Units Conversion Guide. . . . . . . . . . . . . . . . . . . . . . . .

849 849 850 851 851 852 853

11.35

11.36 11.37 11.38

818 819 819 820 820 820 821 821 822 823 825

Subject Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 855 About the Author . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 869

1

1

Polymeric Materials

1.1

Introduction to Plastic Materials

Plastic materials are the ultimate tribute to man’s creativity and inventiveness. Plastics are true man-made materials. Like any other materials, they have their origin in nature. The structure of plastic materials is based on basic chemical elements such as carbon, oxygen, hydrogen, nitrogen, chlorine, and sulfur. These elements are extracted from the air, water, gas, oil, coal, and even from living plants. It was man’s inspiration to take these elements and combine them through various chemical reactions in an almost unending series of combinations to produce the rich variety of polymers known today as plastics. It is possible to create different polymers from different combinations of elements and create almost any property desired for an end product. These new polymers have similar properties to existing conventional materials but they offer greater design freedom and cost incentives for manufacturing. There are some plastics with signiﬁcant property improvements over existing materials, while other polymers can only be described as unique materials with exceptional properties previously unknown to the industrial world. There are plastics that will melt at 200 °F, while other plastic materials can withstand up to 1,000 °F. The heat shields that protect astronauts travelling in space are plastic materials based on the technology known as “ablative plastics”. There are polymers used for shields that can stop a bullet. There are ﬂexible plastic ﬁlms that protect grocery products and there are rigid plastics rugged enough to serve as support beams in a building. Plastics are among the best electrical insulating materials known to mankind. However, we ﬁnd another type of special plastic material capable of conducting electricity. Plastic composite materials are used for golf club shafts, while other ﬂexible polymers are used as upholstery materials for furniture. There are impact resistant and transparent polymers used as windshields for airplanes, automobiles, and shower doors. There are also transparent packaging materials used to protect consumer items. The number of permutations possible in combining chemical elements to create plastics with different properties is almost endless. It is this diversity that has made plastics so applicable to such a broad range of end uses and products today. This polymer diversity makes it difﬁcult to grasp the idea of a single family of materials that can provide an inﬁnite range of properties, characteristics, and transformation processes.

1.1.1

Beginning of Plastics

Plastic materials have played an important role in the development of this modern civilization. These polymers have an extensive versatility of properties and process automation while offering several cost advantages. It is surprising to realize that a little more than a century ago there were no such plastic materials any where around the world. The plastics industry dates its beginning back to 1868, when John Wesley Hyatt mixed pyroxylin, made from cotton and nitric acid, with camphor to create an entirely different and new product called “Celluloid”. This material was the ﬁrst commercial plastic material. The

7

7 7

Figure 1-1 First photographic celluloid film

2

1 Polymeric Materials development of celluloid was in response to a competition sponsored by a manufacturer of billiard balls. It came about to overcome a shortage of ivory used to produce billiard balls. With the need for a new material and a production method for this application, celluloid was developed and the plastics industry was born.

Figure 1-2 First phenolic applications

Celluloid quickly moved into other markets, including new applications such as shirt collars, cuffs and shirt fronts, dolls, combs, buttons, and window curtains used on early automobiles. However, the most important celluloid application was the ﬁrst photographic ﬁlm used by Eastman to produce the ﬁrst motion picture ﬁlm in 1882. This material is still in use by the motion picture industry today, under its chemical name of cellulose nitrate. The plastics industry took its second major step 41 years later. Dr. Leo Hendrik Baekeland introduced the first phenol formaldehyde “Phenolic” in 1909. This was the ﬁrst plastic material to achieve world acceptance. What is more important, he also developed techniques for controlling and modifying the phenolformaldehyde reaction. This technology made it possible to produce useful items, such as marbleized clock bases or electric iron handles, under heat and pressure from phenolic. This process of liquefying the material to form various shapes under heat and pressure is the same process that is still in use by the industry to produce thermoset plastic materials. The third major step in plastic’s development took place in the 1920s with the introduction of cellulose acetate. This polymer was similar in structure to cellulose nitrate but safer in processing and use. Urea-formaldehyde can be processed like phenolic, but into light colored articles that were more attractive than the phenolic’s black and brown colors. Polyvinyl chloride (PVC) became the second largest selling plastic for such applications as ﬂooring, upholstery, wire, and cable insulation, tubing, hoses, and ﬁttings. Polyamide or nylon (Du Pont’s trade name) was ﬁrst developed as a ﬁber material. Nylon represents one of the most important new developments in the plastic industry. The research development work of W. T. Carothers in the late 1920s made possible the introduction of the nylon technology. The tempo of plastic’s development picked up considerably in the 1930s and the 1940s. Each decade newer, more exciting, more versatile plastics came into existence. In the 1930s, the acrylic resins were introduced for signs and transparent articles. The introduction of polystyrene made this polymer the third largest selling plastic for house wares, toys, and for applications in the packaging industry. Melamine resins were also introduced for use in dishware, paints, and wet strength paper. Melamine later became a critical element (as a binder) in the development of decorative laminate kitchen counter tops, table tops, and panels. During the World War II years of the 1940s, the demand for plastics accelerated as did research into new plastics that could aid in the defense effort. Polyethylene, today the most important type of plastic, was a war time development that grew out of the need for a superior insulating material for applications such as radar cables. The thermoset polyester resins were also introduced a decade later. Radical changes in the boat building industry were also a war time development introduced for military use. Acrylonitrile-butadiene-styrene (ABS) is best known today as the plastic material used for applications such as appliance housings,

3

1.1 Introduction to Plastic Materials refrigerator liners, safety helmets, tubing, telephone handsets, and luggage. The original ABS research work was a crash program during the war for the development of synthetic rubber. By the beginning of the 1950s, plastics were on their way to being accepted by designers and engineers as basic industrial materials. This decade also saw the introduction of polypropylene following the Nobel Award winning work of Karl Ziegler in Germany and Giulio Natta in Italy for “ordering” the molecular arrangement of plastics. Also highlighting this decade was the development of acetal and polycarbonate; two plastics that, along with nylon, came to form the nucleus of a subgroup in the plastic’s family known as the “Engineering Thermoplastics”. Their outstanding impact strength, thermal and dimensional stability enabled engineering plastic resins to compete directly with metal materials. The 1960s and 1970s also had their share of new plastics’ introductions. The most important contribution was the thermoplastic polyesters used in exterior automotive parts, under the hood applications, and electrical and electronic components. Polyester bottles internally coated with high nitrile barrier resins (outstanding resistance to gas permeation) developed the new drink bottle packaging applications. During this time span, another subgroup of the plastic’s family called “High Performance Plastics” found new markets; this group includes such materials as polyimide, polyamide-imide, aromatic polyester, polyphenylene sulﬁde, and polyether sulfone. These materials historically met their objectives in the demanding thermal needs of aerospace and aircraft applications; reinforcing the vision of the plastic’s industry that the future is, indeed, plastics.

1.1.2

Polymer Families

Plastic materials are the result of the combination of carbon elements reacting with oxygen, hydrogen, nitrogen, and other organic and inorganic elements. These polymers have the ability to change into a liquid (melt), and are capable of being formed into shapes by the application of heat and pressure. Plastics are a family of materials, not a single kind of material. Plastics have an extensive number of polymers and compounds with each kind of material having its own unique and special type of properties. Most plastics fall into one of the following groups: thermoplastics, thermoplastic elastomers, liquid injection molding elastomers, thermosets, and thermoset rubbers. Thermoplastic resins consist of a long chain of molecules, either linear or branched, having side chains or unattached groups to other polymer molecules. Usually, the commercial shapes of the thermoplastic materials are pellets, granules or powders. These materials can be repeatedly melted by heat under pressure so they can be formed, then cooled and harden into the ﬁnal desired shape. Chemical changes do not take place during the transformation process. Figure 1-3 shows a simple analogy for molding plastic resins, a wax block that can be liquiﬁed by heat, poured into a mold, then cooled to become a solid again. Thermoplastic elastomer (TPE) resins are rubbery materials with the characteristics of a thermoplastic and the performance properties of a thermoset rubber. TPEs are processed using the same thermoplastic equipment and methods, such as extrusion, injection molding, and blow molding. Liquid injection molding compounds are a family of unique products. Generally, these materials use two liquid formulations in a 1 : 1 ratio. These compounds

Figure 1-3 Thermoplastic materials analogy, wax candle

4

1 Polymeric Materials produce precision elastomeric molded parts efﬁciently. They use a liquid metering, mixing, and delivery system, a specially modiﬁed injection molding machine, and a high temperature precision mold. Thermoset materials have a reactive portion between the chain cross link and the long molecule’s network during polymerization. The linear polymer chains bond together to form a three-dimensional network. Therefore, once polymerized or hardened, the material cannot be softened by heating without degrading some linkages of the material. Thermoset materials in commercial form are supplied as resins, powders, and liquid monomer mixtures or as partially polymerized molding compounds. In this uncured condition, they conform to the ﬁnished shape with or without pressure and can be polymerized with chemicals or heat.

Figure 1-4 Thermoset plastics analogy, concrete

Figure 1-4 shows one of the analogies for the thermoset materials as the chemical transformation of concrete. When the cement powder blends with water and sand, the mixture becomes a thick paste compound. This mixture is then transferred to a cavity for curing and hardening to become a solid object (concrete). The chemical reaction transforms the product into concrete. The transformation processes of the concrete items are irreversible. Reprocessing concrete forms or returning to cement, sand, and water are not possible. The concrete becomes a new, different and strong material. Thermosetting materials are not reprocessable or recyclable. Thermosetting rubber materials are not covered in this book, because the technology, chemistry, part design, mold design, and processing are completely different and too intensive to review in this plastic product design handbook.

1.2

Thermoplastic Polymers

1.2.1

Classiﬁcation of Polymers by Performance

Thermoplastic resin classiﬁcations divide the polymers into four family groups based on their application performance. The ﬁrst is the commodity resins, which have a large consumption volume, extensive application end uses, low material cost, and limited property performances. The commodity resins include polystyrene (PS), polyethylene (PE), styrene acrylonitrile copolymer (SAN), cellulose nitrate (CN), polybutene (PB), bismaleimide (BMI), unsaturated polyester (UP), and polyvinyl chloride (PVC). The second group is classiﬁed as intermediate resins. These resins have mechanical, thermal, chemical, and electrical properties generally that are higher than the commodity resins. The basic matrix properties remain constant when modiﬁcations are made to change speciﬁc mechanical properties of the compound. This intermediate category of resins includes acrylics, thermoplastic oleﬁn (TPO), polyphenyleneoxide (PPO), thermoplastic vulcanizate (TPV), melt procesable rubber (MPR), high impact polystyrene (HIPS), ionomers, melamine formaldehyde (MF), polyxylene, polypropylene (PP), acrylonitrilebutadiene-styrene (ABS), styrene-acrylonitrile (SAN), polyphenylene ether (PPE), polyurethane (PUR), urethane hybrid, polyester alkyd (PAK), styrenic block copolymer thermoplastic elastomers (TPR), and ultra high molecular weight polyethylene (UHMWPE). The third group is classiﬁed as engineering resins. The level of mechanical properties that qualify as engineering grade is somewhat arbitrary; a tensile

1.2 Thermoplastic Polymers strength that is not lower than 7,000 psi with a minimum modulus of elasticity of 350,000 psi are reasonable criteria. Engineering resins are fundamentally unmodiﬁed resins, whose properties are improved by compounding. A compounded resin is deﬁned as a material containing a matrix (basic resin), additives, a reinforcing ingredient, such as ﬁber glass or minerals, heat and ultraviolet stabilizers, ﬂame retardants, and other additives. Several types of compounded resins are formulated to improve speciﬁc properties required for the application. A conventional ﬁber glass reinforced resin contains from 10 to 55% glass ﬁber. The glass ﬁbers are only 0.125 in long and a coating with a coupling agent is added to the glass ﬁber to obtain a bond with the matrix. The engineering resins are acetal, polyamide (nylon, PA), polycarbonate (PC), polybutylene terephthalate (PBT), polyethylene terephthalate (PET), glass ﬁber reinforced polypropylene, block copolyester TPE, polyamide TPE, liquid injection molding silicone, diallyl phthalate, epoxy (EP), and cyanate. The forth group is the high performance engineering resins. These resins in this category have the highest resistance retaining a high percentage of their useful mechanical properties at high temperatures, providing a longer service life of the product. They also maintain properties at higher electrical frequencies without sacrificing their chemical resistance properties when exposed to corrosive elements. These resins are also inherently ﬂame retardant, with UL-94 ﬂammability ratings of V0 and 5V. The high performance engineering resins include high temperature nylon (PA), liquid crystal polymers (LCP), polysulfone (PSU), fluoropolymers, polyetherimide (PEI), polyaryletherketone (PAEK), polyphenylene sulﬁde, silicone, and polyimide ((PMR). With few exceptions, high performance engineering resins do not have the higher Izod impact strength at room temperature that many engineering thermoplastic resins have. Competition among resin producers to capture markets has created an engineering resin supply of thousands of resin grades. The result is that there is usually more than one choice, and often several choices, available to meet the end product performance requirements. A total assessment of various grades includes product design and processing considerations. To select the best material for an application, compare the different properties and processing characteristics of several thermoplastic resins, which may meet the application requirements. However, product designers must make their resin selection based on other important characteristics of the resins, such as part design ﬂexibility, and process performance. Generally, most resins have different performance characteristics that could create molding problems and/or part failures. Part design and process characteristics can diverge when the resin properties between grades are matched. When additives, such as ﬂame retardants and stabilizers, are compounded into the resin, the characteristics of the matrix are modiﬁed, sometimes with a loss of some of the properties. In addition, when ﬁber glass or minerals are added to increase the mechanical strength properties, processibility becomes more difﬁcult, because the rheology or viscosity of the compounded product increases, decreasing the melt ﬂow rate (higher injection pressures and melt/mold temperature are required).

5

6

1 Polymeric Materials The economics involved in thermoplastic resin selection are complex, because the resin price is not usually the most important factor. When a thermoplastic resin is used for injection molding a close dimensional tolerance end product, the following engineering requirements are essential in reducing manufacturing costs: • Part geometry needs to be designed to produce molded components with maximum productivity. • A precision mold must be designed and constructed for fast running cycles and molding the maximum number of cavities automatically. • It is also important to select the best type, capacity, and running conditions of the injection molding machine. • Then, it is necessary to establish efﬁcient injection molding process setup conditions for the resin being used. It is also essential that the molding and maintenance organizations are well trained, using an up-to-date technical training program provided by a qualiﬁed instructor. Material costs become increasingly signiﬁcant with higher volume and less critical product requirements. Here, the resin cost represents a high percentage of the molded components ﬁnished cost. For example, common items such as business machine housings, plumbing (faucets, valves, tubing, and shower heads), kitchenware and appliance components (refrigerator liners, washing machine impellers, and vacuum cleaner housings) are price sensitive products. For these markets, the resin cost becomes a very competitive business aspect of manufacturing.

1.2.2 Amorphous molecules

Figure 1-5 Amorphous polymer molecular structure

Polymeric materials are an aggregate of long-chained molecular structures. There are two different states, one of which is comprised of high polymer compounds arranged in a crystalline structure (crystalline polymers) and the other in the form of ﬂexible molecular chains that are entangled (amorphous polymers). However, there is no crystalline plastic material exhibiting solely a crystalline structure, but rather they have a mixed structure in which crystalline sections and amorphous sections coexist. The ratio of crystalline sections is called the material’s crystallinity. Figures 1-5 and 1-6 show the difference between these families of polymers.

1.2.3

Amorphous molecules

Molecular Structure of Plastic Materials

Acrylonitrile-Butadiene-Styrene (ABS)

Crystalline structure

The ABS resins have a well balanced set of properties for molding close dimensional control articles with an outstanding surface ﬁnishing, good impact resistance, and metal plating characteristics. ABS resins belong to a very versatile family of thermoplastic polymers. They are produced by combining three monomers: acrylonitrile, butadiene, and styrene. The chemical structure of these monomers requires each monomer to be an important component of the ABS resins. Acrylonitrile contributes heat resistance, chemical resistance, and surface hardness to the system. The butadiene contributes toughness and impact resistance, while the styrene component contributes processibility, rigidity, and strength. Figure 1-6 Semi-crystalline polymer molecular structure

ABS plastics are two-phase systems. Styrene-acrylonitrile (SAN) forms the continuous matrix phase. The second phase is composed of dispersed polybutadiene

7

1.2 Thermoplastic Polymers

General Properties of ABS Speciﬁc gravity

1.05

Tensile modulus @ 73 °F (Mpsi)

0.3

Tensile strength @ yield (Kpsi)

5.0

Notch Izod impact @ 73 °F (ft-lb/in)

2.50–12.0

Thermal limits Service temp. (°F)

167–185

Shrinkage (%)

0.4–0.7

Tg (°F)

185–240

Vicat point (°F)

237

Process temp. (°F)

410–518

Mold temp. (°F)

122–176

Drying temp. (°F)

176–185

Drying time (h)

2.0–4.0

particles, which have a layer of SAN grafted onto their surface. The binding matrix layer of SAN makes this polymer’s two phases compatible. The property balance of ABS is controlled by the ratio of the monomers and by the molecular structure of the two phases. Stabilizers, lubricants, colorants, and other additives can be added to the system, and while this makes the production of ABS very complex, it allows great ﬂexibility in product property design. As a result of the unique morphology of ABS, hundreds of different products have been developed and are available commercially. ABS resins are grouped into two major divisions: injection molding and extrusion grades. The primary difference between these grades is their melt viscosity, which is signiﬁcantly lower for injection molding resins. Within each division of ABS polymers, there are the corresponding classes of grades. Standard ABS grades are grouped by impact strength into medium, high, and very high impact grades. The standard ABS versions are available in a low surface gloss, a high surface gloss, and an ultra-high surface gloss. Specialty ABS grades include high heat, plating, clear, ﬂame retardant, and structural foam grades. Standard grades of ABS generally meet the Underwriter’s Laboratories (UL) rating for slow burning (UL-94 HB). Flame retardant materials have UL-94 V0 at thicknesses as low as 0.062 in and UL-94 5V at thicknesses as low as 0.125 in. Clear ABS grades use methyl methacrylate providing light transmission of 72% and a haze level of 10%. Alloys of ABS-PVC are available in high and low gloss grades. Alloys of ABS-PC are available in injection molding and plating grades. ABS-SMA heat resistant alloys are available in injection molding, extrusion, and plating grade versions. Alloys of ABS-PA are also available in injection molding grades. ABS is an excellent choice for use in alloys and blends. When the plastics are combined, the positive features of each can be maintained, or even enhanced, while the negative features of each can be reduced. ABS-polycarbonate (ABSPC) and ABS-polyvinyl chloride (ABS-PVC) are well-established alloys. More recent innovations have resulted in ABS-styrene-maleic anhydride (ABS-SMA) and ABS-polyamide (ABS-PA) alloy products. ABS offers superior processibility and appearance as well as low cost, along with a good balance of engineering properties.

Figure 1-7 Portable power tool housing

8

1 Polymeric Materials Advantages • Good impact resistance (toughness) and rigidity properties • Low creep • Good dimensional stability • High strength properties • Metal coatings have excellent adherence to ABS • Transformed by conventional thermoplastic methods • A light-weight plastic material Disadvantages and Limitations Figure 1-8 Faucet chrome plated shell

• ABS is resistant to acids (except concentrated oxidizing acids), alkalis, salts, essential oils, and a wide range of food and pharmaceutical products. It is, however, attacked by many solvents, including ketone and ester. • Low dielectric strength • Only low elongation available • Low continuous service temperature • While the mechanical property of the ﬁnished part is not sensitive to moisture, its presence during processing can cause part appearance problems. Maximum suitable moisture levels of 0.2% for injection molding and 0.03% for extrusion can be reached using a dehumidifying air dryer. Typical Applications • Refrigerators: Doors and food liners for the interior surface of the refrigerator. Medium impact extrusion and molding grades, including clear ABS, are used in crisper pans, breaker strips, shelves, shelf supports, evaporator parts trays, and kick plates.

Figure 1-9 Brief case external covers

• Small Appliance Housings and Power Tool Applications: These include hair dryers, curling irons, blenders, electric can openers, coffee makers, food processors, electric fans, vacuum cleaners, electric drills, leaf blowers, and lawnmower decks. • Automotive Applications: Instrument panels, armrests, interior trim panels, seat belt retainers, glove compartment doors, and lift gates. ABS plating grades are used in wheel covers, grilles, headlight, mirror housings, and decorative trim. • Drain: Waste, vent pipes, pipe ﬁttings, and pool ﬁlter housings. • Telecommunications: Telephone housings, mobile phones, typewriter housings, and keyboard keys. • Business and Consumer Electronics: Videocassettes, televisions, audiovisual equipment, computer housings, ﬂoppy disks, printers, and copiers. • Household Items: Countertops, sinks, and tub surrounds, roof mounted air conditioning units. – Recreational: Motorcycle fairing, sailboats, airplanes, campers, hard-sided luggage, and picnic cooler liners.

Figure 1-10 Videocassette housing

• Other Applications: Briefcases, cosmetic cases, household packaging, toys, and photographic equipment.

9

1.2 Thermoplastic Polymers

1.2.4

Acetal (POM, Polyacetal)

General Properties of Acetal Homopolymers Speciﬁc gravity

1.42

Tensile modulus @ 73 °F (Mpsi)

400.00

Tensile strength @ yield (Kpsi)

10.0

Notch Izod impact @ 73 °F (ft-lb/in)

1.30

Thermal limits service temp. (°F)

230 (short) 195 (long)

Shrinkage (%)

1.9–2.3

Tm (°F)

350

Tg (°F)

–90

Process temp. (°F)

375–450

Mold temp. (°F)

140–200

Drying temp. (°F)

N/A or not required

Drying time (h)

N/A or not required

General Properties of Acetal Copolymers Speciﬁc gravity

1.42

Tensile modulus @ 73 °F (Mpsi)

360.00

Tensile strength @ yield (Kpsi)

8.5

Notch Izod impact @ 73 °F (ft-lb/in)

1.20

Thermal limits service temp. (°F)

200 (short) 175 (long)

Shrinkage (%)

2.0–2.5

Tm (°F)

330

Tg (°F)

–90

Process temp. (°F)

340–420

Mold temp. (°F)

125–185

Drying temp. (°F)

N/A or not required

Drying time (h)

N/A or not required

Acetal resins provide a well balanced set of properties including a hard selflubricated surface, excellent chemical resistance, strength, stiffness, and toughness over a broad temperature range. The acetal homopolymer was ﬁrst introduced in 1960 as a semi-crystalline form of polymerized formaldehyde forming a linear chain of molecules of oxymethylene. In the homopolymer process, the formaldehyde is separated from the water and purified to CH2O gas, which is then polymerized to the polyoxymethylene molecule. In this case, the molecule is stabilized by a reaction with acetic anhydride to give acetate end groups. The acetate capped homopolymer is less resistant to attack by base, but it has a higher melting point and mechanical advantages in strength, stiffness, toughness, hardness, creep, and fatigue than the copolymer acetal.

10

1 Polymeric Materials In the acetal copolymer process, the formaldehyde is ﬁrst converted to the cyclic structure of three formaldehyde molecules, trioxane. The trioxane is separated, puriﬁed and reacted with a comonomer (ethylene oxide) to prepare polyoxymethylene that has randomly distributed –CH2–CH2 groups in the chain. This resultant raw polymer is then treated with heat and base to degrade the ends of the molecules back to a –CH2–CH2 “block” point at each end. This leaves a molecule resistant to further degradation by basic environments. The end-capping of homopolymer and copolymer chains is necessary to prevent the irreversible depolymerization of the polymer backbone during melt processing by the thermal “unzipping” of the –H–O–CH2–O–CH2– end group to a formaldehyde monomer. The homopolymer grades exhibit the highest crystallinity, with good strength, stiffness, and impact resistance. The melting point is 350 °F. The high crystallinity also provides good chemical resistance, with little or no effect seen in the material after direct exposure to common hydrocarbons, aldehydes, ketones, alcohols, and fuels. The homopolymer is also resistant to aqueous solutions with a PH range from 4–10. The homopolymer is recommended for continuous service temperature in air and water up to 220 °F. The copolymer grades have random ethyleneoxy or n-butyleneoxy units scattered throughout the polymer backbone. These comonomer units slightly disrupt the crystallinity of the polymer in the solid state, leading to slightly lowered short term strength, stiffness, impact resistance, and a melting point of 330 °F. They have good solvent resistance, because the polymer chains are not capped with an ester group and aqueous solution PH resistance is extended over a range from 4–14. The copolymer is recommended for continuous service temperatures both in air and water up to 200 °F. Acetals are strong, stiff and tough over a broad temperature range. They have good surface lubricity and a low coefﬁcient of friction against metals, ceramics, and other plastics. They are creep and fatigue resistant because of low cold ﬂow characteristics. The balanced properties and good solvent resistance of acetal make the ideal candidate to replace materials such as metals, thermoset polymers, wood, and ceramics. Special technology has led to impact-modiﬁed homopolymer and copolymer grades, but increased impact resistance is offset by decreased strength and stiffness. Many grades also have the approval from the Food and Drug Administration for repeated food contact, National Sanitation Foundation and Canadian Standards Association for potable water applications and Underwriter’s Laboratories UL-94 HB ratings for ﬂammability. Special grades are also available with US Department of Agriculture approval for direct contact with meat and poultry products and Dairy and Food Industries Supply Association approval for contact with dairy products. The following grades of acetal resins are commercially available: • UV stabilized and weatherable grades,

• Low wear and low coefﬁcient of friction grades, • Toughened and impact modiﬁed grades, • Fiber glass reinforced grades,

• Filled mineral, glass bead, milled glass ﬁlled grades, and • High melt ﬂow grades.

11

1.2 Thermoplastic Polymers The outstanding characteristics of this polymer include stiffness, which permits the design of parts with large areas and thin cross sections; high tensile strength and creep resistance under a wide range of temperatures and humidity conditions; high fatigue resistance and resilience for applications requiring springiness and toughness. Acetal has achieved importance in applications because of a good balance of properties. Two types of acetals are available. One is a homopolymer resin with higher mechanical properties, higher end use temperatures, and higher melt ﬂow index, and the other is a copolymer resin with better processing characteristics and impact resistance. Advantages • High mechanical properties, tensile strength, rigidity, and toughness • Glossy molded surfaces • Low static and dynamic coefﬁcients of friction • Retains electrical and mechanical properties up to 250 °F • Low gas and vapor permeability • Approved for applications used in contact with food • Excellent chemical resistance to common hydrocarbons, aldehydes, ketones, alcohols, and fuels Disadvantages and Limitations • Poor resistance to acids and bases • High mold shrinkages • Subject to UV degradation, if special acetal grades are not used • Flammable (UL-94 HB) • Excessive process melt temperatures over 450 °F can result in signiﬁcant thermal degradation of the material, with the release of formaldehyde gases • Violent thermal degradation (explosion) if acetal melt is contaminated with PVC • Difﬁcult to bond when the acetal surface is not treated Typical Applications • Industrial: Conveyor links and slats, cams, bearings, wear stops, hose connectors, valve bodies, pumps (housings, pistons, valves and impellers) and gears. • Automotive: – Fuel handling systems: Filler caps, level sensors, ﬂoats, pumps and reservoirs – Trims: Seat belt buckle housings, window cranks, shift lever handles, knobs, buttons, mounting clips, visor mounting brackets, levers, exterior door pulls, mirror housings, brackets, bumper strip end plugs, and antenna bases – Instrument panel components: Cluster gears, bearings, housings, cable connectors, slide plates, and panel locks – Under the hood components: Fans, fan blades, dashpot housings, and tubing connectors

Figure 1-11 Automotive gasoline tank level sensor unit (Courtesy: Du Pont)

12

1 Polymeric Materials • Appliances: – Refrigerators: Shelving clips, brackets, bearings, and gears – Washers and dryers: Gears, bearings, wear strips, instrument housings, and hose connectors – Dishwashers: Rack rollers, spray nozzles, soap dispensers and ﬁlter supports

Figure 1-12 Food processor conveyor belt (Courtesy: Du Pont)

Figure 1-13 Textile injection molded components (Courtesy: Du Pont)

• Home Electronics: – Keyboards: Key caps, plungers, guides, and base plates – Telephones: Push buttons, gears, bearings, and springs – Modular components: Clips, peg boards, connectors, wear strips for drawers, latch springs, and clamps – Audio and video tape players and recorders: Tape hubs, guide rollers, cams, gears, bushings, and bearings • Plumbing: – Water meters: Housings, cams, gears, dials, and pressure plates – Faucets: Underbodies, cartridges, stems, packing nuts, and waterways – Water softeners: Pump housings, pistons, impellers, and valves – Filters: Bodies, plates, and screens – Pressure regulators: Bodies, stems, knobs, and pressure plates – Potable water distributors: Fittings, drain valves, stop valves, and metal pipe adapters • Consumer: – Personnel care: Mascara, perfume, and deodorant containers, combs, aerosol valves, soap dispensers, and cosmetic applicator handles – Small appliances: Motor gears, cams, bearings, pumps, glue applicators, housings, and springs – Toys: Shells, frames, gears, bearings, cams, springs, wheels, and connectors – Sporting goods: Ski bindings, gears, bearings, guides, wear plates, clamps, pump components, valves, and buckles • Hardware: Drapery and venetian blind guides, hangers, rollers, bearings, furniture casters, slide plates and locks, tool holders, bearings, gears, and housings

Figure 1-14 Gears appliance motor (Courtesy: Du Pont)

• Irrigation: Sprinkler nozzles, arms, gears, housings, waterways, pump housings, impellers, pistons, metering valve bodies, knobs, stems, and internal components • Agriculture: Shift levers and housings, hydraulic connectors, bearings, gears, and seed application disks

1.2.5

Polymethyl Metacrylate (Acrylic, PMMA)

Polymethyl metacrylate (acrylic) polymers have outstanding optical properties, weatherability, and a full range of transparent, translucent, and opaque colors. Acrylics are comprised of polymers and copolymers in which the major monomeric belong to two families of ester-acrylates and methacrylates. Hard, clear acrylic sheets are made from methyl methacrylate, molding and extrusion resins are made in a continuous solution from methyl methacrylate copolymerized with small percentages of other acrylates or methacrylates. Methyl methacrylate is produced by a dual step process in which acetone and hydrogen cyanide react to form acetone cyanohydrin. The compound is then

13

1.2 Thermoplastic Polymers

General Properties of Generic Unﬁlled PMMA Speciﬁc gravity

1.17

Tensile modulus @ 73 °F (Mpsi)

0.38

Tensile strength @ yield (Kpsi)

7.50

Notch Izod impact @ 73 °F (ft-lb/in)

0.30–0.50

Thermal limits service temp. (°F)

190 (short) 150 (long)

Shrinkage (%)

0.3–0.6

Tg (°F)

230

Vicat point (°F)

184

Melt ﬂow rate (g/10 min)

0.8–2.0

Process temp. (°F)

410–575

Mold temp. (°F)

140–190

Drying temp. (°F)

165

Drying time (h)

2–4

heated with methanol in the presence of concentrated sulfuric acid to yield the monomer. Acrylic monomers polymerize by free-radical processes initiated by peroxides in the polymerization process. A monomer initiator active at elevated temperature, this reaction is vigorous and liberates tremendous heat that must be dissipated. The formulations differ in molecular weight and in their principle properties (ﬂow rate, heat resistance, and toughness). Special formulations are available that provide matte surfaces or that absorb or transmit ultraviolet light, and a full range of transparent, translucent and opaque color resins. High impact acrylic grades for injection molding and extrusion are available. These compounds are composed of an acrylic hard phase and an acrylic modiﬁer as the soft phase. Acrylic polymers have outstanding optical properties and weatherability. Colorless acrylic is capable of transmitting white light up to 92%, with the remaining 8% being the reﬂection loss, and has haze values of only 1–2%. Acrylics have outstanding resistance to the effects of sunlight and long-term exposure to the elements. The low strain optic coefﬁcient of acrylics, coupled with their ability to be molded with very low stress, makes an ideal material for video disks. Sheet extruded from acrylic base impact modiﬁed grade has excellent thermoforming characteristics and can be stiffened by applying glass reinforced polyester to the inside surface with a spray gun to produce bathroom whirlpool tubs. High ﬂow grade has the best transparency, because it does not contain acrylonitrile, making it suitable for medical applications in which transparency is of prime importance. Acrylic plastics can be cleaned with solutions of inorganic acids, alkalis, and aliphatic hydrocarbons. However, chlorinated and aromatic hydrocarbons, esters, and ketones will attack the acrylic plastics. Advantages • Excellent optical clarity • Excellent surface hardness

14

1 Polymeric Materials • Excellent weatherability and resistance to sunlight • Rigid with good impact strength • Excellent dimensional stability and low mold shrinkage • Thermoforming increases bi-axial toughness Disadvantages and Limitations Figure 1-15 Acrylic – medical devices

• Poor solvent resistance; attacked especially by ketones, esters, chlorocarbons, aromatic hydrocarbons, and Freon® • Combustible, continuous service temperature limited to 160 °F • Flexible grades unavailable • Moisture produces dimensional variations of the molded articles Typical Applications • Automobile: Tail lights, parking light lenses, decorative emblems, medallions, and name plates

Figure 1-16 Acrylic – typical applications

• Household: Light ﬁxtures, picture frames, and decorative articles • Transparent Items: Available in a rainbow of sparkling colors, ideal for packaging, jewelry, and signs • Electronics: Used on print circuit board coating applications

1.2.6

High Temperature Nylon (HTN)

General Properties of HTN – 30% GR @ 50% R.H. Speciﬁc gravity

1.44

Tensile modulus @ 73 °F (Mpsi)

1,500

Tensile strength @ yield (Kpsi)

32.0

Notch Izod impact @ 73 °F (ft-lb/in)

1.80

Thermal limits service temp. (°F)

440 (short) 315 (long)

Shrinkage (%)

0.2–0.6

Tg (°F)

257

Tm (°F)

570

HDT (°F) @264 psi

510

Process temp. (°F)

580–620

Mold temp. (°F)

260–300

Drying temp. (°F)

175

Drying time (h)

2–16

There is a cost and performance gap between the engineering thermoplastic resins such as polycarbonate, nylon 6/6, acetal, and the ultrahigh performance polymers such as PEI, PEEK, LCP, and a group of relatively new resins with the convenient generic name of high temperature nylons. These HTN resins are chemically related (but not identical). The following types of HTN resins are commercially available worldwide:

1.2 Thermoplastic Polymers • Nylon (4,6) • Polyphthalamide PPA (6T/6I/66) • Nylon (6,T/D,T) • Nylon (6/6T) • Nylon (6T/6I) • Nylon (66/6T) HTN, an aliphatic aromatic polyamide line, offers many grades, variously employing glass ﬁbers, mineral ﬁllers, ﬂame retardants, and impact modiﬁers. HTN also provides an additional 50 °F or more of high temperature service as compared to standard nylons. Although molders running nylon 6 and 6/6 will not have to change molds or machines to run HTN, they will certainly come across some processing differences. Extremely close attention must be paid to drying. Drying conventional nylons requires a moisture content of no more than 0.2%. A sealed HTN bag comes predried, but with the recommendation to predry at 175 °F for 2–16 or more hours. As with all nylons, part dimensioning must take into account both shrinkage and ambient moisture absorption by the molded parts. The mold shrinkage rates of these semi-crystalline HTN polymers are similar to the conventional nylons. Shrinkage rates range from 0.18–0.22 in/in for the unreinforced grades to 0.02–0.06 in/in for the reinforced grades. Moisture absorption will reduce shrinkage. Hot runnerless molds can be used, but individual temperature control of drops is recommended. Molds run hot, hotter than for conventional nylons, which would require electric or oil heating. The HTNs are semi-crystalline resins. By increasing the melt temperature during processing, it raises the degree of crystallinity in the polymer, consequently giving a boost to post-mold dimensional stability and an improvement of the chemical resistance of the material. There is an increase in melt viscosity with the HTN, it does not run like water, as the standard easy ﬂow nylons do, but the advantage is less mold ﬂashing with lower temperature molds (180 °F). One of the characteristics of the HTN resins is the fast molding cycle. These materials set up (harden) very quickly in the mold cavity before the molded parts are ejected. This family of materials may carry the name nylon, but their high performance drastically exceeds the characteristics of the traditional nylon materials. The HTN (6,T/D,T) series of polyphthalamides has a melt temperature of 570 °F and provides excellent heat resistance for heat soldering owing to its partial aromatic nature. The FR grades are ﬂame retardant resins that have achieved UL-94 V0 rating for a 0.0031 in wall thickness. Advantages • Lower moisture absorption than nylon 6 and nylon 6/6 • Good dimensional stability • Excellent mechanical properties • Low creep characteristics • Excellent chemical resistance • High temperature performance, HTN uses part polyester and part nylon, which provides 50 °F higher melt temperature than nylon 6/6

15

16

1 Polymeric Materials • High glass transition temperature (257 °F) • Good injection molding processibility Disadvantages and Limitations • Its main weakness is low elongation. • Low crystallization (needs 300 °F mold temperature). Figure 1-17 Distribution transformer, various components (Courtesy: Du Pont)

• Problems with predrying the resin. Impact resistance and ﬁnished part appearances are adversely affected by excessive moisture during processing. Maximum suitable moisture levels of 0.2% for injection molding; needs a drying time of 2–16 h (depending on the type of compound); a dehumidifying dryer is recommended. Typical Applications • Transformer tri-clamp • Windows lift motor housing • Automotive safety: air bag sensor housings, ABS (brake sensors), speed and temperature sensors • Solenoid coil bobbins

Figure 1-18 Insulator electrical motor (Courtesy: Du Pont)

• Connector lamp sockets Automotive engine and transmission bearings, thrust washers, timing gears, valve stems and retainers, seal rings, piston skirts, and lifter wear pads

1.2.7

Ionomer Polymers

General Properties of Generic Ionomer Polymers Speciﬁc gravity

0.94

Tensile modulus @ 73 °F (Mpsi)

14–67

Tensile strength @ yield (Kpsi)

3.2–3.8

Notch Izod impact @ 73 °F (ft-lb/in)

15.0 – no break

Thermal limits service temp. (°F)

–110 +120

Melt ﬂow rate (g/10 min)

1.1–5.0

Tm (°F)

178–205

Vicat point (°F)

135–160

HDT (°F) @ 66 psi

160–185

Process temp. (°F)

400–500

Mold temp. (°F)

40–120

Drying temp. (°F)

160–180

Drying time (h)

2–4

Ionomers are thermoplastic resins that contain metal ions along with organic chain molecules. These ions are either sodium or zinc, serve as “reversible” cross linking networks, and result in high levels of resilience and impact resistance. Ionomers are based on ethylene copolymers and contain carboxylic acid groups.

17

1.2 Thermoplastic Polymers Ionomers are supplied either as unmodiﬁed resins or as composites. The unmodiﬁed ionomers are ﬂexible polymers that can be melt-processed at low temperatures. The ionomers can be reinforced with glass ﬁbers to increase the strength and stiffness or ﬁlled with high loading of minerals without sacriﬁcing useful toughness. Other additives are also used to increase the heat deﬂection temperature properties. The most outstanding characteristics of both the unmodiﬁed and High Performance (HP) ionomers are their impact toughness and resistance to failure at low temperatures. In the case of a HP, for instance, the notched Izod impact is approx. 8 ft-lb/in at –110 °F; there are no clean breaks in room temperature Izod impact tests. Unmodiﬁed ionomers also have demonstrated outstanding performance over many years in uses involving repeated, severe impacts, notably golf ball and bowling pin covers. The degree of cut resistance shown in golf ball applications is outstanding. The environmental resistance of ionomers is also of signiﬁcance in many end uses. Ionomers are insoluble at normal end use temperatures and therefore resist etching by solvents. These unmodiﬁed ionomers are used for see-through applications because of their transparency or low haze properties. The selection of ion types is based on the speciﬁc property required. Zinc ionomers absorb less water and have better impact strength than the sodium types, but are inferior in transparency and oil resistance. The properties combine very high toughness with comparatively low modulus. The tensile strength of the polymer exhibits a well deﬁned yield point, followed by substantial cold draw; the stress increases during the cold draw, and consequently, the energy to break is very high. The stress strain characteristic combines features of a crystalline polymer and a cured elastomeric product. The low heat deﬂection temperature of unmodiﬁed ionomers requires end use testing for high elevated temperature service applications. Like other ﬂexible polymers, the ionomer polymers creep when loaded over long time periods. Mechanical properties data of a typical HP ionomer show that outstanding impact resistance is combined with moderate stiffness and yield strength. The heat deﬂection temperature range of 160–185 °F is adequate for many high temperature end uses, but careful testing is needed before commercial adoption. Because of their glass content, moldings of HP ionomers lack the glossy surfaces of many unreinforced plastics. The signiﬁcance of this appearance factor depends entirely on the transparency of the product in use. The density of unmodiﬁed ionomers ranges from 0.94–0.96 g/cm3, while the HP ionomers have a density of approx. 1.05 g/cm3.

Figure 1-19 Golf ball (Courtesy: Du Pont)

Advantages • High impact resistance • Outstanding cut resistance • Flexible and transparent material • Excellent low temperature properties • Low melt processing temperatures • Excellent UV resistance • Excellent solvent resistance at room temperature

Figure 1-20 Bowling pin (Courtesy: Du Pont)

18

1 Polymeric Materials Disadvantages and Limitations • Low heat deﬂection temperatures (160–185 °F) for HP ionomers • Poor surface appearance of HP ionomers • Poor creep resistance under continuous load • Limited to low temperature applications • Poor impact resistance and ﬁnished part appearances are molding problems caused by moisture absorption during processing. The resin must be predried, using a dehumidifying hopper dryer. Figure 1-21 Ski boots (Courtesy: Du Pont)

Typical Applications • Packaging: Films • Sporting Goods: Golf balls, bowling pins, ski boot’s components, sport footwear, roller skates, skis, and bowling machine components • Automotive Components: Air dams, bumper guards, metal-ionomer (foamed) laminates for decorative fascia, large tonnage brackets for engine and transmission shipment between assembly plants • Industrial: Hand tools, machine components, steps for marine ladders, and buoy applications

1.2.8

Liquid Crystal Polymer (LCP)

General Properties of Generic LCP – 30% GR Polymers Speciﬁc gravity

1.62

Tensile modulus @ 73 °F (Mpsi)

2.25

Tensile strength @ yield (Kpsi)

23.0

Notch Izod impact @ 73 °F (ft-lb/in)

1.30

Thermal limits Service temp. (°F)

400–465

Shrinkage (%)

0.2–0.5

Tm (°F)

630

Tg (°F)

250–355

HDT (°F) @ 264 psi

600

Process temp. (°F)

660–680

Mold temp. (°F)

150–230

Drying temp. (°F)

250–300

Drying time (h)

1–4.0

LCPs are a family comprised of diverse polymers, which lack chemical homogeneity found in the high temperature nylons. Most commercial LCPs are copolyesters, copolyamides, or polyester-amides, although many other linkages are possible. LCP structures range from partially aliphatic to wholly aromatic polymers. The melting point characteristics, the end use temperature properties, the chemical and solvent resistance, ﬂammability, processibility, and cost are different for each product compound. LCPs are processed in the liquid crystalline state. All thermoplastic LCPs show one-dimensional order, which results from the semi-rigid, essentially linear

19

1.2 Thermoplastic Polymers architecture of the molecules. They are found in either the solid or nematic states. Because melt viscosity increases by a factor of two to ten when the polymer passes from the nematic to the isotropic state, processing is extremely difﬁcult as melt temperatures approach the clearing point. One other important point is that the melt point is a very low enthalpy transition, so there is little crystallization, allowing fast cycling and low mold shrinkage. The chemical structure common to all melt processed LCPs is that of pare hydroxy benzoic acid. There are three classes of LCPs: The ﬁrst is a general purpose class, which exhibits exceptional ease of processing, dimensional stability, molded part repeatability, chemical resistance, ﬂame resistance, strength, and stiffness. The second is a more temperature resistant variant, which is not quite as easy to process, but has good dimensional stability, good chemical resistance, excellent ﬂame resistance, and very good strength and stiffness. The third, a lower temperature performance class, is less expensive and less easily processed, but has good dimensional stability, with good strength and stiffness, less chemical resistance, and moderate ﬂame resistance. Advantages • Cycle repeatability, tight tolerances, high melt ﬂow for thin wall parts, warpfree parts, low mold shrinkage, and fast cycles • Molded article dimensional changes are minimal when the item is exposed to high temperatures (vapor phase or infrared soldering), without annealing • Low, controllable coefﬁcient of thermal expansion that can be matched to glass, ceramics, and metals • LCP has inherent resistance to burn and very low smoke generation • Excellent chemical resistance to all organics, acids, and bases • Exceptional strength, stiffness, and toughness Disadvantages and Limitations

Figure 1-22 Surface mounted connector housings (Courtesy: Du Pont)

• Weld lines are very weak, because inherent stiffness of the molecules prevents their diffusion across the melt front to strengthen the weld. • Unﬁlled or unmodiﬁed LCP resins are so easy to orient that the surface can be abraded. The resins must be handled with extreme care. • Differences in coefﬁcient of thermal expansion caused by different degrees and directions of molecular orientation can cause warpage. • The extremely low die swell results in a tendency to produce surface jet marks when small gates are used.

Figure 1-23 Electrical surface mounted bobbins (Courtesy: Du Pont)

• Impact resistance problems are caused by excessive amounts of moisture during processing. Maximum moisture level of 0.01% for molding is recommended, using a high temperature dehumidifying hopper dryer. Typical Applications • Dual oven handles • Electrical and electronic components • Vapor phase and infrared solderable connectors, sockets, relay and capacitor housings, active and passive molded printed wiring board components, thin walled coil forms (bobbins), brackets

Figure 1-24 Medical needle-free syringe device (Courtesy: Du Pont)

20

1 Polymeric Materials • Precise dimensions, typically required for miniaturization, where the pin spacing is very small • Semiconductor components, in which ionic contaminants are not tolerable

1.2.9 Figure 1-25 Automotive speed sensor housing (Courtesy: Du Pont)

Polyamide (PA, Nylon)

General Properties of Generic Nylon 6 – 33% GR Resins @ 50% RH Speciﬁc gravity

1.30

Tensile modulus @ 73 °F (Mpsi)

0.8

Tensile strength @ yield (Kpsi)

13.0

Notch Izod impact @ 73 °F (ft-lb/in)

3.5

Thermal limits service temp. (°F)

300 (short) 195 (long)

Shrinkage (%)

0.2–0.6

Water absorption (%) @ 24h and 73 °F

1.1

Tm (°F)

410

HDT (°F) @ 264 psi

400

Process temp. (°F)

440–550

Mold temp. (°F)

140–200

Drying temp. (°F)

175

Drying time (h)

2–20

General Properties of Generic Nylon 6/6 – 33% GR Resins @ 50% RH Speciﬁc gravity

1.38

Tensile modulus @ 73 °F (Mpsi)

0.9

Tensile strength @ yield (Kpsi)

18.0

Notch Izod impact @ 73 °F (ft-lb/in)

2.5

Thermal limits service temp. (°F)

390 (short) 265 (long)

Shrinkage (%)

0.2–0.6

Water absorption (%) @ 24h and 73 °F

0.7

Tm (°F)

491

HDT (°F) @ 66 psi @ 264 psi

500 480

Process temp. (°F)

530–580

Mold temp. (°F)

80–200

Drying temp. (°F)

175

Drying time (h)

1–16

This polymer was commercially introduced in the 1930s by Du Pont as a result of the signiﬁcant research work of W. H. Carothers. Polyamides are most commonly regarded as synonymous with nylon (Du Pont’s trade name), that is, synthetic polymers that contain an amide group –CONH– as a recurring part of the chain.

21

1.2 Thermoplastic Polymers

General Properties of Generic Nylon 6/12 – 33% GR Resins @ 50% RH Speciﬁc gravity

1.32

Tensile modulus @ 73 °F (Mpsi)

0.9

Tensile strength @ yield (Kpsi)

20

Notch Izod impact @ 73 °F (ft-lb/in)

2.5

Thermal limits service temp. (°F)

305 (short) 200 (long)

Shrinkage (%)

0.2–0.6

Water absorption (%) @ 24h and 73 °F

0.16

Tm (°F)

414

HDT (°F) @ 264 psi

410

Process temp. (°F)

450–550

Mold temp. (°F)

80–200

Drying temp. (°F)

175

Drying time (h)

1–4

Nylons are made from (a) diamines and dibasic acids, (b) w-amino acids, or caprolactam. Nylon gradess are identiﬁed by a/b, which correspond to the number of C-atoms in the monomer’s diamine (a) and the number of w-amino acids (b). The most common types of resins are nylon 6, nylon 6/6, and nylon 6/12. They are made from hexamethylene diamine and the 12-carbon acid, dodecanedioic acid, or HOOC(CH2)10COOH. The molecular weights of nylons range from 11,000 to 34,000. The melting points range from 410 to 491 °F. Nylon 6/6 and nylon 6 are the most important commercial products. Other nylons are 6/9, 6/10, 6/12, 11, and 12. The more C-atoms, that is, the lower the concentration of amide groups, the lower the melting point. Nylons are modiﬁed by use of monomer mixtures leading to copolymers. These are normally less crystalline, more ﬂexible, and more soluble than the homopolymers. Additives are used in nylons to improve thermal and photolytic stability, facilitate processing, increase ﬂammability resistance, enhance hydrolytic resistance, and increase lubricity. Modiﬁcation is an important asset. Fiber and mineral reinforcement are widely used. Blending with elastomeric modiﬁers has yielded nylons with improved toughness. Toughened nylon alloys or blends in the dry as-molded condition have a notched Izod impact strength of 2.5–4.5 ft-lb/in versus 1.0 ft-lb/in for unmodiﬁed nylon 6/6 or nylon 6. The supertough nylon 6/6 exhibits Izod impact values higher than 15 ft-lb/in (50% R.H.) and ductile behavior and high Izod strengths independent of test variables. Nylon is resistant to oils, greases, solvents, and bases. Nylon also has fatigue, repeated impact toughness, and abrasion resistance, plus a low coefﬁcient of friction. Nylon also has high tensile strength properties, creep resistance, and retains most of its mechanical and electrical properties over a wide temperature range. Its limitations are high moisture pickup, with resulting changes in dimensional and mechanical properties. UL temperature’s rating for continuous service ranges from 195–265 °F.

22

1 Polymeric Materials The following types of nylon compounds are available: Lubricated, nucleated, heat stabilized, ultraviolet stabilized, hydrolytically stabilized, ﬂame retarded, glass reinforced, Kevlar® ﬁber reinforced, mineral reinforced, toughened, melt ﬂow modiﬁed, electrically conductive, and several nylon alloys. Advantages • Excellent toughness and impact resistance • Excellent abrasion resistance • Low coefﬁcient of friction Figure 1-26 Office chair structure components (Courtesy: Du Pont)

• High tensile strength properties, creep resistance, and retention of mechanical and electrical properties over a wide temperature range. • Excellent resistance to oils, greases, solvents, and bases • Processed by all thermoplastic methods Disadvantages and Limitations • High moisture pick-up causes dimensional changes of the molded part, based on the type of polymer. • Requires UV stabilization.

Figure 1-27 Automotive engine valve cover (Courtesy: Du Pont)

• Electrical and mechanical properties are greatly inﬂuenced by polymer type, moisture content, and chemical composition. • Low impact resistance and poor finished part appearances are caused by excessive amounts of moisture during processing. Maximum suitable moisture level of 0.2% for injection molding and 0.03% for extrusion is recommended using a dehumidifying hopper dryer. Typical Applications

Figure 1-28 Automotive speed sensor (Courtesy: Du Pont)

• Transportation: This segment represents the largest single market for nylons. Applications for unreinforced materials include electrical connectors, wire jackets, emission canisters, and light duty gears for windshield wipers and speedometers. Toughened nylons are used as stone shields and trim clips. Glass reinforced nylons are used for engine fans, radiator headers, brake and power steering ﬂuid reservoirs, valve covers, sensors, and fuel injectors. Mineral reinforced resins are used for mirror housings and tire hub covers. A combination of glass and minerals serve in exterior parts, such as fender extensions. • Electrical and Electronics: Flame retardant nylons, including those complying with Underwriter’s Laboratories UL-94 V0 requirements, play a major role in the electrical market (plugs, connectors, bobbins, wiring devices, terminal blocks, antenna mounting devices). • Appliances: Nylons are used not only for electrical components, but also for mechanical parts, housings, and other applications in power tools, washers, and various small appliances. • Telecommunication Applications: Relays, ﬁttings, and connections.

Figure 1-29 Automotive transmission speed sensor (Courtesy: Du Pont)

• Industrial: Hammer handles, mowing machine parts, unlubricated gears, bearings, anti-friction parts, and a variety of applications requiring snap ﬁts or spring load assembling.

23

1.2 Thermoplastic Polymers • Food and Textile Processing Equipment: Pumps, valves, meters, agricultural and printing devices, business and vending machines. • Consumer Products: Toughened nylon applications such as ski boots, ice and roller skate supports, racket sports equipment, bicycle wheels, kitchen utensils, toys and photographic equipment. • Nylon Films: They are widely used for packaging meats and cheeses and for cook-in bags and pouches. Nylon ﬁlms are also used as an enclosure for the thermoset fabrication of small airplane wings. • Wire and Cable Jacketing: They are used mostly as a protective layer over primary insulation. • Nylon Tubing: They are used to convey brake ﬂuids, refrigerants, or as lining for ﬂexible cables.

Figure 1-30 Marine radiator end caps (Courtesy: Du Pont)

• Extrusion: Sheets, rods and machining stock shapes. • Nylon 6/12 Mono-ﬁlaments: They ﬁnd extensive applications in brush bristles, ﬁshing lines, ropes, and sewing threads. They are also used for cloth stiffening, rugs, women’s nylons, and ﬁlter screens. • Nylon 11 or 12 is used to powder coat metals by means of electrostatic-spray or ﬂuidized-bed for food handling and the pharmaceutical industries. • Encapsulated nylon for spline shafts, automotive engine timing gears.

1.2.10

Polyetherimide (PEI)

General Properties of Generic PEI – 30% GR Resins Speciﬁc gravity

1.50

Tensile modulus @ 73 °F (Mpsi)

1.30

Tensile strength @ yield (Kpsi)

24.50

Notch Izod impact @ 73 °F (ft-lb/in)

1.90

Thermal limits service temp. (°F)

392 (short) 356 (long)

Shrinkage (%)

0.2–0.7

Vicat point (°F)

426

Tg (°F)

420

HDT (°F) @ 66 psi @ 264 psi

412–415 408–420

Process temp. (°F)

640–800

Mold temp. (°F)

150–350

Drying temp. (°F)

250–300

Drying time (h)

4.0–6.0

PEIs are amorphous, high performance thermoplastic polymers that have been in use since 1982. Their chemical structure consists of repeating aromatic imide and ether units. PEIs are characterized by high strength and rigidity at room and elevated temperatures, long-term high heat resistance, highly stable dimensional and electrical properties, and broad chemical resistance. PEI resins can be meltprocessed using typical thermoplastic processing and thermoforming equipment for high volume production. Unmodiﬁed PEI is amber in color and transparent

Figure 1-31 Electrical solenoid coils encapsulated (Courtesy: Du Pont)

24

1 Polymeric Materials and exhibits inherent ﬂame resistance and low smoke generation without the use of halogenated or other types of ﬂame retarding additives. The amorphous structure of PEIs contributes to their excellent dimensional stability, low shrinkage, and highly isotropic mechanical properties. Their high glass transition temperature (Tg) of 420 °F and high performance strength and modulus characteristics at elevated temperatures are provided by the very rigid imide groups in their chemical structure. The high Tg allows PEI to be used intermittently at 392 °F and permits short-term excursions to even higher temperatures. Higher strength and stiffness at elevated temperatures are achieved with glass or carbon ﬁber reinforcement. PEI unreinforced resins are rated as having 338 to 356 °F continuous end use temperatures by Underwriter’s Laboratories, Inc. and are listed as UL-94 V0, down to 0.010 in thickness. PEI resins are available either in unmodiﬁed form or reinforced with 10, 20, 30, and 40% glass ﬁber. Also available are grades with carbon reinforcement for high strength and static dissipation, and a series of products with internal lubricants. Other grades and blends are offered for blow molding, structural foam, and extrusion processes. PEI has been formulated to meet speciﬁc market needs, such as electromagnetic interference shielding capability and Federal Aviation Administration (FAA) heat release requirements. PEI also can be spun into ﬁber for use in advanced composite systems. Advantages • Excellent chemical resistance to most hydrocarbons, non-aromatic alcohols, and fully halogenated solvents • Good chemical resistance to mineral acids • Good hydrolytic stability to boiling water and steam autoclaving for sterilization (low water absorption) • Good ultraviolet radiation resistance • Good resistance to gamma radiation • Excellent electrical properties, suitable for use as a dielectric material with highly sensitive electronic components Disadvantages and Limitations • Short-term reaction when used with mild bases. Partially halogenated solvents, such as methylene chloride can be good solvents for PEIs. • The resin must be dried to less than 0.05% moisture content. • Molding process temperatures of 640–800 °F are required. • High mold temperatures are required (oil or electrical heat elements). • Price of the resin is high. Typical Applications • Automotive: Temperature and fuel sensors, air handling devices, and metallized reﬂectors Figure 1-32 Computer dual 50 pins connector

• Electrical and Electronics: Connectors, printed circuit boards and integrated circuit chip carriers, burn-in sockets, ﬂexible circuitry, bobbins, and explosion proof boxes

25

1.2 Thermoplastic Polymers • Packaging: Steam resistant, thermal stability, microwave resistant and compliance to Food and Drug Administration requirements • Aircraft: PEIs are used as interior materials both in sheet and molded forms. The resin is available in a variety of formulations primarily to improve chemical resistance and impact properties to meet both FAA regulations for heat release and airframe manufacturer standards for smoke and toxicity. FAA compliant blow molding grades are also available. • Industrial Applications: Corrosion resistant ﬂuid and air handling components, mechanical couplers, and threaded fasteners. • Medical: Surgical staplers and other tool housings, handles, non-implant devices, and trays.

1.2.11

Figure 1-33 Spider motor armature insulator

Polyarylate (PAR)

General Properties of Generic PAR Polymers Speciﬁc gravity

1.22

Flexural modulus E (Mpsi)

0.3–0.35

σY (Kpsi)

10.0

Izod impact @ 73 °F (ft-lb/in)

4.0–4.75

Thermal limits service temp. (°F)

210

Shrinkage (%)

0.9

Vicat point (°F)

392

Tg (°F)

379

HDT (°F) @ 264 psi

300–345

Process temp. (°F)

675–735

Mold temp. (°F)

250–300

Drying temp. (°F)

250–285

Drying time (h)

3.0–8.0

Polyarylates (PAR) are a family of tough polyesters with heat, creep, and warpage resistance, ultraviolet stability, ﬂammability resistance (without additives), and good electrical properties. PARs are aromatic polyesters derived from aromatic dicarboxylic acids and diphenols. Although both amorphous and crystalline resins have been developed, the amorphous resins have been the preferred polymer for most of the applications. The mechanical properties of amorphous PARs are similar to polycarbonates (rigid, strong, and tough). Heat deﬂection temperatures of 0.25 in thick amorphous PARs tested at 264 psi range from 300–345 °F. Good retention of notched Izod impact strength, after exposures of up to 300 °F for 336 hours. In addition, PARs are able to sustain loading stresses of 1,100,000 psi at 210 °F with less than 1% total strain deformation. Amorphous PARs have also shown excellent low temperature resistance and are useful as an adhesive down to –240 °F. They can withstand thermal cycling at temperatures ranging between –240 to +212 °F. Liquid crystalline PARs have a very high heat deﬂection temperature of 671 °F.

26

1 Polymeric Materials Amorphous PAR materials offer outstanding weatherability properties. Because excellent ultraviolet resistance is inherent in these resins, there is no need for additional ultraviolet absorbers. Under the inﬂuence of ultraviolet radiation, the polymer surface rapidly rearranges to form an o-hydroxybenzophenone structure, which provides stability against ultraviolet radiation. Mechanical as well as optical properties are retained after weathering exposure. Amorphous PARs offer good performance in all areas, including low ﬂame spread, high oxygen index, low smoke density and relatively low amounts of toxic gases. PAR products of combustion will not contain halogen, nitrogen, sulfur, or phosphorous components commonly found in ﬂame retardant compounds. Amorphous PARs show unusual ﬂexural recovery or elastic rebound, the combination of this elastic recovery and thermal resistance offer opportunities for hot comb and high temperature lens applications. Unmodiﬁed, amorphous PAR materials show similar chemical resistance as other amorphous resins, such as PC. Alloying with other resins can improve the resistance to chemicals such as gasoline and ethanol. The crystalline versions of PAR show good resistance to most chemical environments. Amorphous polyarylates injection molding melt temperatures range from 675–750 °F, depending on the formulation of the alloy. The most important condition for successful PAR processing is the dryness of the resin, with no more than 0.02% moisture level being acceptable. Resins are unstable in the presence of moisture at processing temperatures. Unsatisfactory part properties result from processing wet resins. Suitable decorating procedures for PAR include screen printing, hot stamping, decal transferring, and painting. No surface pretreatment is needed. Advantages • Good heat and cold temperature resistance, impact strength, and dimensional stability • Excellent high gloss surface used for metallization applications • Excellent long-term weather resistance • Excellent gasoline and oil resistance • Excellent ﬂammability resistance, colorability, and transparency • Excellent electrical shielding ability • Excellent gamma sterilizability Disadvantages and Limitations • The resin must be dried to moisture levels below 0.02%. • Heavy wall thickness drastically reduces the impact strength property of the molded part. • Polymer melt degradation occurs at high melt temperatures and on extended melt residence time in the processing equipment. • Slow melt ﬂow characteristics. • Solvent resistance is only fair. Figure 1-34 Gas knob components

• Properties become unstable when submerged in hot water

27

1.2 Thermoplastic Polymers Typical Applications • Automotive: Headlight housings and brake light reﬂectors, exterior trim parts, and alloy PAR body panels • Safety Equipment: Fire helmets, helmet face shields, mining light covers, and trafﬁc signal lenses • Electrical and Electronic: Electrical connectors, relay housings, coil bobbins, switches, and fuse covers

Figure 1-35 Gasoline product components

• Exterior Construction and Lighting: Glazing, skylights, and transparent panels

1.2.12

Polyetherether Ketone (PEEK)

General Properties of Generic PEEK Polymers Speciﬁc gravity

1.3–1.6

Flexural modulus E (Mpsi)

0.6–3.2

σY (Kpsi)

13.0–39.0

Izod impact @ 73 °F (ft-lb/in)

1.2–1.6

Thermal limits service temp. (°F)

482

Shrinkage (%)

1.0–1.8

Tm (°F)

645

Tg (°F)

289

Thermal degradation

805

Process temp. (°F)

660–790

Mold temp. (°F)

250–430

Drying temp. (°F)

300–350

Drying time (h)

3.0

Polyetherether ketone (PEEK) is a subgroup of ketone polymer materials that is similar to polyarylether ketone (PEAK) and polyetherketone ketone (PEKK). Because of their semi-crystalline nature, PEEK resins demonstrate an excellent balance of physical properties, including strength at elevated temperature, chemical resistance, and hydrolytic and thermal stability. They offer the highest level of thermal resistance together with thermoplastic processing capability. Polyetherether ketone resins are produced as unreinforced resins, glass ﬁber reinforced resins, and carbon ﬁber reinforced resins. PEEK materials retain their mechanical properties at very high end use temperatures. The retention of ﬂexural modulus and tensile strength properties is excellent, especially for the ﬁber reinforced resins. Another characteristic of PEEK-based resins is their outstanding thermal properties, the UL-94 rating for PEEK has been observed to be V0, without the addition of any ﬂame retardant additives. PEEK resins exhibit heat deﬂection temperatures that make them perform well under both low and high stress conditions; the UL continuous use temperature is higher than 480 °F. The electrical property values are good for PEEK resins.

Figure 1-36 Air quick disconnect couplings

28

1 Polymeric Materials Chemical resistance at elevated temperatures in various aggressive environments shows that reinforced PAEK is a chemically resistant material. The retention of tensile properties for unreinforced PAEK resins is greater than 75%. Glass reinforced PAEK resins exhibit a much higher retention rate of tensile properties but show some loss of tensile strength properties. This problem may be caused by the effects of the ﬁber glass coupling agent’s bonding strength with the matrix. Unreinforced PAEK resins chemically react to acids and bases. Figure 1-37 Various molded products

Reciprocating screw injection molding machines adapted to the high melt temperature and rheology of PAEK and PEEK are used to process these resins. Care must be exercised when working at these temperatures, because polymer melt degradation may occur beyond 805 °F. Its processibility is due to its semicrystalline characteristics after reaching its melting point of 645 °F. Advantages • Excellent ﬂexural modulus and tensile strength retention • Outstanding thermal properties • Low ﬂammability and smoke generation

Figure 1-38 Electronic wafer baskets

• UL-94 V0 rating, without ﬂame retardant additives • UL continuous use temperature is 480 °F • Good electrical properties Disadvantages and Limitations • High resin cost • Limited type of commercial resins available • Unreinforced resins may react to acids and bases • Injection molding machines must be adapted to the high melt temperature and rheology of the polymer

Figure 1-39 Aircraft bracket support

• Polymer degradation occurs beyond the melt temperature of 805 °F Typical Applications • Aircraft and Aerospace: Engine components, cabin interior material, air ducts, and exterior parts • Electrical and Electronics: Wire and cable • Chemical Processing: Pump components and oil seals • Industrial Equipment: Journal bearing surfaces

1.2.13

Polycarbonate (PC)

Polycarbonate was introduced in 1958. PC is an amorphous engineering thermoplastic material with exceptionally high impact strength, transparency, high temperature resistance, and dimensional stability. Melt ﬂow rate is one of the most important properties of polycarbonate. PC is produced by reacting Bisphenol A and carbonyl chloride in an interfacial process. This reaction is carried out under basic conditions in the presence of an aqueous and an organic phase. Molecular weight is controlled by a phenolic

29

1.2 Thermoplastic Polymers

General Properties of generic PC – 30% GR Polymers Speciﬁc gravity

1.40

Tensile modulus @ 73 °F (Mpsi)

1.25

Tensile strength @ yield (Kpsi)

19.00

Notch Izod impact @ 73 °F (ft-lb/in)

1.7–3.0

Thermal limits service temp. (°F)

220 265

Shrinkage (%)

0.15–0.6

Tm (°F)

267–495

Tg (°F)

293–300

Vicat point (°F)

305–310

Process temp. (°F)

430–620

Mold temp. (°F)

175–250

Drying temp. (°F)

250–260

Drying time (h)

2.0–4.0

Figure 1-40 Automotive headlamp assembly

chain stopper. Trifunctional monomers are added for increased melt strength for extrusion and blow molding applications. Polycarbonate does not have a true melting point as other crystalline polymers. It does, however, have a high glass transition temperature of 300 °F. Polycarbonate is characterized by an exceptionally high notched Izod impact strength of 12–17 ft-lb/in on a 0.125 in wall thickness. The impact strength, measured by a falling dart or a Dynatup® (dynamic impact testing equipment), is retained to temperatures as low as minus 60 °F. The heat deﬂection temperature of polycarbonate, at 264 psi, is 260–270 °F, this HDT can be increased to close to 320 °F when the matrix is compounded with a high heat polyphthalate carbonate. PC has low and predictable mold shrinkage characteristics. It also has good creep resistance and low moisture absorption properties. The molded surface of PC has a high gloss ﬁnishing. This material can be compounded to produce resins for sterilizability, ﬂame retardance, and stain resistance.

Figure 1-41 PC/ABS alloy – volvo bumper

Polycarbonate has high corona resistance and insulation resistance properties, as well as a dielectric constant that is independent of temperature. Polycarbonate must be dried before processing at a drying temperature of 250–260 °F for 3–4 hours, using a dehumidifying hopper dryer. This drying is essential to reduce the moisture content to 0.02% to prevent hydrolysis during molding.

Figure 1-42 Drinking mugs and pitchers

Polycarbonate requires a high melt temperature, from 430–620 °F. The mold temperature should range between 175 and 250 °F to control surface ﬁnish. The components should be molded stress-free, using a high injection molding pressure, low injection speeds to ﬁll the cavities, and long curing time. A great number of grades have been developed for special purposes by compounding ﬁber glass and mineral reinforcements with various additives. These ingredients enhance the thermal stability, ultraviolet stability, tensile strength, stiffness, and ﬂame retardants. In addition, the color, shielding ability, Food and Drug Administration compliance, gamma sterilizability, low temperature impact, solvent resistance, wear resistance, and foameability are obtained in these resins.

Figure 1-43 Binocular housing

30

1 Polymeric Materials Polycarbonate is a versatile compounding material and is used as a basic matrix component in the production of several alloy resins. For example, polycarbonate and polyester (polyethylene terephthalate and polybutylene terephthalate), polycarbonate and ABS (acrylonitrile-butadiene-styrene), and polycarbonate and SAM (styrene-maleic anhydride) are used in alloys and blends. Advantages • High impact strength • Low ﬂammability Figure 1-44 Air-conditioner appliance grille

• Electrical shielding ability • Gamma sterilizability • Wear resistance • High heat deﬂection temperatures • Good dimensional stability • Good electrical properties • Processable by all thermoplastic methods Disadvantages and Limitations • Soluble in selected chlorinated hydrocarbons • Exhibits crazing in acetone and is attacked by bases

Figure 1-45 Coil bobbins

• Its surface is relatively soft and therefore can be scratched • Ultraviolet resistance of non-ultraviolet-stabilized polycarbonate is poor. The molded part tends to yellow after long-term exposure to ultraviolet light • Heavy wall thickness drastically reduces the impact strength property of the molded part • Snap-off undercuts for assembly are not recommended • Poor stress cracking resistance • Subject to melt degradation at high processing temperatures and extended residence time, especially for the ﬂame retardant resins • Solvent resistance is only fair Typical Applications

Figure 1-46 Motorcycle helmet shell

• Electronic and Business Equipment: Business machine housings, computer parts and peripherals, connectors, terminal blocks, and telecommunication components • Appliances: Food processors, electrical kitchen components, power tool housings, refrigerator drawers, and vacuum cleaner components • Transportation: Tail and head lights, signal light lenses and housings, runway markers, blow molded spoilers, instrument panels, and seat backs • Safety and Sports: Sports helmets, recreational vehicle hoods, windshields, head lights, boat propellers, and sun glass lenses • Food Service: Microwave cookware, serving trays, mugs, pitchers, water bottles, baby bottles, and institutional storage containers

31

1.2 Thermoplastic Polymers • Medical: Tubing connectors, dialysis components and devices, blood oxygenators, ﬁlter housings, lenses, gamma sterilization appliances, and surgical staplers • Sheet Products: Institutional and mass transit glazing signs, aircraft interior panels, and greenhouse windows • Industrial: Mailboxes, material handling containers, and highway delineations

1.2.14

Modiﬁed Polyphenylene Oxide (PPO)

General Properties of Generic PPO – 30% GR Polymers Speciﬁc gravity

1.25

Tensile modulus @ 73 °F (Mpsi)

1.10

Tensile strength @ yield (Kpsi)

14.50

Notch Izod impact @ 73 °F (ft-lb/in)

1.7–3.0

Thermal limits service temp. (°F)

200–240

Shrinkage (%)

0.20–0.60

HDT (°F) @ 66 psi @ 264 psi

280–320 275–317

Tg (°F)

302

Vicat point (°F)

230–298

Process temp. (°F)

520–600

Mold temp. (°F)

160–220

Drying temp. (°F)

200–250

Drying time (h)

2.0–4.0

Polyphenylene oxide materials are rigid, amorphous, tough, and dimensionally stable at a wide range of high temperatures. The chemical composition of the homopolymer is a poly(2,6-dimethyl-1,4-phenylene ether) or a poly[oxy-(2-6dimethyl-1,4-phenylene)]. Modiﬁed PPO resins are the result of compounding PPO with polystyrene (PS) plus various additives. Unmodiﬁed PPO resins are high temperature polymers with glass transition temperatures (Tg) of approximately 302 °F. They are very difﬁcult to process, even at melt temperatures of 520–600 °F. The high softening point and melt viscosity of PPO are in contrast to the low softening point and melt viscosity of PS. The unique thermodynamic compatibility of PPO and PS, particularly high impact PS, or HIPS, permits these two polymers to be blended in all proportions. The blends exhibit lower viscosity and improve the processibility compared to pure PPO, as well as a single Tg (between HIPS at 212 °F and PPO at 302 °F, at all ratios). This permits tailoring of the viscosity and toughness properties of a blend to speciﬁc requirements. With the addition of rubber impact-modiﬁed HIPS provides the PPO blend with an Izod impact value higher than the blends of PPO and crystalline PS. Most blends contain 20–80% PPO, with 40–50% being the most common mixture. The HIPS impact modiﬁer and various additives represent the compounding ingredients of the resin. To improve the chemical resistance and thermal property of the compound, PPO has been alloyed with polyamide (nylon 6/6). To increase ﬂexural modulus and reduce mold shrinkage, glass and mineral ﬁlled modiﬁed PPO resins are available.

Figure 1-47 Laundry dryer housing

32

1 Polymeric Materials The PPO modiﬁed blend formulations have the following characteristics: high tensile strength, high heat resistance, and low moisture absorption properties. Heat deﬂection temperatures range from 180–310 °F. Small creep rates are detected at room temperature. There is a low coefﬁcient of thermal expansion. Notched Izod impact values range from 1.3–8 ft-lb/in. The dimensional stability of modiﬁed PPO at elevated temperatures and stress can be maintained in a number of environments, such as strong alkalies and bases, detergents, and hot water. Filled and ﬂame retardant (FR) grades that have Underwriter’s Laboratories (UL) recognition are available. The high melt strength of modiﬁed PPO makes possible injection molding, structural foam molding, extrusion, blow molding, and thermoforming processes. Modiﬁed PPO blends make excellent insulators, because they have low loss and dissipation factors over broad temperature and frequency ranges. These materials also have inherently low dielectric constants and high dielectric strengths. High humidity conditions have little effect on these properties. Secondary operations can be performed on parts made from modiﬁed PPO resins. However, to obtain good adhesion, application of a urethane, alkyd, or acrylic base primer is recommended before painting. For electroplating, a preactivator step is necessary before catalyst deposition on the surface. Sputtering, vacuum metallizing, and hot stamping can be used for decorative purposes. Assembly techniques include the use of solvents, ultrasonic welding, vibrational welding, hot melt adhesives, thread cutting, and thread forming screws and metal inserts encapsulation. Advantages Figure 1-48 Electrical connectors

• High mechanical strength • High heat and moisture resistance • Small creep rates at room temperature • Low coefﬁcient of thermal expansion • Good notched Izod impact • Excellent electrical properties • Good dimensional stability Disadvantages and Limitations

Figure 1-49 Office cabinet housing

• Most hydrocarbon and aromatic base substances such as esters, oils, grease, or alcohols cause stress cracks or soften PPO blends • Ultraviolet (UV) applications are not recommended • They are very difﬁcult to process, even at high melt temperatures Typical Applications • Blow Molded Products: These resins are used for ofﬁce furniture, automotive steering column covers, and appliance doors and ducts • Flame Retardant Products: They include business machine housings, decks, and enclosures

Figure 1-50 Automotive body panel

• Automotive: They include instrument panels, wheel covers, fuse blocks, trim, and windshield wiper blades

33

1.2 Thermoplastic Polymers • Electrical Applications: They include ﬁberoptic connectors, ceiling boxes, control housings, and load centers • Industrial: Pumps, impellers, shower heads, chemical process equipment, and ﬁlter bodies • Metal Plated: Modiﬁed PPO performs well in electromagnetic interference and radio frequency interference (EMI/RFI) shielded enclosures

1.2.15

Polybutylene Terephthalate (PBT)

General Properties of Generic PBT – 30% GR Polymers Speciﬁc gravity

1.53

Tensile modulus @ 73 °F (Mpsi)

1.35

Tensile strength @ yield (Kpsi)

17.50

Notch Izod impact @ 73 °F (ft-lb/in)

0.90

Thermal limits service temp. (°F)

200–250

Shrinkage (%)

0.30–2.30

HDT (°F) @ 66 psi @ 264 psi

421–510 385–437

Tg (°F)

113–140

Tm (°F)

437

Process temp. (°F)

470–530

Mold temp. (°F)

110–200

Drying temp. (°F)

250–300

Drying time (h)

2.5–5.5

Polybutylene terephthalate is a high performance, semi-crystalline resin, one of the toughest and most versatile of all engineering thermoplastics. Strong and lightweight, this polyester is characterized by low moisture absorption, excellent electrical properties, broad chemical resistance, lubricity, durability, mechanical strength, and heat resistance. These properties are stable over a broad range of temperature and humidity conditions. The resin is commonly supplied with ﬁber glass and/or mineral reinforcements. Polybutylene terephthalate is produced by the transesteriﬁcation of dimethyl terephthalate with butanediol. This reaction takes place by a catalyzed melt poly-condensation, resulting in a repetition of the molecular unit. PBT can be alloyed with 10–30% nylon to facilitate glass reinforcement. Moisture absorption can be reduced, processibility and mechanical properties boosted simultaneously, when PBT is alloyed with 15–25% low density polyethylene (LDPE). The addition of graft containing butadiene polymers, as well as urethanes or copolyester elastomers, reduces notch sensitivity. An enhanced gloss surface also can be achieved when PBT is blended with PET.

Figure 1-51 Automotive motor fan supports (Courtesy: Du Pont)

Advantages • Excellent resistance to water up to 122 °F temperature. It also resists most aqueous salt solutions, weak acids and bases, most organic solvents (aliphatic hydrocarbons, glycol, ethers, high molecular weight esters, and ketones), gasoline, and cleaning solutions at room temperature. This material is not attacked by most oils and greases at temperatures up to 140 °F.

Figure 1-52 Electric iron handle (Courtesy: Du Pont)

34

1 Polymeric Materials • Excellent electrical properties, including high dielectric strength and insulation resistance. It has superior arc resistance, consistent dielectric constant, and a low dissipation factor over a range of temperatures and humidity. The dielectric constant, dissipation factor is nearly independent of frequency and temperature. Underwriter’s Laboratories (UL) continuous use temperature indices for the electrical and mechanical strengths range from 248–265 °F. Disadvantages and Limitations

Figure 1-53 Motorcycle side panels (Courtesy: Du Pont)

• Sensitive to alkaline oxidizing acids, hot water, strong bases, aromatic solutions, and ketones above ambient temperatures. Solvents include triﬂuoroacetic acid, phenol with chlorinated aliphatic hydrocarbons, and hexaﬂuoroisoprophenol. • Unmodiﬁed PBT is notch sensitive. High levels of glass ﬁller make the material brittle. • Part warpage caused by different cooling rates resulting from part geometry or tool design lead to differential shrinkage. Typical Applications • Building and Construction: Housewares, lawn and garden • Automotive: Grilles, body panels, fenders, wheel covers, components for door handles, mirrors, and windows. Under the hood distributor caps, rotors, ignition components, head lamp systems, windshield wiper assemblies, water pumps, and brake systems

Figure 1-54 Automotive distributor cap (Courtesy: Du Pont)

• Electrical and Electronics: Switches, relays, motor housings, tube sockets, photoelectric cell receptacles, ballast housings, fuse cases, key caps for computer keyboards, chip carriers, and connectors • Telecommunications: Network connector devices, junction boxes, and ﬁberoptic buffer tubing • Material Handling: Conveyor chain links, gear wheels, bearings, monorail hangers, dunnage, collapsible part racks, foamed pallet containers, tote bins • Industrial Applications: Tubing, ﬁttings, conduits, roof anchors, plumbing, pump impellers, sprinkler bodies and levers, pipe caps, pressure vessel housings, cable protection, and valve components • Consumer Uses: Food processor blades, vacuum cleaner parts, fans, gears, furniture, iron handles, toaster side panels, hair dryer housings, coffee makers

1.2.16

Polyethylene Terephthalate (PET)

Polyethylene terephthalate (PET) is a versatile polymer widely used in synthetic ﬁbers, industrial and packaging ﬁlms, injection molding, blow molding, and thermoforming. High strength products are possible because of the ability of the material to be oriented and crystallized.

Figure 1-55 Electric motor end frame (Courtesy: Du Pont)

Unoriented PET melt crystallizes or hardens slowly during its production. The maximum crystallization rate is obtained when the melt temperature reaches 482–490 °F. In its amorphous state, PET tends to be brittle at room temperature; then it softens above its glass transition temperature of approx. 158 °F. Molten PET is subject to hydrolytic degradation and so must be rigorously dried before melt processing. A maximum moisture content of 0.02% is recommended.

35

1.2 Thermoplastic Polymers

General Properties of Generic PET – 30% GR Polymers Speciﬁc gravity

1.67

Tensile modulus @ 73 °F (Mpsi)

1.50

Tensile strength @ yield (Kpsi)

22.00

Notch Izod impact @ 73 °F (ft-lb/in)

1.60

Thermal limits service temp. (°F)

392

Shrinkage (%)

0.20–0.90

HDT (°F) @ 66 psi @ 264 psi

392–490 435

Tg (°F)

158

Tm (°F)

482–490

Process temp. (°F)

510–565

Mold temp. (°F)

150–250

Drying temp. (°F)

250–275

Drying time (h)

2.0–4.0

Figure 1-56 Encapsulated stator motor (Courtesy: Du Pont)

PET is prepared by a reaction of either puriﬁed terephthalic acid (PTA) or dimethyl terephthalate (DMT) with ethylene glycol (EG). The high viscosity melt is converted into amorphous clear pellets by rapid quenching and cutting in water. Thermoplastic polyesters are long chain, unbranched molecules that are produced by the condensation reaction between a dibasic organic acid or ester and a glycol. PET has the stiffest polymer chain possible for a thermoplastic polyester with an outstanding combination of strength, stiffness, and melt ﬂow properties. These special properties of PET are the basis for its broad application in areas such as electrical and electronics components, ﬁbers, ﬁlms, and drink bottles. However, the chain stiffness also retards the rate of crystallization from the melt that makes PET, by itself, unsatisfactory as a resin for injection molding. The PET injection molding grades include a proprietary crystallization system that provides the rapid crystallization needed for good moldability at normal molding temperatures.

Figure 1-57 Wheelchair structure (Courtesy: Du Pont)

Reinforced grades continue to be introduced as improvements in crystallization rates, impact, warpage and other physical properties are made. Most of the PET compounded resins are glass ﬁber/mineral reinforced. Glass ﬁber and mica are used for low warpage applications. Reinforced grade resins are available with ﬁber glass contents ranging from 15–55%. Several ﬂame retardant resins meeting UL-94 V0 and V5; high impact and foamed grades are also produced. Specialty copolymers based on polyester and polycarbonate blends of PET with PBT have also been introduced. Advantages • Excellent resistance to water up to 122 °F temperature; also resists most aqueous salt solutions, weak acids and bases, gasoline, and cleaning solutions. PET is not attacked by most oils and greases • Excellent electrical properties, including high dielectric strength and insulation resistance, consistent dielectric constant, low dissipation factor. Underwriter’s Laboratories (UL) continuous use temperature index for electrical, mechanical strengths ranges from 302–392 °F

Figure 1-58 Beverage bottles (Courtesy: Du Pont)

36

1 Polymeric Materials Disadvantages and Limitations • Sensitive to alkaline oxidizing acids, hot water, strong bases above 122 °F temperature • PET requires drying of the resin from 2–4 hours at 275 °F to reach a 0.02% of moisture before processing • Part warpage caused by different cooling rates, which are a result of part geometry or tool design, leads to differential shrinkage Typical Applications Figure 1-59 Electrical coil bobbin (Courtesy: Du Pont)

• Automotive: Body panels, spoilers, door handles, distributor caps, rotors, ignition components, head lamps, and windshield wiper • Electrical and Electronic: Switches, relays, motor housings, tube sockets, ballast housings, fuse cases, chip carriers, and connectors • Consumer Uses: Vacuum cleaner parts, fans, gears, furniture, iron skirts, hair dryer housings, and coffee makers

1.2.17 Figure 1-60 Transformer spool and housing (Courtesy: Du Pont)

Polyethylene (PE)

General Properties of Generic Unﬁlled HDPE and LDPE Polymers

Speciﬁc gravity

HDPE

LDPE

0.94

0.91

Tensile modulus @ 73 °F (Mpsi)

0.2

0.05

Tensile strength @ yield (Kpsi)

3.75

2.25

Notch Izod impact @ 73 °F (ft-lb/in)

No break

No break

Thermal limits service temp. (°F)

158–176

140–167

Shrinkage (%)

1.1–1.4

1.1–1.4

Vicat point (°F)

255

–

Tg (°F)

–150

–150

Tm (°F)

257–275

212–230

Process temp. (°F)

400–535

360–530

Mold temp. (°F)

50–140

50–140

Drying temp. (°F)

150

150

Drying time (h)

3.0

3.0

The experimental development and testing of polyethylene polymers in the 1930s for use as a high frequency insulation for radar cables during World War II gave impetus to its commercial production. Polyethylene is available in a range of ﬂexibilities, depending on the production process. High density polyethylene (HDPE) is the most rigid of the three basic types of PE resins (HDPE, low density polyethylene (LDPE), and linear low density polyethylene (LLDPE)). HDPE can be formed by a wide variety of thermoplastic processing methods and is particularly useful where moisture resistance and low cost are required. Polyethylene is limited by low end use temperature characteristics. There are three basic manufacturing processes for making HDPE. The slurry particle reactor process (the most widely used method), the gas phase process, and the new metallocene catalyst technology. This last process also may be used

37

1.2 Thermoplastic Polymers with standard catalysts to combine the advantages offered by both polyoleﬁn polymerization catalysts. A new class of catalysts for polyoleﬁns based on nickel and palladium reportedly can produce a very broad range of molecular weights and branching. Ability to incorporate a wide variety of polar and nonpolar comonomers allows the production of EVA and EMA copolymers in low pressure processes. Combinations of reactors also are used to make HDPE. The tandem reactors produce what are called bimodal molecular weight HDPE. The slurry process and gas phase process can also be used to make HDPE. Unlike LDPE, LLDPE lacks a long chain branching. Another difference is that the ethylene is copolymerized with butene, octane or hexene comonomers in the reactor. LLDPE has a narrower molecular weight distribution than LDPE. A ﬁnal distinction is that LLDPE exits the reactor in a powder form and the granules are compounded with additives in an extruder and pelletized. Blow molding uses the largest amount of HDPE. About 35% is used to make blow molded products. Extruded products consume about 30%, and injection molding accounts for about 20%. High molecular weight (HMW) and HDPE resins have high strength; they are used in packaging ﬁlms, sheets, pipes, and large blow and rotation molding items. Low density polyethylene (LDPE) is the second largest PE capacity in the US. Extrusion is the dominant process used with LDPE resins. Extruded products, principally ﬁlms, account for 75% of LDPE sales. LDPE remains the dominant PE resin for extrusion coating and for food packaging ﬁlms. LDPE is also used for injection molding, blow molding, and rotational molding. High molecular weight (HMW) LDPE ﬁlm grade resins produce high gloss, high clarity ﬁlm that exhibit good toughness and heat sealability. High pressure, high temperature polymerization reactors, both tubular and autoclaves, are used to make LDPE. The molecular structure of LDPE is characterized by long side branches that give the resins their combination of ﬂexibility, clarity, and processability. 1.2.17.1

Ethylene Vinyl Acetate (EVA)

General Properties of Generic EVA Polymers Speciﬁc gravity

0.93

Flexural modulus E (Kpsi)

2.5–14.0

σY (Kpsi)

0.50–2.50

Izod impact @ 73 °F (ft-lb/in)

No break

Thermal limits service temp. (°F)

–

Shrinkage (%)

–

HDT (°F) @ 264 psi

–

Vicat point(°F)

140–200

Tm (°F)

189–223

Process temp. (°F)

285–435

Mold temp. (°F)

60–105

Drying temp. (°F)

120–14-

Drying time (h)

8.0

Figure 1-61 Tennis shoes (Courtesy: Du Pont)

Figure 1-62 Automotive carpet sound backing (Courtesy: Du Pont)

38

1 Polymeric Materials These copolymers are very ﬂexible, tough materials with good adhesion properties. The vinyl acetate content ranges from 1–50% and the grades with more than 20% VA contents are called high EVAs. All thermoplastic end use fabrication processes can be used with EVA products. The resins also are used to make hot melt adhesives and can be compounded with other polymers. 1.2.17.2

Figure 1-63 School bus toy

Ethylene N-Butyl Acrylate (ENBA)

These copolymers are very ﬂexible, tough, and have good tear strength, low temperature toughness, and optical properties. The butyl acrylate content of these resins ranges from 5–20%. ENBA copolymers are used in ﬂexible packaging and extrusion coating applications. These polymers require higher heat processing temperatures, and they have low temperature properties and heat seal properties not provided by EVA copolymers. Optical properties of ENBA copolymers are comparable to those of EVA copolymers. 1.2.17.3

Ethylene Methyl Acrylate (EMA)

These copolymer resins have properties very similar to those of ENBA copolymers, but are available in FDA approved grades. As a result, the EMA copolymer resins are used in ﬂexible food packaging where high clarity, low temperature toughness, and good tear strength are required. Figure 1-64 Blow molded crane and sandbox toy

1.2.17.4

Ethylene Ethyl Acrylate (EEA)

These copolymer resins are very tough and ﬂexible polyoleﬁns, which have excellent low temperature properties. The ethyl acrylate content of these resins ranges from 15–30%. The compounds with the higher percentage have properties similar to rubber materials. EEA copolymers, which are made in modiﬁed high pressure PE reactors, can be injection molded, extruded, and blow molded. The resins are used to make hot melt adhesives and are compounded with other polymers. 1.2.17.5

Figure 1-65 Resin bulk shipping container

These resins have an average molecular weight of 3–6 million. They offer excellent abrasion resistance, a low coefﬁcient of friction, low temperature impact strength and good chemical resistance. However, these resins are difﬁcult to process by conventional processing techniques. Compression molding generally is used to make UHMWPE parts or they are machined from rods, bars, or plates. 1.2.17.6

Figure 1-66 Kitchen ware container

Ultrahigh Molecular Weight Polyethylene (UHMWPE)

Metallocene Polyethylene (MPE)

These are made using standard catalyst technology. They claim to offer better strength, improved heat sealing characteristics, better moisture and oxygen barrier characteristics, high clarity, and greater toughness. Also, the high metallocene catalyst activity is said to virtually eliminate catalyst residues. The MPE resins are more expensive and more difﬁcult to process than LDPE and LLDPE resins. Advantages • Low resin cost • Good impact resistance at temperatures from 104–194 °F • Good properties at low temperatures

39

1.2 Thermoplastic Polymers • Moisture and oxygen resistance • Good chemical resistance • Food grades are available • Readily processed by all thermoplastic methods Disadvantages and Limitations • High coefﬁcient of thermal expansion

Figure 1-67 Low voltage electrical cables

• Poor UV and weathering resistance • Subject to stress cracking • Difﬁcult to bond • Flammable Typical Applications • Packaging: Packaging ﬁlms, rigid and semi-rigid packaging products • Transportation: Automotive fuel tanks • Medical: Hygiene products, medical application trays, and containers • Consumer: Toys, blow molded bottles, bottle caps, household goods, kitchen utensils • Appliances: Portable containers, outdoor furniture, and irrigation • Industrial: Pipes, connectors, buckets, containers, processing equipment, and hardware items for construction • Electrical: Wire and cable insulations

1.2.18

Polytetraﬂuoroethylene (PTFE)

General Properties of Generic PTFE Polymers Speciﬁc gravity

2.13–2.20

Flexural modulus E (Kpsi)

58–108

σY (Kpsi)

3.50–4.50

Coefﬁcient of friction (73 °F)

0.02

Thermal limits service temp. (°F)

500.0

Water absorption (%)

< 0.01

Tm (°F)

648

Tg (°F)

257–266

HDT at 264 psi (°F)

132.0

Process temp. (°F)

690–710

Dielectric constant

2.10

Dielectric strength (V/mil)

500–600

Dissipation factor

< 0.0003

Figure 1-68 Car windshield fluid reservoir

40

1 Polymeric Materials Polytetrafluoroethylene (PTFE) was the first fluorocarbon resin invented by Roy J. Plunkett and introduced by E. I. Du Pont in the late 1940s. The properties of PTFE are unique and versatile. The ﬂuorocarbon resin is created from chloroform by a series of reactions. PTFE is composed of long, straight chains of ﬂuorinated carbons. It is the chemistry of PTFE that provides the unique characteristics of the polymer. The straight backbone of carbon atoms symmetrically surrounded by ﬂuorine atoms provides the resin with its unique chemical, electrical, and thermal properties. Other ﬂuoropolymer resins include ﬂuorinated perﬂuoroethylene-propylene copolymer (FEP), perﬂuoro alkoxy alkane (PFA), ethylenetetraﬂuoroethylene (ETFE), polyvinylidene ﬂuoride (PVDF), chlorotriﬂuoroethylene (CTFE), ethylene chlorotriﬂuoroethylene (ECTFE), polyvinyl ﬂuoride (PVF), and other less important polymers. The maximum end use temperature for PTFE in continuous service is 500 °F. When exposed to ﬂame, PTFE will burn, but it does not continue to burn when the ﬂame is removed, because it has an exceptionally high limiting oxygen index and will not support combustion in air. Like most ﬂuoropolymers, PTFE has outstanding electrical insulation properties. Its dielectric constant and loss factor are very low and uniform across a wide temperature and frequency range. PTFE is probably best known for its low coefﬁcient of friction, lower than most other materials. It does not adhere to anything. PTFE is ﬂexible, strong and tough at temperatures as low as –320 °F. It is an opaque white or bluish-white color, but it can be formed into very thin and transparent ﬁlm. Its tensile strength ranges from 3,500–4,500 psi and its ﬂexural strength is 2,000 psi. The ﬂuorocarbon polymers are available in granular, ﬁne powder, and water base dispersion forms (8 μ in). Granular resins are used for molding and ram extrusion. Fine powders are paste-extruded into thin sections. Dispersions are used for metal coatings and to impregnate porous structures such as fabrics. The molding techniques used for granular PTFE are similar to those for ceramics. The resin starts its transformation at ambient conditions, at pressures between 2,000 and 5,000 psi. The self supporting gel is then sintered at temperatures ranging from 690–710 °F under a controlled temperature cycle. To make tapes or sheets of PTFE (up to 0.250 in thick), a cylinder is molded and veneered. Sintering always follows compression of powder, whether the process is used to mold PTFE into sheets or billets through simple compression molding. To form multiple, simply shaped items, it is recommended to use automatic molding, or isostatic molding equipment. A ﬂexible membrane and hydraulic pressure are used to make complicated shapes. The molecular weight of PTFE is 10 million or higher and it is about 90% crystalline before molding. It will withstand continuous exposure to a temperature near absolute zero (–525 °F) without reaction. The polymer has an initial melting point of 648 °F and although it does not char, it does begin to depolymerize at about 1,142 °F. Polytetraﬂuoroethylene is also highly resistant to blending with ﬁllers and reinforcers. However, one or more fillers can be added to PTFE in some circumstances to prevent creep or cold ﬂow from occurring when a load is applied to the soft material. Fillers are particularly desirable when PTFE is used for parts in a dynamic operation in which a high rate of wear may occur. Bonding between the PTFE and the ﬁller particles is not possible. The ﬁller particles

1.2 Thermoplastic Polymers are just suspended and depend on the relative movement of PTFE particles to improve the ﬂexural modulus, creep, and wear, while decreasing the physical properties of PTFE. The ﬁllers include ﬁber glass, carbon, graphite, bronze, and molybdenum disulﬁde. The PVDF resin can be processed in a typical thermoplastic extruder. To extrude thin wall tubing from ﬁne powder PTFE material, the ﬁne powder should be mixed with a volatile organic lubricant, such as naphtha or mineral spirits, to lubricate the ﬁne powder. The lubricated ﬁne powder is then cold-formed into a cylindrical billet, loaded into an extruder barrel, and driven by hydraulic or pneumatic ram pressure through the die. This process produces a heated continuous extrudate when a reciprocating ram slowly moves the resin through a heated die tubing. As it leaves the die, it passes through a multi-zone oven that removes the solvent and sinters the resin. Only highly corrosion resistant alloys should be used in contact with the melt during processing polytetraﬂuoroethylene. There are two general classes of polymers based on their process capabilities, the ﬂuoropolymers that are able to use traditional melt processing technologies and the ﬂuoropolymers that need special types of processing techniques. Polytetraﬂuoroethylene has a high melt viscosity property and the polymer never becomes ﬂuid. At temperatures above 620 °F it becomes a self supporting gel. Because of this unusual melt characteristic, traditional thermoplastic processing equipment cannot be used. The ﬂuoropolymer families can also be classiﬁed not only by their processing requirements, but by their chemical composition and properties. The characteristics of each ﬂuoropolymer family are described below. 1.2.18.1

Fluorinated Perﬂuoroethylene Propylene (FEP)

A copolymer of hexaﬂuoropropylene and tetraﬂuoroethylene is similar to PTFE in its chemical inertness, dielectric characteristics, coefﬁcient of friction, and resistance to high heat. Its maximum temperature in continuous use is 400 °F. With a lower abrasion resistance than PTFE, it can be processed by most thermoplastic techniques and is available in pelletized grades for extrusion and molding. 1.2.18.2

Perﬂuoro Alkoxy Alkane (PFA)

It is produced when an alkoxy side chain is added onto a base of tetraﬂuoroethylene. It is similar to PTFE in thermal and chemical resistance, low coefﬁcient of friction, and abrasion resistance. It has the same temperature range of PTFE, up to 500 °F. It can be extruded or injection molded, but to avoid stresses, PFA should be processed at high temperatures (700 °F) at slow processing rates. 1.2.18.3

Ethylenetetraﬂuoroethylene (ETFE)

It is a 1 : 1 alternating copolymer of ethylene and tetraﬂuoroethylene. It is resistant to chemicals, exhibiting only a slight swelling when exposed to some chlorinated solvents. Its service temperature ranges from –32 to 300 °F. It has excellent resistance to impact, cut through, abrasion, weather, and radiation. It also has a low dielectric constant and uniform electrical properties. The resin is processed by extrusion and injection molding, as dry nano particles, is used in ﬂuidized bed, rotation molding, and electrostatic coating applications.

41

42

1 Polymeric Materials 1.2.18.4

Polyvinylidene Fluoride (PVDF)

This resin is one of the most rigid and abrasion resistant of the melt processible ﬂuoropolymers. It has good chemical resistance, particularly to permeation by halogens (bromine). It can be attacked by a number of chemicals, such as hot sulfuric acid and hot amines. Some coating grades are solubilized by acetone and ethyl acetate. It has a higher dielectric constant than does PTFE, but a more limited service temperature range. Polyvinylidene ﬂuoride resin is produced by emulsion or suspension polymerization of vinylidene ﬂuoride. It can be extruded and injection molded on typical PVC machines. It is available as dry powder, pellets, and as a solvent base dispersion for coatings. 1.2.18.5

Chlorotriﬂuoroethylene (CTFE)

This resin is a melt processible ﬂuoropolymer that substitutes every fourth ﬂuorine atom on the PTFE chain with chlorine. This ﬂuoropolymer can be used at cryogenic temperatures as low as –350 °F. It retains resilience after exposure to most chemicals, although it is swollen (but not attacked) by chlorinated solvents. It can be molded and extruded, but can be degraded by thermal and mechanical means. The electrical dissipation factor and dielectric constant are higher than PTFE. CTFE has excellent resistance to permeation by chemicals and can be made fully transparent, which lends to its use in packaging applications. 1.2.18.6

Ethylene Chlorotriﬂuoroethylene (ECTFE)

This resin is produced by copolymerizing chlorotriﬂuoroethylene and ethylene. Because of its 1 : 1 alternating copolymer, ECTFE has greater strength, wear and creep resistance than PTFE. It exhibits excellent chemical resistance to acids, bases, and solvents. It is attacked by hot amines and swells slightly in polar organic solvents. It matches PVDF in rigidity and abrasion resistance and surpasses it in impact strength and chemical resistance. Ethylene chlorotriﬂuoroethylene is the easiest to process, whether by melt extrusion, injection, or compression molding, or by dry powder coating and rotational molding. Advantages • Highly inert, no substance has been found that will dissolve PTFE • Nonﬂammable • Excellent resistance to nuclear radiation, ultraviolet rays, and ozone • Excellent chemical resistance, PTFE is unaffected by virtually all chemicals, acids, bases, and solvents • Low creep characteristics • Excellent electrical properties • Good mechanical strength at ambient temperature • Excellent permeability • Excellent oxidative stability • Higher temperature capability (500 °F) • Anti-stick characteristics • Low coefﬁcient of friction Figure 1-69 High performance seals (Courtesy: Du Pont)

• Low mold shrinkage • Low water absorption

43

1.2 Thermoplastic Polymers Disadvantages and Limitations • Poor adhesive characteristics • Low mechanical strength at high temperatures • Comparatively high resin cost • Toxic by-products upon thermal decomposition • Not processable by common thermoplastic methods • Requires corrosion resistant steel for processing equipment

Figure 1-70 Submarine ball valves (Courtesy: Du Pont)

Typical Applications • Aerospace Industry: Low weight and stress crack resistance of PTFE, FEP, and ETFE are used for hose and tubing for hydraulic, fuel, oil, pneumatic, and oxygen systems. Dielectric strength, low dissipation factor, and low dielectric constant are used for printed circuit board laminates in military defence and commercial ﬂight communications equipment. Resistance to severe chemicals, low coefﬁcient of friction, and compressibility is used for sealing applications in turbine engines, alternators, and rotary actuators.

Figure 1-71 Automotive hoses for fuel and air conditioning (Courtesy: Du Pont)

• Automotive: Power steering and transmission seals and rings. Mechanical control cable and fuel hose linings, head gasket coatings. Shaft, compressor, and shock absorber seals. ETFE is extruded into wire harness conduit to protect wire assemblies from high engine temperatures and abrasion. It is also used in shielding applications to serve as a barrier against heat in underthe-hood applications. • Petrochemicals: Lining pipes, valves, pumps, tanks, tubing, ﬁttings, column packing, and processing equipment.

Figure 1-72 Stadium roof covering membrane (Courtesy: Du Pont)

• Medical: Vascular grafts, cardiovascular patches, surgical membranes, soft tissue patches, sutures, ligaments, catheters, and piping systems for the laboratories. • Semiconductor Manufacturing: PFA and PVDF are used for piping, pumps, and valves in clean room ﬂuid handling systems. Filter cartridges with PTFE or PVDF membrane elements are used for ﬁlters, which are supported by other components that are injection molded from PFA or PVDF. Wafer carriers, processing cassettes, wet bench equipment, rinse tanks, counter tops, sinks, and 55 gallon drum liners. PFA is used for packaging and shipping clean room processing chemicals, storage, and transportation of waste chemicals. • Electronic and Electrical: Code approved plenum cable, aircraft and ﬁre alarm equipment, cables and printed circuit board substrates. • Other Applications: Nonstick cookware, coating and impregnation of valve and pump packing, coating on glass cloth provides outstanding weathering and ultraviolet radiation resistance. PTFE coated glass fabric is used as building roof covering material. Fluoropolymers are used as a breathable waterproof in protective clothing, ﬁltration screening, and mist separators.

Figure 1-73 Internal liner valve, corrosion resistant (Courtesy: Du Pont)

44

1 Polymeric Materials

1.2.19

Polyphenylene Sulﬁde (PPS)

General Properties of Generic PPS – 30% GR Resins Speciﬁc gravity

1.38

Tensile modulus @ 73 °F (Mpsi)

1.70

Tensile strength @ yield (Kpsi)

22.00

Notch Izod impact @ 73 °F (ft-lb/in)

1.10

Thermal limits service temp. (°F)

450 (short) 390 (long)

Shrinkage (%)

0.6–3.5

Vicat point (°F)

530

Tm (°F)

540

HDT (°F) @ 66 psi @ 264 psi

534 507

Process temp. (°F)

600–715

Mold temp. (°F)

275–320

Drying temp. (°F)

300

Drying time (h)

3.0–6.0

Polyphenylene sulﬁde is an engineering thermoplastic material with an excellent combination of properties: thermal stability and unusual insolubility, chemical resistance and inherent ﬂame resistance. It is produced commercially by the reaction of 1,4-dichlorobenzene with a suitable sulfur source, such as sodium sulﬁde. PPS is a semi-crystalline, aromatic polymer composed structurally of a series of alternating pare-substituted phenylene rings and divalent sulﬁde moieties. Polyphenylene sulﬁde possesses good thermal stability in which isothermal weight losses as a function of time were measured at several temperatures. These results indicate good resistance to thermal degradation up to 700 °F. Polyphenylene sulfide possesses excellent solvent resistance, being almost totally insoluble in organic solvents below 390 °F. PPS is affected only by high temperature exposure to a few organic solvents, strong mineral acids, and strong oxidizing environments. Oxidizing agents, such as peracetic acid and aqueous sodium hypochlorite, oxidize the sulﬁde moiety to sulfoxide and/or sulfone groups. Because of their chemical structure and ability to char when exposed to an external ﬂame, PPS is inherently ﬂame resistant. It possesses a high oxygen index and a low radiant ﬂame spread index and it is classiﬁed as UL-94 V0 and UL-94 V5. The auto-ignition temperature is 1,000 °F. Glass ﬁber reinforcement produces injection and compression molding compounds that exhibit high tensile strength, good ﬂexural strength, high heat deﬂection temperature, low elongation, and very low impact strength. A variety of pigments and mineral ﬁllers can be added to produce several injection molding compounds, with mechanical properties ranging between the unﬁlled grades and the 50% glass ﬁlled compounds.

Figure 1-74 Automotive water pump impeller

Both ﬁlled and unﬁlled PPS grades exhibit the inherent ﬂame resistance and excellent chemical resistance characteristics of the base resin. Results of longterm heat aging of moldings are consistent with the good thermal stability of the polymer. PPS/glass ﬁber and PPS/mineral ﬁber compounds show good retention

45

1.2 Thermoplastic Polymers of tensile properties under long-term temperature and load exposure. Because of the characteristics of PPS, aging at higher temperatures gives an even more impressive retention of properties owing to a “case hardening” surface effect caused by curing of surfaces exposed to high temperature air. Polyphenylene sulfide and its compounds possess good overall electrical (insulating) properties. The glass ﬁlled compound has a low dielectric constant and dissipation factors are retained over a broad frequency range. In addition, the dissipation factor remains low at temperatures up to 390 °F. Insulating properties are obtained by measuring the volume resistivity and the insulation resistance. These properties are also retained after exposure to high humidity environments. Arc resistance of formulations containing only glass reinforcement is not impressive but can be increased dramatically by proper choice of mineral ﬁllers. The combination of good electrical properties and high temperature resistance has led to UL approval of PPS compounds for use at temperatures (390–450 °F) The addition of a small amount of carbon black provides an adequate ultraviolet protection to the compound.

Figure 1-75 Automotive exhaust sensor

Advantages • Capable of extended usage at 450 °F • Good radiation, solvent, and chemical resistance • Excellent dimensional stability and low water absorption • Non-burning Disadvantages and Limitations

Figure 1-76 Automotive air brake valve

• Difﬁcult to process (high melt temperature) • Comparatively high resin cost • Fillers required to achieve good impact strength • Attacked by chlorinated hydrocarbons Typical Applications • Electrical and Electronic: Sockets, coil forms, bolt yokes, motor brush holders, connectors, integrated circuit, capacitor encapsulations, switches, electronic watch bases, and relay components • Mechanical and Chemical: Pump housings, impeller diffusers, pump veins, end plates, valve components, oil ﬁeld hardware, heat shields, boiler sensors, and ﬂow meters • Appliance: Housings, handles, internal components, microwave oven components, computer disc drives, and heat insulators • Automotive: Emission control system, censoring devices, fuel, ignition and braking systems, light sockets, cooling system, air conditioning units, and generator and alternator components

Figure 1-77 Computer pin connectors

46

1 Polymeric Materials

1.2.20

Polypropylene (PP)

General Properties of Generic Unﬁlled PP Homopolymer Speciﬁc gravity

0.90

Tensile modulus @ 73 °F (Mpsi)

0.17

Tensile strength @ yield (Kpsi)

4.00

Notch Izod impact @ 73 °F (ft-lb/in)

0.5–18.0

Thermal limits service temp. (°F)

212

Shrinkage (%)

0.5–2.0

Vicat point (°F)

320

Tm (°F)

329–338

HDT (°F) @ 264 psi

120–140

Process temp. (°F)

390–525

Mold temp. (°F)

85–175

Drying temp. (°F)

175

Drying time (h)

2.0–3.0

Polypropylene was introduced in the late 1950s and is the fastest growing commodity thermoplastic in the world. It is a versatile polymer used in applications from ﬁbers, ﬁlms, appliances, to automobile bumpers. PP continues to displace other materials, such as ﬁber glass, mineral reinforced thermoplastics and metals, in a variety of applications. Figure 1-78 Variety of injection molded products

Polypropylene is manufactured by polymerizing propylene monomer with a titanium based catalyst, a second co-catalyst (triethylaluminum) is added to initiate the polymerization reaction and hydrogen is used in the reactor to control polymer molecular weight. This reaction is produced using a slurry or gas phase type of process. There are three PP structures: isotactic, syndiotactic, and atactic. The principal structure of PP is isotactic semi-crystalline in a helical form. This structure has good mechanical properties, such as stiffness and tensile strength. These properties can be further increased with nucleating agents or with ﬁllers, such as talc, calcium carbonate, or ﬁber glass.

Figure 1-79 Snap-top closure

Syndiotactic PP is produced by the monomer units inserted alternately head-totail. This structure is more ﬂexible than the isotactic form but has better impact resistance and clarity. Atactic PP (hard wax amorphous monomer) is a by-product of the manufacturing process. This product is used in rooﬁng tars and adhesives for the shoe industry. All forms of PP are susceptible to oxidation caused by the presence of a tertiary hydrogen. Polypropylene is stabilized against thermal degradation by the addition of primary and secondary antioxidants. Neutralizing agents are also added to stabilize the low levels of chloride ash generated during manufacturing. Other special additives are used, such as antistatic agents, slip agents, and UV stabilizers.

Figure 1-80 Compact disk jewel boxes

Polypropylene is sold commercially as homopolymers, random copolymers, or impact resistance copolymers. Physical properties range from high strength, stiffness, to a ﬂexible polymer with lower strength but greater toughness.

47

1.2 Thermoplastic Polymers The homopolymer has the highest melting point and stiffness with a wide range of melt ﬂow properties. The copolymers incorporate small amounts of ethylene that lower the crystallinity, producing improved impact resistance properties, more ﬂexibility, but a lower melting point. Impact resistant copolymers are copolymerized in a reactor by adding ethylene. The copolymer acts as a plasticizer and is evenly dispersed throughout the homopolymer matrix to form a heterophasic polymer. This copolymer has very high impact resistance even at low temperatures. High impact resistance copolymers are produced by compounding the pre-blend of copolymer, additives, and ethylene-propylene or EPDM rubber. Advantages • Lighter or low density polymer (0.90 g/cm3) • High melting point (329–338 °F) • End use temperatures of 212 °F • Good chemical resistance to hydrocarbons, alcohols, and non-oxidizing reagents • Good fatigue resistance (integral life hinge closures) • Processed by all thermoplastic methods: injection molding, compression molding, blow molding, extrusion, cast ﬁlms, and thermoforming

Figure 1-81 Automobile injection molded dash panel

Disadvantages and Limitations • Degraded by UV • Flammable, but ﬂame retardant grades are available • Attacked by chlorinated solvents and aromatic solutions • Difﬁcult to bond • Several metals accelerate oxidative degrading Typical Applications • Flexible packaging ﬁlms • Biaxially oriented packaging ﬁlms

Figure 1-82 Hot water dispenser bottles

• Stretched and oriented monoﬁlament, tapes for textiles, carpeting, insulated medical fabrics and woven carpet backing • Automotive interiors, bumpers, spoilers, air vent systems, under the hood components, internal wheel guards, and bellows • Hygiene products, household goods and medical application trays, strainers, and containers • Consumer products, such as closures, over caps, trigger sprayers, rigid and semi-rigid packaging, video cassette cases, toys, and electrical hardware • Appliance housings and components, outdoor furniture, and luggage • Injection blow molded stretch bottles with excellent stiffness, impact resistance, and clarity

Figure 1-83 Electric rice cooker

48

1 Polymeric Materials

1.2.21

Polystyrene (PS)

General Properties of General Purpose PS Polymers Speciﬁc gravity

Figure 1-84 Watch packaging

1.05

Tensile modulus @ 73 °F (Mpsi)

0.45

Tensile strength @ yield (Kpsi)

6.0

Notch Izod impact @ 73 °F (ft-lb/in)

0.25–0.60

Thermal limits service temp. (°F)

158 (short) 122 (long)

Shrinkage (%)

0.05–0.80

Vicat point (°F)

200–227

Tm (°F)

212

HDT (°F) @ 264 psi

190

Process temp. (°F)

390–480

Mold temp. (°F)

50–175

Drying temp. (°F)

160–200

Drying time (h)

2.0–3.0

Polystyrene (PS) has been known for well over 100 years, but its molecular nature was not clariﬁed until about 1920, when the work of Staudinger described the material’s molecular structure. Polystyrene has been commercially produced since the late 1930s. PS is one of the most popular commodity amorphous thermoplastic resins; it has a broad range of balanced properties and an attractive price. Polystyrene is divided into semi-crystalline (general purpose) polystyrene (GPPS), rubber modiﬁed medium and high impact polystyrene (MIPS and HIPS), and expandable polystyrene (EPS). The raw materials for polystyrene are ethylene and benzene which react in a process to form ethylbenzene, which is further processed into styrene monomer; other feed stocks are acrylonitrile and butadiene rubber. A polymerization process using a thermal or catalyzed reaction of styrene monomer is used to produce an amorphous polymer. Other materials are added to the process such as rubbers, plasticizers, release agents, and stabilizers to give the polymers the desired characteristics. The formulation may also include colorants, ﬂame retardants, UV stabilizers, or impact modiﬁers. GPPS is typically selected for its clarity, rigidity, and suitability for many applications. HIPS and MIPS are used where more ﬂexibility or impact resistance is required; HIPS and MIPS contain butadiene rubber as the copolymerization agent to increase toughness, it also makes it opaque in color. Advantages

Figure 1-85 Perfume bottle

• • • • • • •

Optical clarity High gloss FDA grades are available Processable by all thermoplastic methods Low cost Good dimensional stability Good rigidity

49

1.2 Thermoplastic Polymers Disadvantages and Limitations • Flammable, but ﬂame retarded grades are available • Poor solvent resistance, attacked by many chemicals • Homopolymers are brittle • Subject to stress and environmental cracking • Poor thermal stability Typical Applications • Single service items, such as plates, glasses, and cups • Packaging items such as cassette boxes and compact disc jewel boxes

Figure 1-86 Kitchen timer

• Consumer durables, such as housewares and cosmetic containers • Blow molded medical and pharmaceutical packaging • Extruded solid sheets, foamed or biaxially oriented sheets for thermoforming; blends with styrene butadiene rubber-block copolymer are used where clarity and toughness are desired • Shower curtain sheets, easy to color with a printable surface • Foamed food packaging articles, such as trays, takeout containers, building insulation, and construction materials • Oriented polystyrene food contact articles, such as cookie containers and chocolate trays • Molded parts and components for refrigerators and other appliances, consumer durables, such as toys, housewares, video cassettes, and micro ﬂoppy diskette housings

1.2.22

Polysulfone (PSU)

General Properties of PSU – 30% GR Polymers Speciﬁc gravity

1.46

Tensile modulus @ 73 °F (Mpsi)

1.35

Tensile strength @ yield (Kpsi)

14.50

Notch Izod impact @ 73 °F (ft-lb/in)

1.10

Thermal limits service temp. (°F)

375 (short) 350 (long)

Shrinkage (%)

0.50–0.80

Vicat point (°F)

363–387

Tg (°F)

373

HDT (°F) @ 66 psi @ 264 psi

360 350

Process temp. (°F)

600–715

Mold temp. (°F)

200–355

Drying temp. (°F)

200–320

Drying time (h)

3.0–4.0

Figure 1-87 Wine rack

50

1 Polymeric Materials Polysulfone (PSU) is produced from Bisphenol A (BPA) and 4,4-dichlorodiphenylsulfone by a nucleophilic process. It is an amorphous engineering thermoplastic material with exceptionally high temperature resistance, high rigidity; it is transparent and dimensionally stable. It exhibits a signiﬁcant reduction in Izod impact strength when notched; its continuous use temperature index is 350 °F. PSU has a glass transition temperature (Tg) of 373 °F. Polysulfone has been approved to be used in several medical and laboratory applications. It exhibits good chemical resistance but is sensitive to polar solvents and other solutions. Polysulfone is in compliance with the gamma sterilizability requirements. Unlike many other polymers, the melt viscosity of PSU is relatively insensitive to shear forces. The low degree of molecular orientation during injection molding slightly affects the physical properties of the molded parts, independent of ﬂow directions of the melt ﬂow. Polysulfone sheets or molded parts may be joined to another PSU surface or to metals by direct heat sealing, adhesive bonding, solvent fusion, hot plate welding, ultrasonic welding, or heat sealing at 700 °F. Solvent fusion is accomplished with a 5% solution of PSU in a dope of methylene chloride, followed by drying at elevated temperatures as the parts are held together. Polysulfone can be plated by an electrolysis process that imparts bond strengths. Both nickel and copper electroless processes can be employed. All molded parts must be annealed before plating. Polysulfones also can be vacuum-metallized using standard equipment and techniques. Polysulfone should be dried before processing to prevent voids, surface streaks (jetting), and splay marks caused by volatilization of absorbed water. Moisture content should be reduced to below 0.05% before processing the resin. Advantages • Heat deﬂection temperature of 345 °F at 264 psi • UL continuous use temperature index of 350 °F • High resistivity and dielectric strength properties • High resistance to burning without ﬂame retardant additives • Excellent thermal stability • Excellent resistance to thermal degradation • Good chemical resistance • Highly resistant to alkalies, salt solutions, and aqueous mineral acids • Conventional thermoplastic techniques are used for transformation Figure 1-88 Artificial heart components

• Both thermosetting and thermoplastic types of polymers are available Disadvantages and Limitations • High resin cost • Poor weatherability, subject to stress cracking • Soluble in chlorinated aliphatics, such as methylene chloride, chloroform, and trichloroethylene

Figure 1-89 Syringe holder device

• It is partially dissolved in esters and ketones

51

1.2 Thermoplastic Polymers • Low notched Izod impact strength • Viscosity is very temperature sensitive • Processing requires high injection molding pressures, high barrel temperatures, and high mold temperatures • It absorbs water (0.85%) Typical Applications • Construction and Building: Plated plumbing ﬁxture and insulators

Figure 1-90 Tile flooring

• Major Appliances: Microwave oven components • Small Appliances: Plated hair dryer handles • Medical: Pans, hospital equipment, electronic parts, packaging enclosures, ﬁlters cases, and laboratory supplies that require sterilization • Housewares: Microwave oven trays, bowls, and covers

1.2.23

Polyvinyl Chloride (PVC)

Figure 1-91 Industrial boots

General Properties of Generic Rigid PVC Polymers Speciﬁc gravity

1.38

Tensile modulus @ 73 °F (Mpsi)

0.35

Tensile strength @ yield (Kpsi)

6.00

Notch Izod impact @ 73 °F (ft-lb/in)

0.40–20.0

Thermal limits service temp. (°F)

221 (short) 140 (long)

Shrinkage (%)

0.10–2.50

Vicat point (°F)

179–216

Tm (°F)

360–390

HDT (°F) @ 264 psi

140–170

Process temp. (°F)

365–400

Mold temp. (°F)

85–140

Drying temp. (°F)

160–180

Drying time (h)

2.0–3.0

Polyvinyl chloride (PVC) was introduced in the early 1930s to become a very popular material in the building and construction industry because of its properties, its competitive cost, wide processing capability, and because it can be recycled.

Figure 1-92 Structural window frames

Figure 1-93 Garage door

The two major categories of PVC resins are available: the homopolymer suspension resins and the dispersion resins. Suspension resins account for more than 90% of the total PVC market. They are produced as white powders containing coarse porous particles. When mixed with additives, the suspension resin becomes a powder blend. Suspension resins can be produced as either rigid resins for non-plasticized applications or ﬂexible resins for plasticized applications. All PVC resins require heat stabilizers to allow processing without degrading and discoloring the polymer. Plasticizers are added to increase the ﬂexibility

Figure 1-94 Flat wire cable

52

1 Polymeric Materials of the compound. They can also improve the heat stability or improve the ﬂame retardancy of the compound. Fillers are used to reduce the cost of the compound and improve dimensional stability (stiffness) and impact strength of the polymers. Suspension PVC resins may be blended using two types of mixers: the high intensity mixer and the ribbon blender. High intensity blenders and twin screw extruders are used for producing rigid resins. Ribbon blenders and single screw extruders are used for producing ﬂexible resins. Figure 1-95 Electric cords

Dispersion resins account for 7% of the total PVC market. They have a very ﬁne particle size of about 1 micron. When dispersed in plasticizers and other liquid ingredients, they form plastisols and organosols that are applied as liquid (paste) coatings and then fused with heat. Advantages • Processed by all thermoplastic methods • Wide range of ﬂexibility, made possible by varying levels of plasticizer • Relatively low cost • Not ﬂammable

Figure 1-96 Packaging

• Good resistance to weathering • Dimensional stability • Excellent resistance to water and aqueous solutions Disadvantages and Limitations • Attacked severely by stronger solvents, such as aromatic hydrocarbons, ketone, esters, and chlorinated solvents

Figure 1-97 Electrical plug

• Limited thermal capability • Thermal degradation of the polymer generates hydrochloric acid • Stained by sulfur compounds • Higher density than many other plastics Typical Applications • Construction: Pipe, conduit, ﬁttings, siding windows, doors

Figure 1-98 Tubing and connections

• Residential and Commercial Building Constructions: Flexible vinyl sheet and tile ﬂooring, wire insulation, electrical tape, rigid molded outlet boxes and covers. Wall coverings, vinyl coated fabrics and furniture upholstery sheeting, shower curtains, window shades and blinds, wall paneling, picture frames, garden hose, lawn edging, swimming pool liners, and weather stripping • Recreational and Sporting Goods: Toys and athletic footwear • Consumer Use: Appliances, luggage, handbags, baby pants, footwear, tablecloths, adhesive tape, labels, notebook covers. and credit cards • Automotive Applications: Instrument door panels, upholstery, arm rests, center consoles, vinyl tops, body moldings, ﬂoor mats, sealants, and caulks

Figure 1-99 Large pipe fitting

• Packaging and Medical Applications: Clear and opaque bottles, ﬂexible and rigid packaging, clear food wrap ﬁlm, medical gloves, blood bags, tubing, and test equipment trays

53

1.2 Thermoplastic Polymers

1.2.24

Styrene Acrylonitrile (SAN)

General Properties of Generic Unreinforced SAN Polymers Speciﬁc gravity

1.06–1.08

Tensile modulus @ 73 °F (Mpsi)

0.40–0.56

Tensile strength @ yield (Kpsi)

9.0–12.0

Notch Izod impact @ 73 °F (ft-lb/in)

0.40–0.60

Thermal limits service temp. (°F)

175 (short) 190 (long)

Shrinkage (%)

0.30–0.50

Vicat point (°F)

219–235

Tg (°F)

212–250

HDT (°F) @ 66 psi @ 264 psi

165–200 170–205

Process temp. (°F)

360–550

Mold temp. (°F)

30–175

Drying temp. (°F)

160–190

Drying time (h)

2.0–4.0

Styrene acrylonitriles (SAN) are random, amorphous, linear copolymers, produced by copolymerizing styrene and acrylonitrile. The characteristics of SAN copolymers are: transparency, excellent thermal properties, good chemical resistance, hardness, dimensional stability, and load bearing capabilities. Two rubber-modiﬁed versions of SAN are oleﬁn-modiﬁed SAN (OSA) and acrylic styrene acrylonitrile (ASA). Both polymers are softer than unmodiﬁed SAN; they are ductile and opaque materials. The two-phase terpolymers are commonly known as weatherable polymers. OSAs are produced by combining with a grafted SAN, saturated elastomeric oleﬁn rubber backbone. ASA is a random amorphous terpolymer, produced either by a mass copolymerization process or by grafting styrene-acrylonitrile to the acrylic elastomer backbone. The resulting terpolymers (OSA and ASA) are tough and ductile. These polymers are extremely weather resistant. UV absorbers and antioxidants are used to enhance weatherability. SAN resins are normally used unmodiﬁed, although SAN can also be reinforced with glass ﬁber to produce materials with extremely high rigidity and enhanced thermal properties. Alloys and blends are also available. The alloy SAN/PVC improves outdoor weatherability and ﬂame retardancy. The alloy SAN/PC enhances processibility, while retaining the physical properties of PC. The alloy SAN/butyl acrylate enhances outdoor stability. The alloy OSA/butyl acrylate offers improvements in processibility, toughness, and retention of properties with weathering. Because these polymers are hygroscopic, they absorb a small amount of moisture. Drying is recommended before processing, especially after long-term storage in a warm, humid environment. Drying for 2–4 hours at 160–190 °F with a dehumidifying hopper dryer system is suggested.

54

1 Polymeric Materials Advantages • SAN is optically clear, hard, and rigid with excellent dimensional stability and good environmental stress cracking resistance. • OSA and ASA are opaque, rigid, ductile products, similar to ABS but with superior outdoor weathering characteristics and surface stability. • OSA is ideal for coextrusion as a protective cap layer over various substrates. • Improved solvent resistance over polystyrene Figure 1-100 Insulin kit case

Disadvantages and Limitations • Hygroscopic polymer requires predrying • Highly susceptible to the effects of structure orientation • Annealing heavy wall thickness parts reduces the molded-in stresses • Low thermal capability (175 °F) Typical Applications for SAN • Automotive: Instrument lenses, instrument panel supports, battery caps, probes, and cases • Construction and Building: Window panels, storm door glazing panels, plumbing ﬁxture knobs, aerosol nozzles, and paint jars

Figure 1-101 Disposable lighter

• Appliances: Washing machine drain connectors, control panels and doors • Medical: Intravenous connectors, ﬁlter cases, and blood dialysis units • Electronics: Video cassette hubs and lenses, telephone parts, terminal boxes, and dust covers • Housewares: Glasses, bowls, mugs, cake covers, brush handles, bristles, and bathroom accessories • Packaging: Display racks, cosmetic cases, lipstick tubes and closures • Miscellaneous: Industrial battery cases, home water ﬁlters, swimming pool pump components, and disposable cigarette lighters Typical Applications for OSA and ASA

Figure 1-102 Coffee dispenser

• Automotive: Body moldings, exterior and interior trim part, bumper parts, recreational vehicle components, and pickup truck caps • Construction and Building: Home siding, trim, down-spouts, gutters, fencing, shutters and window frame components • Recreational and Leisure: Swimming pool and pump components, snowmobile housings, and outdoor furniture

1.3 Thermoplastic Elastomers (TPE)

1.3

Thermoplastic Elastomers (TPE)

A thermoplastic elastomer (TPE) is a rubbery material with the characteristics of a thermoplastic and the performance properties of a thermoset rubber. TPEs are processed by using the same thermoplastic equipment and methods,, such as extrusion, injection molding, and blow molding. The TPE resins combine the properties of rubber and thermoplastics. Charles Goodyear started the rubber industry in the 1840s, and John Wesley Hyatt, who is considered the father of the plastic industry, introduced the ﬁrst plastic material in 1860. The rubber and the plastic industries have each grown and prospered in different directions, developing their materials, technology, fabrication methods, and markets. The fast development and commercialization of applications using thermoplastic elastomer resins have forced an interaction between the rubber and the plastic industries and the barriers between these two industries are shrinking continuously. The ﬁrst TPE introduced to the industry was the thermoplastic polyurethane elastomer resin during the late 1950s. The styrene butadiene and styrene isoprene block copolymer resins were introduced in 1965. The introduction of the thermoplastic block copolyester elastomer resins was in the early 1970s. The block copolyester TPE was a signiﬁcant technical innovation, increasing the use of these materials in new and critical applications. The next group of TPEs to be introduced was a blend of rubber and plastic (polypropylene and EPDM rubber) by the late 1970s. The elastomeric alloy thermoplastic vulcanizates (TPV) were introduced in the 1980s. The TPV vulcanization is based on oil dispersion, ﬁllers, and an elastomeric phase (EPD) in a continuous phase of (PP) polyoleﬁn. The elastomeric alloy melt processible rubbers (MPR) were introduced in 1985. The thermoplastic polyamide elastomers (ﬂexible nylons) are another high performance class of TPEs, introduced in the late 1980s. Advantages Thermoplastic elastomers offer a variety of advantages over the thermoset rubbers: • Most TPEs are fully formulated, compounded, and ready for processing. Mixing, blending, or compounding is not required. • TPEs have the processing simplicity of a thermoplastic, giving more efﬁcient processing conditions and signiﬁcantly lower cost for the ﬁnished component. • TPEs offer fast molding cycles, reducing the manufacturing cost of the molded part. The TPE molding cycle is measured in seconds, compared to minutes for compression molding and vulcanization of thermoset rubbers. • Recycling of reject parts and runners (reground). The scrap material from the thermoset rubber process usually is discarded. Some thermoset manufacturing methods could generate large amounts of scrap materials equal or greater to the weight of the ﬁnished part. The reground from TPE processes can be recycled (25% maximum ratio) to give ﬁnished parts with properties as good as the virgin material.

55

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1 Polymeric Materials • Lower energy consumption as a result of lower injection molding cycle characteristics obtained while processing TPEs. • Better quality control and closer part dimensional tolerances obtained by injection molding parts made from TPE resins. • Lower manufacturing costs because of greater reproducibility and consistency in properties of the TPE resins. • TPEs have a lower density than most thermoset rubbers, giving additional cost savings. Both rubber and plastic are processed on a volume basis; however, both materials are bought on a weight basis. Disadvantages Thermoplastic elastomers have some disadvantages over thermoset rubbers: • New Technology: Signiﬁcant thermoset rubber technology innovations have been developed for the industry (better materials, efﬁcient equipment, automatic running molds, and faster processing cycles) to reduce manufacturing costs of the articles. • Unfamiliar Processing Equipment: The equipment required to process TPE resins is familiar to the thermoplastic processors, but this equipment is foreign to the conventional thermoset rubber fabricators. • Drying TPE resins before processing requires additional new drying equipment for the rubber processors. They also need to learn the critical operations required to process TPEs. Conventional rubber fabricators have never used any type of dryer to process thermoset rubbers. The thermoplastic fabricators are very familiar with the dryers and the manufacturing problems caused by processing a wet TPE resin. • The manufacturers offer a limited number of low hardness TPEs. The great majority of commercially available TPEs have a hardness above 60 Shore “A”. The number of available commercial TPE grades with hardness as low as 35 Shore “A” is limited. These special TPE soft grades are slowly becoming commercially available for some types of TPE alloys. • Melting the TPE resins at elevated temperatures and high injection pressures is required for the injection molding process. A thermoset rubber, on the other hand, requires low injection pressure and low temperature provided by the mold.

1.3.1

Thermoplastic Elastomer Families

The property characteristics of thermoplastic elastomers are based on the chemistry and morphology of their matrix. Thermoset rubber materials generally are reinforced with carbon black. In TPEs, the polymer matrix itself provides this reinforcement. Sometimes, it is necessary to modify the matrix structure’s molecular chain ramiﬁcations to obtain the required stiffness or hardness for the product. Chemical composition and morphology provide a rational, convenient means of categorizing the existing commercial thermoplastic elastomers. The following commercial TPE families are available: • Thermoplastic polyurethane elastomers (TPU) • Styrenic block copolymer thermoplastic elastomers (SBS)

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1.3 Thermoplastic Elastomers (TPE) • Thermoplastic elastomer alloys: – Polyoleﬁn thermoplastic elastomeric (TPO) – Elastomeric alloy thermoplastic vulcanized (TPV) – Melt processible rubbers (MPR) • Block copolyester thermoplastic elastomers • Polyamide thermoplastic elastomers To compare the performance characteristics between the different TPE resins and the thermoset rubbers, we need to analyze the performance required for applications and the manufacturing costa. The cost and performance characteristics of the generic TPE categories are rated in the following order: starting with styrenic block copolymers, polyoleﬁn blends, elastomeric alloys, thermoplastic polyurethanes, block copolyester elastomer, and polyamide elastomer. The thermoset rubbers also increase in both cost and performance similar to the TPE materials. When comparing thermoplastic elastomers to the corresponding thermoset rubbers, it is important to remember that the processing costs of the TPEs are signiﬁcantly lower than those for thermoset rubbers, because of the efﬁcient, fast, and automatic molding process.

1.3.2

Thermoplastic Polyurethane Elastomer (TPU)

General Properties of TPU Elastomeric Polymers Speciﬁc gravity

1.12–1.21

Tensile modulus @ 73 °F (Mpsi)

5–12 (soft) 15–145 (hard)

Tensile strength @ break (Kpsi)

1.5–7.5 (soft) 5–114 (hard)

Compression set w/o annealing (%)

60–80

Elongation break @ 73 °F (%)

450–600 (soft) 160–450 (hard)

Shore durometer hardness

70A – 98A (soft) 40D – 75D (hard)

Thermal limits service temp. (°F)

230 (short) 176 (long)

Shrinkage (%)

1.3–1.5 (soft) 0.6–0.85 (hard)

Tg (°F)

250 (soft) 320 (hard)

Process temp. (°F)

380–430 (soft) 390–450 (hard)

Mold temp. (°F)

50–150

Drying temp. (°F)

200–230

Drying time (h)

2.0–4.0

Thermoplastic polyurethane elastomers are a member of the family of polyurethanes, which were discovered by Otto Bayer in 1937. Originally, polyurethane elastomers (formed by the casting technique) were considered a material that was chemically crosslinked like the thermoset materials. Not until the late 1950s was it found that essentially linear polyurethane elastomers based on

58

1 Polymeric Materials 4,4-diphenylmethane diisocyanate could be processed like thermoplastics. The segmented structure of the polyurethane elastomer contained crystalline hard segments. This led to a completely new model for the scientiﬁc understanding of the elastomeric materials. The basic ingredients of thermoplastic polyurethanes (TPUs) are diisocyanates and long-chain and short-chain diols. The diisocyanates and short-chain diols become the basis of the hard segment structure, while the long-chain diols provide the basis of the soft segments. Because the hard and soft segments do not mix, TPU exhibits a two-phase structure. The properties of TPU can be attributed to the formation of domain micro structures. The basis of polyurethane chemistry is the reaction of isocyanates with various active hydrogen elements present in the compounds. The higher functionality of the isocyanates and the active hydrogen present in the comnpound must be used to obtain higher weight molecular polyurethane. Primarily, TPU linear polymers, consisting of a large number of urethane groups, are synthesized by the condensation of diisocyanates with short-chain diols and polyester or polyether diols. Thermoplastic polyurethanes can be categorized into polyester and polyether types. Polyester-based TPUs generally have better physical properties, such as thermo-oxidative stability and oil resistance. At a similar hardness, polyetherbased TPUs exhibit better low temperature properties, such as hydrolytic stability and resistance to microbial attack. A wide variety of TPUs, with hardness ranging from 70–90 Shore “A” and 40–75 Shore “D”, are available in either pellet or granular form. TPUs consist of an amorphous phase (soft block) and a crystalline phase (hard block), owing to the incompatibility of the adducts of short-chain versus longchain diols with diisocyanates. The hard segments are generally dispersed in the amorphous phase (continuous phase). In a typical TPU, the hard segments determine the hardness, the ﬂexural modulus, the tear strength, and the upper use temperature, while the soft segments of the compound determine the elastic characteristics and the low use temperature properties. The load bearing properties of TPUs are a direct function of the hardness; the higher the hardness of a TPU, the better its load bearing properties. The compression set property of a polymer is its elastic recovery behavior under a speciﬁc loading or a speciﬁc deﬂection at various times and temperatures. Typical compression set values of annealed TPUs range from 25–50% and from 60–80% without annealing. Compression set under 25% is obtained with a deﬂection at 158 °F cured for 22 hours. TPUs are known to exhibit excellent resistance to abrasion. However, the abrasive wear of a TPU is considerably affected by the surface heat buildup during the test, which is believed to be related to the coefﬁcient of friction, stress loading, and contact areas. The abrasive wear of a lubricated TPU is generally lower than that of an unlubricated one. The thermal stability of TPUs is strongly dependent on the structures of the isocyanates and chain extenders. Most TPUs are decomposed slowly at 302– 398 °F and at a measurable rate at 398–482 °F. The TPUs exhibit a loss of mechanical properties and discoloration upon exposure to sunlight. The UV stability of TPUs can be improved by the addition of UV stabilizers. Carbon black-pigmented TPUs have also been found to exhibit better UV stability.

59

1.3 Thermoplastic Elastomers (TPE) Although the hydrophilic characteristics of TPU materials have prevented their use in applications where consistently high electrical insulation resistance properties are required, the use of TPUs as a protective covering for various cable applications has been increasing because of their ﬂexibility and abrasion resistance. One advantage of TPUs is their ease of processing. They can be inexpensively processed into simple or complex semi-ﬁnished and ﬁnished goods by a variety of methods. The extrusion and injection molding techniques are two of the most important processes for the fabrication of TPUs. Other methods include calendering and solution fabrication. Thermoplastic polyurethanes absorb moisture rapidly upon exposure to atmospheric air; so they must be dried before they can be converted into acceptable ﬁnished goods. Excessive moisture content in TPUs will result in molding and extrusion difﬁculties. Blisters, splay marks, bubbles, foamy melt, nozzle drool, and poor physical properties are typical injection molding problems, if the TPU resins are processed wet without drying. During the extrusion process, bubbles, poor surfaces, wave forms, surging, and degradation problems are observed. Although the amount of moisture that can be tolerated in a TPU during processing may vary with the application, it is generally suggested that the moisture content be below 0.07%. Therefore, proper drying of the TPU just before processing is strongly recommended. One recommended way of drying TPUs is to use hopper dryers, which make use of hot air having a dew point below 0 °F and supply drying air at a ﬂow rate greater than 1 lb/h The drying temperatures for TPUs normally range from 203–230 °F with residence times of 2–4 hours. Generally, the softer the TPU is, the lower the drying temperature should be. Prolonged drying should be avoided to prevent TPU discoloration. Advantages • Good physical properties

• Good thermo-oxidative stability

• Good oil, grease, and gasoline resistance • Moderate temperature resistance • Good hydrolytic stability

• Good resistance to microbial attack

Figure 1-103 Snorkels

• Excellent abrasion, wear, and tear resistance Disadvantages and Limitations • High resin cost

• Low aliphatic hydrocarbon resistance

• Low aromatic hydrocarbon resistance • High compression set properties

• Most solvents cause swelling and/or degradation • Requires UV stabilizers for outdoor applications • Poor electrical properties

• Requires drying the resin below 0.07%moisture content before the extrusion process

Figure 1-104 Swim fin

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1 Polymeric Materials Typical Applications • Automotive: Bushings and seals, steering gear parts, shock-absorber bumper, fender extension, membranes of hydro-pneumatic suspension system, grill, bumper ﬁller, and valance panel • Hosing and tubing: Irrigation hose, garden hose, ﬁre hose, hydraulic hose, sewer hose, fuel line hose, fuel line tubing for snowmobiles and small gas engines, pneumatic tubing, and medical tubing • Wire and cable: Air gun control cable, audio wire, camera cable, computer cable, head-set wire, interior building wire, marine cable, welder cable, military cable, communication wire, and jacketing Figure 1-105 In-Line skate wheels

• Casters: Shopping carts, food service carts, and hospital cart wheels • Film and sheet: Elastic leg and waist bands for disposable diapers, life vests, life rafts, life jackets, football liners, seals, gaskets, diaphragms, balloons, air mattresses, water beds, rainwear, surgical drapes, wheel chain pads, conveyor belts, disposable gloves, adhesive ﬁlm for foam and fabrics, protective covers for industrial usage, and foam cushions • Industrial: Shoe soles and heels, belts, drive gears, wheels

Figure 1-106 Caster wheels

• General Purpose: Reinforced and unreinforced drive belts, packaging seals and gaskets, asphalt paving equipment, soft face hammer heads, wear strips, ski masks, foot holders for wind surﬁng boards, ski pole handles, mining screens, boots, and soles

1.3.3

Styrenic Block Copolymer (SBS)

General Properties of Generic SBS Speciﬁc gravity

0.90–1.28

Flexural modulus @ 73 °F (Mpsi)

140–1,400

Tensile strength (psi)

560–3,500

Compression set w/o annealing (%)

–

Elongation at break @ 73 °F (%)

250–820

Shore durometer hardness

36 A – 95 A 36 D – 60 D

Thermal limits service temp. (°F)

–110 to 140

Shrinkage (%)

0.3–0.6

Melt point (°F)

330

Process temp. (°F)

350–440

Mold temp. (°F)

75–90

Drying temp. (°F)

160

Drying time (h)

2.0–3.0

Styrenic thermoplastic elastomers represent a unique class of materials introduced about 1965. They have many of the physical properties of vulcanized rubbers (softness, ﬂexibility, resilience) but are processed as thermoplastics. Styrenic thermoplastic elastomers can be processed on conventional plastic processing equipment, e.g., with injection molding, and scrap usually can be

61

1.3 Thermoplastic Elastomers (TPE) recycled. The high production rate and low resin cost have given the styrenic thermoplastic elastomers a signiﬁcant boost in the thermoplastic elastomers market growth. However, because the transition to the ﬁnal form is reversible, some end use properties of styrenic thermoplastic elastomers (e.g., compression set, solvent resistance, and upper service temperature) are usually not as good as those of thermoset rubbers. Styrenic thermoplastic elastomers are used in areas where these end use properties are less important (e.g., footwear, wire insulation, adhesives, polymer blending, sealants, and coatings). Styrenic thermoplastic elastomers have thermoplastic properties because of their structure. They are multi-phase compositions in which the phases are chemically bonded by block copolymerization. At least one phase is a styrenic polymer that is hard at room temperature but becomes ﬂuid when the polymer is heated, whereas another phase is a softer material that is rubber-like at room temperature. Most polymers are thermodynamically incompatible with other polymers, and such mixtures separate into two phases. This characteristic exists even when the polymeric matrix is part of the same molecule, as in the case of the poly(styrene-b-elastomer-b-styrene) block copolymers. Most of the polymer molecules have their end polystyrene segments in different domains. At room temperature, these polystyrene domains are hard and act as crosslinks, tying the elastomer chains together in a three-dimensional network. But in thermoplastic elastomers the domains lose their strength when the material is heated or dissolved in solvents, allowing the polymer to ﬂow. When the melt is cooled or the solvent is evaporated, the domains become hard again and the network regains its original integrity. The most important property of these polymers for commercial purposes is their resemblance, at least at room temperatures, to vulcanized rubbers. Styrenic block copolymers have a tensile strength from 560–3,500 psi and elongation at break from 250–820% at room temperature. There are two reasons for these high values. The ﬁrst one is that the hard polystyrene domains act as reinforcing ﬁller, and the second takes into account the increased tensile strength, resulting from the slippage of entangled chains. The materials containing a constant polystyrene structure have demonstrated that the tensile module and tensile strengths of the styrenic block polymers are not molecular weight dependent. As long as the polystyrene molecular weight structure is high enough to cause the formation of a strong bond and a well separated structure domain. These materials must be protected against oxidative degradation and in some cases against sunlight also, depending on their end use. Combinations of hindered phenols and thiodipropionate synergists are effective antioxidants. Combinations of benzotriazoles and hindered amines are effective UV stabilizers.

Figure 1-107 Tape measurement grip

For products that do not have to be clear, titanium dioxide or carbon black pigments also provide effective UV protection. Many types of styrenic thermoplastic elastomers have been produced to meet speciﬁc end use requirements. General purpose styrenic block copolymer, soluble when compounded, improved stability. They are used for wire and cable coatings and medical applications. The hard and rigid compounds are used for injection molding and extrusion applications. Like most conventional vulcanized rubbers and unlike most thermoplastics, the styrenic thermoplastic elastomers have no commercial applications when the product is just a pure polymer. Depending on the particular requirements for each end use, they are compounded with other polymers, oils, ﬁllers, and

Figure 1-108 Industrial boot skirt

62

1 Polymeric Materials additives. In almost all cases, the products contain less than 50% of the block copolymer. Advantages • Good mechanical properties and high elongation at room temperature • Low cost resins

Figure 1-109 Diagnostic tester housing

• Thermoplastic processing equipment can be used (injection molding and extrusion) • Reground material can be recycled • Excellent impact modiﬁer compounding ingredient Disadvantages and Limitation • Poor compression set properties • Poor solvent resistance • Poor high temperature resistance • Usage only for noncritical applications

Figure 1-110 Fish knife handle

Typical Applications • Replacements for vulcanized rubber • Injection molded articles requiring modest end use temperatures for the footwear, toys, and furniture industries • Extrusion of sheets, tubing, and wire coating • Adhesives, sealants, and coatings • Polymer blends • Impact modiﬁer ingredient for other compounds • Viscosity index enhancer for lubricating oils • Modiﬁers for thermoset materials

1.3.4

Polyoleﬁn Thermoplastic Elastomer (TPO)

TPOs resemble other TPE resins; they have properties of a thermoset rubber, but are processible in standard thermoplastic equipment. TPO products cover a range of properties, bridging the gap between rubber and plastics. The ability to use thermoplastic processing equipment allows for greater productivity and economy. The recycling of reground materials, such as runners, sprues, and rejected parts, provides a signiﬁcant economic advantage. TPO materials are deﬁned as compounds of various polyoleﬁns, semi-crystalline plastics, and amorphous elastomers. The most common types of TPOs are composed of polypropylene (PP) and ethylene-propylene rubber (EPR). EPR may be either a copolymer of only ethylene and propylene monomers or a third monomer, the diene monomer, which provides a small amount of unsaturation in the polymer chain for sulfur crosslinking. This is called ethylene-propylenediene-monomer rubber (EPDM). Other polyoleﬁn polymers commonly used in TPO compounds include: low density polyethylene (LDPE), high density

63

1.3 Thermoplastic Elastomers (TPE)

General Properties of Generic TPO Speciﬁc gravity

0.88–0.98

Flexural modulus @ 73 °F (Kpsi)

2.7–300.0

Tensile strength @ yield (Kpsi)

0.95–4.00

Izod impact @ 73 °F (ft-lb/in)

1.00 No break

Elongation at break @ 73 °F (%)

20.0–600

Shore durometer hardness

40 D – 70 D

Thermal limits service temp. (°F)

180–212 (Stiff grade)

Shrinkage (%)

0.7–2.1

Melt point (°F)

290–330

Process temp. (°F)

360–500

Mold temp. (°F)

50–150

Drying temp. (°F)

150–225

Drying time (h)

3.0–6.0

polyethylene (HDPE), linear low density polyethylene (LLDPE), copolymers of ethylene with vinyl acetate (EVA), ethylacrylate (EEA), methyl acrylate (EMA), semi-crystalline copolymers of propylene, ethylene, and polybutene. TPO products are compounds of polyoleﬁn polymers. Like most thermoplastic elastomers, they are composed of hard and soft domains. The exact size and shape of these domains determines the properties of the compounds. The properties of the product may be determined as much by the process of producing the compound as by the composition. An unlimited number of formulations for TPO compounds are possible because of the wide variety of polyoleﬁn polymers used. Each of these compounds has its set of properties that may be useful in speciﬁc application. In most TPO compounds, the hard domain is isotactic propylene homopolymer or an isotactic propylene copolymer with a minor amount of ethylene as the comonomer. The ethylene monomer may be distributed either randomly in the copolymer or as blocks. Some of the segments of the polymer chain are composed of ethylene and propylene copolymers, while the other segments of the polymer chain are almost totally propylene homopolymer. The homopolymer and block copolymers have a crystalline melting point of 311–329 °F where as the random copolymers have a melting point of 293–311 °F. The relatively high melting point of the polypropylene portion of the TPO results in products that retain many of their mechanical properties at temperatures approaching the melting point of the polypropylene resin. Higher impact strength at low temperature is achieved by increasing the amount of ethylene present in the copolymer. Ethylene also reduces the rigidity of the copolymer. The crystalline melting point is depressed more sharply with increasing ethylene content in random than in block copolymers. The soft domain of the polymer chain is ethylene-propylene rubber (EPR) or ethylene-propylene-dienemonomer rubber (EPDM). Rubber materials with nearly equal amounts of ethylene and propylene are totally amorphous. The softest EPDM rubber grades are the most efﬁcient impact modiﬁer additives used in TPOs.

64

1 Polymeric Materials Some of the other additives used in TPO compounds function by modifying the rubber phase. Hydrocarbon oil increases the softness and ﬂexibility of a TPO compound by swelling the rubber phase. Polyethylene also increases the volume of the rubber phase, but does not have the softening effect of oils. High density polyethylene improves the impact strength at low temperatures while maintaining stiffness. In addition, a wide variety of other ingredients are used in compounding TPOs. These ingredients may include ﬁllers, reinforcing agents, lubricants, plasticizers, heat stabilizers, antioxidants, UV stabilizers, colorants, ﬂame retardants, foaming agents, ﬂow modiﬁers, or processing aids. Figure 1-111 Roller skate wheels

Figure 1-112 Air conditioner system

Figure 1-113 Tailgate bumper

TPOs can be formulated to combine strength and toughness. TPO products are available from 40–70 Shore “D” hardness, with a ﬂexural modulus ranging from 2,700–300,000 psi. Several factors determine the upper limit of use temperature for a TPO product. The melting point of the hard domain polymer is most important for short-term exposures. For most TPOs, the melting point of polypropylene is a limiting factor. Polypropylene homopolymers melt from 320–350 °F. Most TPO (hard/stiff) compounds will retain useful properties intermittently at temperatures up to 280 °F. For long-term exposure, the resistance of the TPO to aging effects at elevated use temperature is as important as the melting point of the hard domain polymers. The oxidative stability of TPOs is a function of the antioxidant and stabilizer additives. The most effectively stabilized TPO (hard/stiff) compounds are formulated with heat stabilizers, antioxidants, and reinforcements to withstand continuous use temperatures up to 250 °F or more. General purpose TPO (hard/stiff) compounds will retain their physical properties when used for extended periods from 180–212 °F. One of the outstanding properties of TPOs is their performance at low temperatures. TPO compounds retain their ﬂexibility at very low temperatures. The softer TPO grades often have a low temperature brittle breaking point of less than –112 °F. Almost all TPO products retain their physical properties when exposed to sunlight and weather. TPO compounds are made without unsaturated polymers in the polymer backbone and these types of materials are not susceptible to degradation by ozone. EPDM is a weather resistant rubber. TPOs also are not particularly susceptible to fungus growth. Many grades have outstanding color retention. These specially stabilized grades, designed for unpainted automotive applications, have been exposed in Florida sun aging tests for more than two years without changing color. All TPO products are unaffected by water and exhibit fair chemical resistance to acids and bases. Hot hydrocarbon solvents tend to soften and swell TPO products. This softening and swelling is typically slight for the harder formulations but severe for the softer products. The chemically inactive surface of a TPO part makes bonding to other materials difﬁcult. Most TPO compounds are good electrical insulating materials. They have good dielectric strength properties and they do not absorb moisture, because they are not hygroscopic. Many TPO parts that are used in automotive applications must be painted with an automotive ﬁnish to match or accent the painted ﬁnish of the other body panels. Paint systems are available that will give excellent adhesion of the paint to the TPO surface. The TPO materials do not have a molded surface that chemically reacts with the primer paint system to develop a strong and durable bond. To achieve adequate adhesion of the primer, the TPO must be modiﬁed in some way to obtain a reactive surface.

65

1.3 Thermoplastic Elastomers (TPE) Advantages • Excellent electrical insulating characteristics • Can be compounded with nylon to increase end use temperatures • Excellent impact resistance • Excellent fatigue resistance • Excellent low temperature brittle point for softer grades

Figure 1-114 Bumper fascia

• Excellent moisture absorption resistance • Not susceptible to degradation by ozone • Not susceptible to fungus growth • UV resistance grades are available Disadvantages and Limitations • Fair chemical resistance to acids and bases

Figure 1-115 Instrument panel shell

• Hot hydrocarbon solvents tend to soften and swell the TPOs. This softening and swelling is slight for the harder formulations but severe for the softer products • Bonding to other materials is difﬁcult • Reactive surface treating and a paint primer are required before painting Typical Applications • Exterior Automotive Applications: Body side trimming, body side molding, bumper covers, bumper end caps, bumper side pieces, fender liners, sight shields, stone deﬂectors, wheel well moldings, and valance panels • Under Hood Automotive Applications: Wiring harness protective sleeve, blow molded air duct, sound absorbing ﬁre wall blanket, automotive high performance audio systems, such as speaker box enclosures, rocker panel covers, rub strips, and scuff plates • Wire and Cable: TPOs are used for a number of low voltage wire and cable applications. They are used for insulation and jacketing for battery booster cables, portable power supply cords, and submersible pump cable • Mechanical Goods: Seals, electrical plugs, extruded sheet, and weather strips • Impact Modiﬁer: Polyoleﬁn and polypropylene are used to improve the low temperature impact strength; TPO also is added to HDPE to improve stress crack resistance

1.3.5

Elastomeric Alloy Thermoplastic Vulcanized (TPV)

Vulcanized elastomeric alloys are thermoplastic elastomers composed of mixtures of two or more polymers that have received a proprietary treatment. TPV compounds tailor their characteristics during the manufacturing process. The improved TPV properties are superior to other major competitive compounds.

Figure 1-116 Bumper fascia

66

1 Polymeric Materials

General Properties of Generic TPV Speciﬁc gravity

0.94–1.34

Flexural modulus @ 73 °F (Kpsi)

2.7–50.3

Tensile strength @ yield (Kpsi)

0.64–4.0

Compression set @ 22 h (%)

16–44 @ 73 °F

Elongation at break @ 73 °F (%)

330–600

Shore durometer hardness

55 A – 87 A 40 D – 50 D

Thermal limits service temp. (°F)

–75 275

Shrinkage (%)

1.5–2.5

Vicat point (°F)

206–290

Process temp. (°F)

360–410

Mold temp. (°F)

50–175

Drying temp. (°F)

150–175

Drying time (h)

2.0–4.0

EA-TPVs are a category of TPEs made of a rubber and plastic polymer mixture in which the rubber phase is highly vulcanized. The plastic phase of a TPV is a polyoleﬁn (polypropylene) and the rubber phase is an ethylene-propylene elastomer. The ﬁrst TPV (1981) was an EA-TPV, in which the plastic phase was polypropylene and the rubber phase was EPDM rubber. A subsequent TPV of polypropylene and nitrile rubber was introduced to the plastic industry. The vulcanization of the rubber phase of an EA-TPV results in numerous property improvements. Insoluble in rubber solvents, the TPV exhibits less swelling in these solvents than TPE rubber and plastic blends. This vulcanization offers several improvements in the properties of EA-TPV: • Increase in tensile strength and modulus • Decrease in compression set • Decrease in swelling caused by oils Vulcanization of the rubber phase also improves the retention of properties at elevated temperatures. The dynamic vulcanization embraces the curing of a rubber composition during its compounding and one of the ingredients of this rubber composition must be a thermoplastic resin. It is important that the mixing be continuous during the compounding step. The temperature reached during the mixing must be sufﬁciently high to melt the resin and effect the chemistry of the crosslinking reaction. Elastomeric alloy TPVs differ from simple elastomer blend TPEs in the degree of vulcanization of the rubber phase. The rubber TPV consists primarily of a ﬁne dispersion of highly vulcanized EPDM rubber in a matrix of polypropylene as the continuous phase. Of equal importance for the improved properties of a TPV is the size of the particles of vulcanized EPDM. As the particle size decreases, the properties progressively improve. Typical grades range from 55–87 Shore “A” and from 40–50 Shore “D”. Service temperature ranges from a minimum of –75 °F to a maximum of 275 °F.

1.3 Thermoplastic Elastomers (TPE) The tear strength and abrasion resistance of elastomeric alloy TPVs is good but not outstanding, the tear strength increases progressively with hardness. The desirability of low or high resilience (hysteresis) will depend on the end use. Uses, where heat generation of the ﬂexing rubber is critical, demand high resiliency; those requiring vibrational damping need low resiliency.An outstanding dynamic property of thermoplastic vulcanizates is their fatigue resistance. TPVs have a maximum use temperature in air of 275 °F for extended periods of time. This temperature is based on air aging data after 1,000 hours, 50% retention of elongation, 70% retention of tensile and hardness reduction. Accelerated outdoor aging tests on black TPVs show that they retain their performance properties for 12 months; colorable EPDM-TPVs are much less resistant to the combined effects of solar radiation, atmospheric oxygen, ozone, and other atmospheric pollutants. The resistance of EPDM-TPVs to polar ﬂuid media is excellent. The environment ranges from a corrosive media, such as concentrated acids and alkalis, to mild media, such as aqueous salts and polar organic and inorganic ﬂuids. As the polarity of the medium decreases, so does the resistance of the EPDM-TPVs, culminating in their poor resistance to halocarbons. Chemical resistance to hydrocarbons is fair; the resistance to saturated hydrocarbons is better; resistance to water, aqueous solutions (both acids and alkalis), and hot oils are poor. These materials cannot be recommended for service in diesel fuel above 158 °F. TPVs have fair to good resistance to industrial and automotive ﬂuids at elevated temperatures. Thermoplastic vulcanizates have excellent electrical insulation properties, an excellent dielectric strength 400 volts/mil at a thickness of 0.080 in. The volume and surface resistivity of EPDM-TPVs are sufﬁciently high to justify their consideration for use as primary electrical insulators, as well as jacketing material. The dielectric constant (speciﬁc inductive capacitance) and power factors are basic limits in the selection of a primary electrical insulation material. These two parameters of ﬂame retarded TPVs at ambient and elevated (212 °F) temperatures and the normal power transmission frequency of 60 Hz make them a good choice for electrical applications. TPVs absorb moisture, causing signiﬁcant processing problems, such as internal voids, rough surface appearance, and poor dimensional control. When using a hot runnerless mold, moisture causes a dangerous buildup of high pressure steam. Therefore, moisture adsorption after exposure for 24 hours, or even less, causes signiﬁcant processing problems. The recommended moisture content limit for these materials is less than 0.3% before processing. Proper drying will eliminate most of the major processing problems for these materials.

Figure 1-117 Soft-touch common handling tools

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68

1 Polymeric Materials It is strongly recommended that TPVs be dried again immediately before use. Desiccant drying systems are recommended for drying the resin from 2–4 hours at 150–175 °F. Hot air drying systems are not recommended, because they remove only the moisture from the surface of the resin, the interior volume of the resin does not become dried, producing uneven drying results. Advantages • Can be injection molded, extruded, compression molded, blow molded, calendered, thermoformed, and extrusion foamed Figure 1-118 Weatherstripping seal

• Good tear strength and abrasion resistance • Excellent fatigue resistance • Excellent resistance to polar ﬂuid media • Excellent electrical insulation properties Disadvantages and Limitations • Poor UV resistance • Poor bonding, surface preparation is recommended • Chemical resistance to hydrocarbons is fair; the resistance to saturated hydrocarbons is better • Poor resistance to halocarbons

Figure 1-119 Equipment grip handle

• Poor chemical resistance to aqueous solutions (both acids and alkalis) and hot oils • TPVs cannot be recommended for service in diesel fuel above 158 °F • Fair resistance to industrial and automotive ﬂuids at elevated temperatures Typical Applications • Automotive: Hose coverings, air inlet duct covers, gaskets, seals, convoluted boots, vibration dampeners, ignition components, window seals, cover systems connecting the distributor to the spark plugs and vibration dampeners

Figure 1-120 Automotive headlamp seal

• Architectural and Construction: Expansion joints, rooﬁng, ﬂooring, and weather seals around doors and windows, soft extruded TPV proﬁle as a window glazing material, metal reinforced TPV weather stripping, and other exterior openings. • Electrical and Electronics: Insulator, jacketing material, wire and cable coverings, electrical connectors, plugs, insulators for electrical and electronic assemblies, computer hardware and software, telephones, electronic appliances, duplicating machines, and ofﬁce equipment • Hose, Tubing, and Sheeting: Sheeting is used for on-site fabrication of rooﬁng membranes, seals, and gaskets. It is also used for different types of hose and tubing, inner liners, jacketing, and hoses reinforced with a textile ﬁber or metal mesh

Figure 1-121 Automotive air duct

• Medical and Food Contact: Applications in direct contact with foods, beverages, pharmaceuticals, and living tissues. These applications employ special grades that comply with the regulations of the USA Food and Drug Administration. Medical devices include syringe plunger tips, drug vial

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1.3 Thermoplastic Elastomers (TPE) stoppers, aerosol valve seals, medical tubing, nursing bed sheets, and liquid dispenser pump diaphragms. All of these applications are derived from the biocompatibility and low toxicity of elastomeric alloy TPVs, both before and after sterilization • Mechanical Rubber Goods: Household appliances, industrial equipment, recreational devices, seals, gaskets, valve seats, grommets, appliance feet, bushings, toys, hand tools, ski pole grips, caster wheels used in grocery carts, toys, and small industrial equipment

1.3.6

Melt Processible Rubber (MPR)

General Properties of Generic MPR Speciﬁc gravity

1.10–1.26

Flexural modulus @ 73 °F (Kpsi)

1.10–1.40

Tensile strength @ yield (Kpsi)

1.03–1.68

Compression set @ 22 h (%)

12–18 @ 73 °F

Elongation at break @ 73 °F (%)

265–350

Shore durometer hardness

60 A – 80 A

Thermal limits service temp. (°F)

–40 to 212

Shrinkage (%)

1.5–3.0

Brittleness temp. (°F)

–132 to –105

Process temp. (°F)

300–380

Mold temp. (°F)

250–300

Drying temp. (°F)

150–200

Drying time (h)

1.0–2.0

With the introduction of the melt processible rubber (MPR) in 1985, it was possible to eliminate vulcanization, increase productivity, recycle the scrap, and minimize the dependency on the rubber compounds. MPR has the range of properties and cost savings needed to replace conventional thermoset rubber. This compound is not a thermoset material. The MPR resin produces parts that look, feel, and perform like vulcanized rubber and can be processed on equipment used for thermoplastics or thermosets. It also offers the combination of resistance to heat, oil, chemicals, and weathering. The melt processible rubber resin has been recognized as the most rubbery-like product among all the TPE families. MPRs are available in a hardness range from 60–80 Shore “A” durometers. MPR alloys are composed of proprietary ethylene interpolymers and chlorinated polyoleﬁns, in which the ethylene interpolymer component has been partially crosslinked. Plasticizers, stabilizers, antioxidants, curing agents, and ﬁllers can be incorporated into the resin to provide ﬂexibility and reinforcement. The MPR polymer network is amorphous, providing a very low ﬂexural modulus, excellent tensile strength, and linearly proportional stress-strain curve. Halogenated polyoleﬁns are intermixed with a large number of structurally coupled ethylene interpolymers. The ethylene interpolymers are used as intermediates in MPRs. They contain functional groups that are strongly protonaccepting to promote hydrogen bonding with the alpha hydrogen atoms of the halogenated polyoleﬁn over the entire composition range. The MPR polymer

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1 Polymeric Materials blends are single-phase systems, because they have a single glass transition temperature. As an amorphous material, the MPR exhibits no crystalline melting point. It gradually softens with increasing temperature, but does not ﬂow at any temperature, unless also subjected to shear. By increasing the shear rate of the MPR, the melt viscosity property of the polymer increases proportionally. The polymer melt only becomes softer, but not a liquid melt. This unique polymer characteristic makes it possible for the melt to ﬂow through small-size gates, runners, and parts with thin wall thicknesses. Increasing the temperature in the range from 300 °F–380 °F has little effect on viscosity. Little is gained in processing MPR at temperatures above 350 °F. Because of its rheology and its amorphous single-phase nature, MPR is obviously not a plastic. It has a fractional melt index, does not melt and exhibits thixotropic behavior. It has a useful plastic processing property in ﬂowing like a thermoplastic on application of shear. Unlike plastics, it displays minimum draw-down because of elastic recovery, even at elevated temperatures. MPR provides a balance of properties, such as durability and performance sufﬁcient to replace thermoset rubbers. MPR is unique in duplicating the feel and recovery of vulcanized rubber and in being processible on both thermoplastic and modiﬁed rubber equipment. It has outstanding oil resistance, heat aging, weatherability, and is recyclable. It has mechanical properties, tear and abrasion resistance, and a continuous temperature similar to mid-performance rubbers. MPR is used as an energy absorbing polymer, but it has limitations in applications requiring very dynamic motion and load carrying capabilities. Advantages • The most rubbery-like product of the TPEs • Blow molding, injection and compression molding, extrusion, calendering processes are possible • Excellent resistance to ultraviolet light • Low operating temperatures below –40 °F • Excellent resistance to parafﬁnic (ASTM oil #1) and aromatic (ASTM oil #3) oils, to gasohol (15% methanol), unleaded gasoline, diesel fuel, JP-4 jet fuel, and kerosene • Good resistance to aromatic fuel B, fuel C, and fuel A • Inert to solvents such as alcohols and ethers, and hydrocarbons, such as hexane, cyclohexane, turpentine, and ﬂuorocarbon (Freon® 112) Figure 1-122 Wrench handle

• Excellent resistance to many lubricants, steering, and brake ﬂuids and other hydraulic ﬂuids Disadvantages and Limitations • Poor chemical resistance to polar ﬂuids, including esters, aldehydes, ketones, and chlorinated solvents • Low swelling in boiling water, glycol, and water coolants, aqueous alcohols, acids, bases, and salts • Poor electrical properties

Figure 1-123 Gasoline nozzle pump

• Continuous operating end use temperature of 212 °F

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1.3 Thermoplastic Elastomers (TPE) Typical Applications • Reinforced Hoses: Low performance, high volume, such as garden hoses and automotive heater hoses • Automotive Hoses: Cover for air condition hoses and fuel lines, ﬁller neck hoses, vent hoses, evaporative loss hoses, truck air brake hoses • Industrial Hoses: Hydraulic wire braid, hydraulic textile braid, oil suction discharge, chemical and oil transfer, agricultural spray, ﬁre extinguisher, fuel delivery, shop vacuum hose, washing machine, and dishwasher hoses • Service Cords: Appliances, electric drills, power drills, saws, typewriters, and ofﬁce machines • Control Cable: Low voltage cables for control switches and relays • Flexible Power Cable Jacket: Airport lighting, industrial plants, oil well, utility power, urban primary, and secondary network cables.

Figure 1-124 Bumper/lamp gasket

• Elastomeric Sheeting: Colorable, ﬂexible sheeting for single ply rooﬁng, elastomeric lining of ponds and pits for industrial waste • Molded Goods: Molded MPR articles of all shapes and forms, that require the rubbery feeling, wear, UV resistance, color and molded to metal adhesion properties • Fabric Coating: Industrial pond and pit liner, automotive fabrics, and different types of clothing • Seals, Gaskets, and Weather Stripping: Building gaskets, highway and bridge gaskets, appliance seals, automotive weather stripping, extruded proﬁles of varying complexity, circumferential gaskets (compression molded or die-cut from sheet stock), coextrusion composite using a MPR seal and a rigid vinyl (PVC) or other hard material as reinforcement

1.3.7

Copolyester Thermoplastic Elastomer

General Properties of Generic Block Copolyester TPE Speciﬁc gravity

1.07–1.43

Flexural modulus @ 73 °F (Kpsi)

4.7–175.0

Tensile strength @ yield (Kpsi)

1.49–7.00

Brittleness temp. (°F)

–157 to –67

Elongation at break @ 73 °F (%)

200–700

Shore durometer hardness

35 D (soft) 82 D (hard)

Thermal limits service temp. (°F)

–40 (soft) 300 (hard)

Shrinkage (%)

0.50–1.60

Tm (°F)

312–433

Process temp. (°F)

360–500

Mold temp. (°F)

90–150

Drying temp. (°F)

212

Drying time (h)

2.0–3.0

Figure 1-125 Automobile steering ring

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1 Polymeric Materials Block copolyester thermoplastic elastomers have a combination of mechanical properties, such as tensile strength, elasticity, creep resistance, and dynamic properties. These polymers are operational over a broad service temperature range without signiﬁcant change in properties. They also have excellent chemical, heat, and oil resistance properties. In 1972, E. I. DuPont Company introduced a new material based on 1,4-butanediol, terephthalic acid and polytetramethylene ether glycol or polypropylene glycol. Because the building blocks behaved like high tensile strength engineering plastics and elastomers, respectively, the product combined these characteristics in proportion to the ratio of hard and soft molecular segments. The fast crystallization rate of the hard segment, combined with exceptional melt stability of the backbone, allowed these polymers to be processed like a thermoplastic material. Block copolyester elastomers stress-strain curves show why these polymers are not like thermoplastics or thermoset rubbers. They are a class of materials combining some of the strength of plastics with some of the elasticity of rubber. A distinguishing characteristic of block copolyester elastomer resins versus other ﬂexible materials is their excellent dynamic performance, which makes these resins suitable for applications requiring long-term spring properties and ﬂex life. Operating within their elastic limits, block copolyester elastomers are creep resistant, supporting loads for a long time without stress relaxation. They can be subjected to repeated cycles of tension and compression without signiﬁcant loss of mechanical strength. Block copolyester elastomers demonstrate low hysteresis loss in dynamic applications. Product components working at low strain levels exhibit complete recovery in cyclic applications with little heat buildup. Typical melt temperatures range from 360–500 °F. As a general rule, stiffer grades have higher melting points and crystallization temperatures, yielding shorter mold cycles than those of the more ﬂexible grades. The melting points between soft and hard grades, for example, 312 °F versus 433 °F, introduce some variations in the process conditions. The process thermal stability of copolyester elastomers permits exposure to elevated temperatures and long residence time without degrading the polymer. The different resins have a signiﬁcant difference in melt rheology. Their viscosity is tailored for speciﬁc processes; the low viscosity melts are used for injection molding and extrusion, while the high viscosity melts are used for extrusion blow molding. Resistance to cut growth during ﬂexing is outstanding, mainly because of high resilience and low heat buildup. Block copolyester elastomers rank high among thermoplastics for impact resistance; most of the compounded grades do not break by conventional notched Izod impact testing equipment. Resistance to abrasion is a complex function of tear strength, coefﬁcient of friction, resilience, heat dissipation, and other properties. If high mechanical strength is required in an abrasive environment, copolyester elastomers will outperform any type of thermoplastic elastomer and rubber. Copolyester elastomers are used in electrical applications with less than 600 V. The following characteristics make them attractive for electrical and electronic applications: good dielectric properties, high mechanical strength, creep resist-

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1.3 Thermoplastic Elastomers (TPE) ance, spring properties, ﬂex fatigue resistance, high impact strength, abrasion resistance, and end use service temperature range between –40 to +300 °F. Block copolyester elastomer compounds have excellent oil resistance; in several applications they have replaced materials such as polyacrylate, silicone, ﬂuorosilicone, SBR, nitrile, and polychloroprene rubber. Block copolyester elastomer stiffer grades offer the best performance in hot hydrocarbon environments, with many resins suitable for use with hot oil, grease, fuels, and hydraulic ﬂuids. Block copolyester elastomer resins are attacked by polar ﬂuids at temperatures above 158 °F, even though the resins are resistant to fuels, they are not suitable for under the hood automotive applications. Hydrolytic stabilizing concentrates are available from Du Pont; this additive is blended with the virgin resin to extend service life in the presence of hot water. Many applications, such as containers, sheets, and seals, require low permeability to standard fuels. Block copolyester elastomers are rather permeable to polar molecules such as water, but are resistant to penetration by nonpolar hydrocarbons and refrigerant gases. Permeability to water vapor makes polyester elastomers useful in coating fabrics for rain resistance. The low permeability to refrigerant gases and propane is important for refrigerant hose and propane gas hose, respectively. Block copolyester elastomers are highly resistant to radiation. Block copolyester elastomer compounds do not contain plasticizer additives in their formulations. The inherent chemical purity of block copolyester elastomers makes them an excellent choice for food contact and medical applications. Block copolyester elastomers are UL-94 HB rated. Master batch additives and compounded grades are available to meet UL-94 V0 ratings. Block copolyester elastomer parts can be produced by injection molding, injection blow molding, blow molding, extrusion, and rotational molding. Film processes include casting, extrusion, and blowing. Block copolyester elastomers are subject to hydrolysis (moisture absorption) at processing temperatures. It is essential that the polymer be dried before processing. The virgin resins are supplied in moisture-proof bags with less than 0.10% moisture. The virgin material can be used directly from sealed bags. However, all reground material and resin from an open shipping bag or container exposed for more than one hour to ambient air, must be dried before processing. The material is hygroscopic and within one hour can absorb enough moisture from ambient air to cause degradation during processing. It is recommended to dry the wet resin from 2–3 h in a dehumidifying oven or hopper dryer at 212 °F.

Figure 1-126 Headphone set spring holder

Reground ratios as high as 100% are possible, however, the reground ratio must take into account the degradation of the reground being blended with virgin material. As a general rule, the level of reground material should be less than 50% to maintain the highest polymer quality. Advantages • Hardness between 35–82 Shore “D” • Excellent dynamic performance • Excellent hysteresis ﬂex life

Figure 1-127 Printer belt drive

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1 Polymeric Materials • Low creep recovery from deformation • Excellent ﬂex fatigue and cut growth resistant • Excellent ﬂexibility over broad temperature range • Good abrasion resistance • Excellent impact resistance at low temperatures Figure 1-128 Pistol holster

• High tensile strength and tear strength • Good chemical and hot oil resistance • Excellent processing characteristics • Recyclable polymer Disadvantages and Limitations • Lower compression set resistance (high deﬂection) than rubber

Figure 1-129 Safety air bag cover

• Lower elastic recovery at high strains compared with rubber • Poor polar ﬂuid resistance at temperatures above 158 °F • Special compounds to optimize properties are difﬁcult to obtain • Addition of UV stabilizers required for outdoor applications • Drying reground or wet resin is required • Different molding conditions are required for each polymer type • High resin cost

Figure 1-130 Wall anchor fastener

Typical Applications Block copolyester elastomers are well suited for applications requiring strength, ﬂexibility, or dynamic properties. They replace rigid plastics and soft materials, such as rubber, where improved shock absorption, impact strength, ﬂexing, sealing, spring characteristics, or silent mechanical operation are required. • Replacement of Metals: Coextruded hinges of copolyester elastomer and PVC, one piece rotomolded wheels, gear replaced cast iron in a textile braiding machine

Figure 1-131 Automotive CVJ boot

• Replacement of Cast Urethane: Automotive body parts including spoilers, air dams, CVJ boots, air bag covers, coil spring connecting brake hoses for trailer trucks, side moldings, seals, hydraulic piston actuators, and grills, railroad and large construction machinery heavy duty shock absorber stops, ﬂuidﬁlled piston actuated shock absorbers for mining trucks, skiing equipment components, irrigation valve diaphragm, and coil phone cords • Replacement of Leather: Footwear, military pistol holsters, straps, soccer balls, motorcycle face masks, and sport clothing • Replacement of Rubber: Hose and tubing, compact garden hose, high pressure hydraulic hose, power drive belts for ofﬁce equipment and computers • Electrical Applications: Including wire insulation, retractile wire, switches, and sealed connectors

Figure 1-132 Automotive suspension

• Specialty Products: Food contact and medical applications

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1.3 Thermoplastic Elastomers (TPE)

1.3.8

Polyamide Thermoplastic Elastomer

General Properties of Generic Polyamide TPE at 50% RH Speciﬁc gravity

0.99–1.14

Flexural modulus @ 73 °F (Kpsi)

27.0–42

Tensile strength @ yield (Kpsi)

3.00–6.90

Shrinkage (%)

0.50–2.50

Elongation at break @ 73 °F (%)

270–715

Shore durometer hardness

60 A – 90 A 25D – 63 D

Thermal limits service temp. (°F)

338–392 (hard grades)

Vicat point (°F)

140–378

Tm (°F)

300–480

Process temp. (°F)

320–537

Mold temp. (°F)

70–140

Drying temp. (°F)

160–230

Drying time (h)

4.0–10.0

The thermoplastic polyamide elastomers consist of hard and soft segments joined by amide linkages. The soft segments of the material could have polyether or polyester in the backbones of the structure. Three different chemical compounds have been developed. The creation of this new family of thermoplastic polyamide elastomers is the result of the research work of three plastics companies. The ﬁrst formulation consists of an adipic acid capped hard segment block of poly (11-amino-undecanoic). The hard block is joined with a soft segment, such as polyoxyethylene glycol, polyoxypropylene glycol, or polyoxytetramethylene glycol, in a polyesteriﬁcation process to form polyetheresteramide (PEEA). The second formulation is a variation, in which there are no ester linkages. The bonds between the hard and soft segments are amides. In this formulation, an amine-terminated soft segment, bis(3-aminopropyl)-polyoxytetramethylene glycol, is reacted with a dimer acid and caprolactam to form the polyetheramide (PETA). The third formulation consists of an acid terminated soft segment that is formed ﬁrst by esteriﬁcation of a polyoxyalkylene or other glycol. This is reacted, along with additional diacid, with a diisocyanate to form the polyesteramide (PESA), in which the hard segments are the amides formed from the additional diacid and the diisocyanate. PESA is the only formulation that incorporates substantial aromatic structures into the backbone to modify its properties. Thermoplastic polyamide elastomers exhibit properties that depend on the chemical composition of the hard (polyamide) and soft (polyether, polyester, or polyetherester) blocks as well as their segment lengths. The composition and molecular weight of the polyamide hard segment determines the polymer melting point, processing temperatures, and mechanical properties at elevated temperatures. The soft segment inﬂuences the ﬂexibility at low temperatures and the chemical and solvent resistance. Polyether soft segments have better low temperature

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1 Polymeric Materials properties and hydrolytic stability, whereas polyester soft segments have better solvent and thermal resistance and oxidative stability. Thermoplastic polyamide elastomers hardnesses range from 60–90 Shore “A” and from 25–63 Shore “D”, at temperatures between –40 °F and 176 °F. These materials have good abrasion resistance, comparable to thermoplastic polyurethane (TPU). Thermoplastic polyamide elastomer’s tensile strength, ﬂexural modulus, and elongation properties are higher than those of any TPE material in the same hardness range. This is a consequence of the load bearing ability of the crystalline portion of the amide hard segment domains. All of these mechanical properties show an improvement after annealing above the Tg of the amide hard segment. The thermoplastic polyesteramide elastomer is a high temperature resistant elastomer that can withstand temperatures up to 338 °F continuously and 392 °F short-term exposure, while still retaining useful mechanical properties. The PESA elastomers have good compression set properties when measured under constant low load bearing conditions. Because of their high ﬂexural modulus and high load bearing capabilities, the high ﬂexural modulus test conditions generate high stress levels. However, when the hard segment domain becomes reorganized, it produces higher (poor) compression set values. The PESA elastomers are very resistant to long-term dry heat aging. The tensile strength properties improved after aging for 5 days at 302 °F (annealing). Similar studies were performed for the ester-based PESA elastomers at 347 °F with only small losses in properties. With the addition of heat stabilizers, polyetheresteramides exhibit improved heat resistance characteristics. The chemical composition of the soft segment is important in determining resistance to humid aging conditions. PESA elastomers are susceptible to hydrolysis attack (reduction of properties by lowering the polymer molecular weight). Appropriate use of hydrolysis stabilizers improves the humid aging characteristics of these polymers. PESA formulations become less susceptible to hydrolysis as their hardness increases, because the ester content is reduced. PETA and PEEA polymers show excellent retention of their original tensile strength properties in humid aging tests. Aromatic polyesteramide elastomers (PESA) exhibit good resistance to UV radiation under moisture conditions. PETA and PEEA polyamide elastomers have less resistance to ultraviolet radiation. However, most of the PEEA grades are available with UV stabilization to endure at least 2,100 hours of outdoor service. The polyamide elastomers’ hard segment greatly affects the chemical and solvent resistance of a segmented block elastomer, because the hard segment domains must maintain the integrity of the polymer. The semi-crystalline amide hard segment in polyamide elastomer (PESA) has low solubility in many solvents. Polyesteramide (PESA) has good resistance to oil, fuel, grease, and phosphatebased hydraulic ﬂuids, but it has poor resistance to chlorinated solvents. Polyamide elastomers must be dried to 0.10% (minimum) or 0.02% (optimum) moisture content (depending on the type of resin) at temperatures from 160– 230 °F with the dew point between –22 to –40 °F for 4–6 hours for virgin resin straight from the bag and for 4–10 hours if the resin is wet, using a dehumidifying hopper dryer system before processing the resin.

1.4 Liquid Injection Molding Silicone (LIM®) It is possible to recycle sprues, runners, and rejected parts of polyamide elastomer resin after regrinding and drying the material. Blending from 10–15% of reground material with virgin resin is recommended to avoid any appreciable change in the characteristics of the molded parts. However, it is not advisable to recycle any reground material in parts that have a very critical functional requirement. Advantages • Hardness range from 60 Shore “A” to 63 Shore “D” • Temperature resistance 338 °F (continuous) and 392 °F (short-term) • Good thermal aging properties • Good chemical and solvent resistance to oil, fuel, and grease • Good impact strength, ﬂexibility at low temperatures • Good tear strength and abrasion resistance • Easy processing by injection molding • It is possible to recycle from 10–15% of reground material. Disadvantages and Limitations • Poor compression set properties • Susceptible to hydrolysis attacks and ultraviolet radiation • Poor chemical resistance to chlorinated solvents • The resin must be dried before processing Typical Applications The end use applications of this new family of thermoplastic elastomers are just beginning to emerge in the following markets: • Hoses, tubing, bellows, and boot enclosures • Wire and cable jacketing, where mechanical toughness, oil, grease, solvent, and chemical resistance are required • Low pressure seals and gaskets • Recreational ski boot components, footballs, and basketballs • Shoe sole applications

1.4

Liquid Injection Molding Silicone (LIM®)

The chemical structures of silicones are different from other polymers. Organic hydrocarbon polymers are based on a backbone of carbon-to-carbon atoms. The silicones have a backbone of silicon-to-oxygen linkages. These linkages are much stronger than those in the organic hydrocarbon polymers; they resemble the Si–O linkages found in inorganic materials, such as quartz, glass, or sand. Silicones are more resistant to attacks by high temperatures, degradation, and oxidation type processes.

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78

1 Polymeric Materials

General Properties of Generic LIM Silicone Rubber Speciﬁc gravity

1.12

Tensile modulus @ 73 °F (psi)

36.0–230.0

Tensile strength (Kpsi)

1.15–1.25

Tear strength (lb/in)

160–250

Elongation at break @ 73 °F (%)

400–750

Compression set 22h @ 350 °F (%)

20–40

Thermal limits service temp. (°F)

–75 to 400

Shore durometer hardness

30 A – 70 A

Dielectric strength (V/Mil)

470–530

Dielectric constant @ 1,000 Hz

3.08–3.17

Shrinkage (%)

3.00

Process temp. (°F)

302–400

Mold temp. (°F)

300–400

Silicones perform well at low temperatures and remain stable at high temperatures; they can be sticky or slippery materials; they can protect against electrical shock; or they can conduct electrical current. Liquid injection molding compounds are a family of specially formulated products used to produce precision elastomeric parts efﬁciently and economically, using specially modiﬁed injection molding equipment. LIM® silicones differ from conventional thermoplastic polymers in the following manner: their viscosity is low, i.e., they have high ﬂow rate properties; the material has the tendency to generate ﬂashing problems during the injection molding process. To avoid the ﬂashing problems, the mold tolerances must be tighter and higher clamping forces are required for the mold. The curing chemistry characteristic makes LIM® silicone a thermoset rubber polymer. For molding, the mold must be heated between 300–400 °F. LIM® cures quickly once the temperature has been reached. Therefore, the hardness of the mold cavities must be high, the molds must operate at elevated temperatures, and the molds must be vented to ensure even ﬁlling of the cavities. LIM® silicones are elastomeric materials that do not shrink when the molded parts are still in the mold cavities. The molded parts’ shrinkage (size reduction) takes place after the parts have been ejected from the mold, when the hot parts start to cool down at room temperature. The molded parts should be ejected from the lower or moving half versus the upper or ﬁxed half of the mold. The mold ejection techniques used for processing rigid thermoplastic materials are not recommended for ejection of LIM® silicone parts from the mold. These techniques could damage the LIM® silicone surfaces that are in contact with the mold ejection mechanism system. Most LIM® silicone rubbers show shear thinning effects. The melt viscosity properties for shear rate values are between 500 and 10,000 s–1 and are critical for maintaining production of injection molded parts. A normal injection molding machine and mold are capable of generating shear rates of this magnitude. The molding operator can adjust the process conditions to mold parts efﬁciently.

1.4 Liquid Injection Molding Silicone (LIM®) The melt temperature affects the viscosity and the melt ﬂow rate. The molder has to inject LIM® (A/B mixture) quickly and completely in the heated mold cavity before the melt temperature accelerates the cure or cross linking process. Higher injection speed and shorter injection forward time will increase productivity. The injection forward time (less than 5 s) should be faster than the mold curing time. Mold temperature should be high enough to reduce the cycle time.

1.4.1

LIM® Silicone Processing

Liquid injection molding was developed from room temperature vulcanizing technology. LIM® is a family of specialty products that use two liquid formulations in a 1 : 1 ratio. These liquid injection molding components are pumped from drums into meter mixing equipment and are automatically colored, catalyzed, and then transferred to the mold without the need of post curing. These compounds are used to produce precision elastomeric parts efﬁciently, at a lower cost, with high volume production. However, a special modiﬁed injection molding machine, a precision high-temperature mold, and a feeding, mixing, and injection system are required. Liquid Components Feeding System • Feeder Pump: Used for feeding and metering discharge pump, the pail pump should have follower plates, driven by either hydraulic or pneumatic systems. • Metering Discharge Pump: Used for metering the constant mixing ratio of the two components by volume, this pump should be a piston or screw hydraulic type. • The pumps should be made of stainless steel, and the internal components should be protected with Teﬂon®; the packing materials should be made from Teﬂon® and Viton® (ﬂuorocarbon rubber); the hoses should be made from Teﬂon®, nylon, or polyester. Other components should be made from stainless steel with a Teﬂon® lining. • Mixer: The mixer is used for blending the matrix with several additive materials. The mixer could be either a static type or a hydraulically driven plasticizing screw used as a mixer. • Injector: The injector supplies pressure and temperature to inject the compounded mixture in the back of the injection molding barrel. The molding machine could be either a hydraulically driven plunger type or a reciprocating plastifying screw. Injection Molding Machine Modiﬁcations • Heater bands on the injection barrel and nozzle are replaced by water cooled jackets • Use corrosion resistant steel needed for the barrel, screw, and check valve • Extension nozzles are recommended to reduce the sprue length • Mold is heated to 300–400 °F with oil or electric heaters • Injection pressures required areranging from 800–3,000 psi • High clamping pressure must be used to prevent ﬂashing

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1 Polymeric Materials • The barrel must be modiﬁed to seal the injectors and water coolant • A valved shut-off nozzle is required to prevent drooling LIM® Mold Requirements The molds are the most important component for injection molding LIM® silicone rubber parts. Processing demands on the molds are severe because of LIM’s low viscosities or high ﬂow rates and high cure temperatures. LIM® molds are generally operated under higher temperatures and higher injection pressures compared to compression and transfer molding processes. At 400 °F, most steels are highly susceptible to galling. It is very important that the steels for LIM® molds be corrosion resistant and hardened. Hardening the steel of the molds allows the use of thinner steel mold plates. Steel hardening is one of the techniques to maintain close tolerances at elevated temperatures and increase the useful life of the mold by reducing steel wear. LIM® Mold Design Recommendations • Chrome plated tool steel, stainless steel, and hardened aluminum • The surface treatments for the LIM® molds are: nickel-TFE coating, chrome plating, and titanium nitride coating.

Figure 1-133 Soft touch encapsulated camera (Courtesy: GE)

• The cold runner cross section should be fully round, its size between 0.062 and 0.125 in diameter. • The type of gates should be pin-point or tapered edge gates. A large full base diameter between 0.031 and 0.087 in, the gate wall surface slightly tapered towards the cavity, and the tapered gate length between 0.050 and 0.100 in. • Mold vents are needed to remove the entrapped air from the cavities and runners. The vents also remove the polymer gas produced inside the plastifying unit. The depth of the vent (cavity and runner cold slot pocket) should range from 0.001–0.002 in, the width between 0.050 and 0.200 in, the land between 0.040 and 0.062 in. The depth of the vent scape channel should range from 0.062–0.093 in until it reaches the edge of the mold. The vents should be located on the upper or moving half of the mold.

Figure 1-134 Face mask seal (Courtesy: GE)

• The temperature of the lower or ﬁxed half of the mold should be 10 °F–30 °F lower than the temperature of the upper or moving half, making sure that the molded parts will remain on the eject side of the mold. • LIM® molds must be thoroughly cleaned and degreased before installation. Lubricants should be applied only to guide pins and bushings and not on the surfaces contacting the mold cavities. LIM® molds must be maintained and stored in a clean environment. LIM® molded parts will stick to a trace of oil or grease in the mold cavities at normal mold temperatures between 300–400 °F. Any type of dirt speck will cause galling of the mold cavity surfaces. Advantages • Excellent thermal stability • Excellent weather resistance • Good electrical properties

Figure 1-135 Various injection molded parts (Courtesy: GE)

• Excellent chemical resistance • High temperature resistance

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1.4 Liquid Injection Molding Silicone (LIM®) • Low temperature impact resistance and ﬂexibility • Hardness range from 30–70 Shore “A” • Low coefﬁcient of friction • Low compression set Disadvantages and Limitations • High material cost • New metering equipment is required • Modiﬁcations or dedicated injection molding machine is required

Figure 1-136 Injection mold vacuum seal (Courtesy: GE)

• Special molds and dual high temperature controller are required • Material recycling is not possible • Limited storage life (one year) for the original unopened resin at temperatures of 73 °F or below Typical Applications • Medical and Health Care: – Foley catheters: Funnel and tip – Surgical irrigation bottle: Pull tabs for surgical, irrigation bags – Respiratory apparatus: Gaskets, nose pieces, diaphragms – Skin contact devices: Ear plugs, military/industrial masks, sports goggles, diving masks, snorkels, camera eye pieces, hearing aid ear pieces, eye glass nose pads – Dental mixing cups: For mixing epoxy compounds – Baby bottle nipples: Nipples, paciﬁers – Cartridge chart rollers: For scanners – Cap liners: For medical and drug laboratory containers

Figure 1-137 Extruded flexible tubing (Courtesy: GE)

• Aerospace and Military: – Electrical connectors: Connector seals – Valves: Safety relief valve for life rafts and life vests – Diaphragms: Pressure sensing device for mines • Appliances, Electrical and Industrial: – Connector seals: Connectors for refrigerators, freezer lights, air conditioners – Bellows: Sealing of outdoor electrical apparatus – Diaphragms: Pneumatic valve actuators – Switch seals: Electrical equipment used outdoors – Nozzle for welding torches: Insulation of welding nozzles – Steam irons: Tubing for water/steam – Seals for oil exploration equipment: Encapsulation down hole instrumentation – Chip carrier trays: For baking semiconductor chips – RFI-EMI shields: Binder for metallic mesh – Camera bellows: For instant cameras – Temperature probes: For toaster ovens • Automotive: – Light bulb holders: For rear lights of heavy duty trucks – Light diffusers: For illumination of dashboard instruments – Connector seals: For electronic control module

Figure 1-138 Parts molded in different processes (Courtesy: GE)

Figure 1-139 Molded roller, washers and stop bottoms (Courtesy: GE)

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1 Polymeric Materials – – – – – – – –

Figure 1-140 Electrical weather-pack seals (Courtesy: GE)

Switch seals: Power window controls Fasteners/shock absorbers: Power seal controls Light gaskets: For water-proof lights Rain gutter gaskets: For PVC gutters Nail gaskets: For automatic nail guns Mercury vapor lamp gaskets: Outdoor illumination seals Wire seals: For telephone function boxes Drip irrigation: Water ﬂow controllers for root irrigation

• Ofﬁce Equipment: – Paper moving rollers: For printers and copy machine – Keyboard pads: Nonmechanical spring action devices for computer terminals and telephones, calculators, etc.

1.5

Thermoset Polymers

Thermoset compounds are organic polymers that cure to a solid and infusible mass by forming an irreversible three-dimensional network of covalent chemical bonds. Thermoset compounds are used in many applications. Construction represents the largest single market area, consuming about half of the compounds produced. Other applications include adhesives for plywood and particle board, binders for insulation, coatings, matrix resins for laminates, and electrical molding products. Thermoset compounds are polymers with a combination of mechanical, thermal, electrical, and chemical resistance properties that allow them to compete with metals, ceramic, and thermoplastic materials. Lacking the strength and stiffness of metals, nearly all thermoset compounds contain particulates or ﬁbrous reinforcements. Fillers, such as calcium carbonate, glass ﬂakes, and wood ﬂour are added to reduce cost and increase the rigidity of the cured product. Fibers, such as glass, carbon, and polyaramid increase its strength, stiffness, and cost. The amount of ﬁllers compounded ranges between 45 and 75%. Compared to metals, thermosets possess corrosion resistance, are lighter weight, have better insulating properties, and can be processed at lower pressures and temperatures. The ﬂow characteristics of uncured thermoset compounds can be used to form large and complex shapes in one mold, allowing part consolidation and elimination of machining costs. The advantages of metal materials are high temperature performance, thermal and electrical conductivity, isotropic properties, ductility, and dimensional control. Compared to ceramic materials, thermosets offer lighter weight, better toughness, and easier processing. Ceramic materials provide improved high temperature performance, excellent chemical resistance, and hardness. For many applications, both thermoset and engineering thermoplastic resins are viable candidates. Selection of the preferred material depends on the speciﬁc combination of required properties and processing characteristics. Thermoset compounds offer advantages in terms of reduced creep and improved solvent/ stress crack resistance. The three-dimensional polymer networks in thermosets also improve the machinability, provide low rates of gas permeability, low mold shrinkage, and high temperature performance. The low initial viscosity of thermoset compounds permits the use of large amounts of ﬁllers or ﬁbrous

1.5 Thermoset Polymers reinforcements, which has led to the development of many low cost compounds. Unsaturated polyester and epoxy compounds are used for large reinforced structures (tanks and boat hulls), because these products can be cured at or near room temperature at ambient pressure. Engineering thermoplastic resins have the advantage in large volume production processes, where the injection molding cycle is fast and automatic. Although cure times below one minute are possible with some thermoset compounds, injection molding cycles for engineering thermoplastics are generally faster and less affected by small changes in processing conditions. The major limitation of thermosets is their poor impact resistance. Consequently, engineering thermoplastics are increasingly being considered for structural applications that require enhanced toughness. Although many thermoset compound properties are dominated by the type and amount of reinforcement rather than the basic matrix type, interest in engineering thermoplastics is high because of the capability of recycling and cost savings in the molding process compared to thermoset compounds. Thermoset compounds have excellent heat, creep, UV, and chemical resistance; they also have excellent dimensional stability, electrical properties, surface ﬁnish and low raw material cost. The key features of the main thermoset resin families are: • Polyesters vary from extremely ﬂexible and resilient to very hard and brittle; from water-sensitive to chemical and UV resistant; from ﬂammable to nonburning. The polymerization depends on maleic or fumaric acid. Allylic compounds from a special class of polyesters are polymerized through the double bond of allyl phthalates. These compounds have superior heat, chemical, insulation, and abrasion resistance, good dimensional stability, and surface ﬁnish. • Phenolics in their unmodiﬁed form are extremely hard and brittle, but with alcohol modiﬁcations, they become ﬂexible. The reinforced grades have the highest creep resistance of all plastics, while featuring low cost, excellent heat and water resistance, good chemical resistance, and electrical and mechanical properties. • Melamines exhibit excellent heat and dimensional stability, electrical insulation, and moisture resistance; they are extremely hard and scratch-resistant. Reinforced melamines are used in electrical applications; however, most applications use the unreinnforced form. • Silicones offer a unique combination of organic and inorganic properties. Their thermal stability from 500–700 °F is outstanding. They also have excellent ﬁre, water, and chemical resistance and excellent electrical properties. Silicones require high pressure for molding and their material cost is high. • Epoxies are relatively low in molecular weight; they are outstanding for their adhesive properties, excellent electrical properties, excellent strength, thermal, and dimensional stability, and excellent chemical and wear resistance.

1.5.1

Polyester Alkyd (PAK)

Polyester alkyds are formed by the reaction of polyhydroxy compounds with unsaturated maleic acids. The resulting unsaturation in the polymer backbone

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1 Polymeric Materials

General Properties of GR PAK Speciﬁc gravity

2.40

Tensile modulus @ 73 °F (psi)

1.90–2.00

Tensile strength @ yield (Kpsi)

3.0–10.0

Elongation at break @ 73 °F (%)

2.0–6.0

HDT (°F) @ 264 psi

350–500

Thermal limits service temp. (°F)

300 (short) 250 (long)

Water absorption @ 24 h (%)

0.10–0.50

Shrinkage (%)

0.10–1.00

Notch Izod impact73 °F (ft-lb/in) @

0.20–3.20

Dielectric strength (V/Mil)

250–530

Dielectric constant @ 50 and 100 Hz

3.0–8.0

Dissipation factor @ 50 and 100 Hz

0.007–0.11

Process temp. (°F)

290–350

can be utilized to cross link the polymer and form a thermoset compound. Polyester alkyd thermoset compounds are used for injection, transfer, and compression molding. A small percentage of high viscosity monomer compounded in the resin yields a relatively high melt ﬂow rate, suited for large and complex parts. This beneﬁt, together with excellent electrical properties, dimensional stability, and low cost make it suitable for a variety of applications in the automotive and electrical industries. Polyester alkyd compounds can be cured without additional molding pressure and do not release water when cured. Therefore, they can be used in a variety of coating applications. Polyester alkyds are primarily electrical materials; they combine good insulating properties at temperatures up to 300 °F for intermittent and 250 °F for continuous use, low resin cost, and good insert molding characteristics for delicate and complex inserts. Polyester alkyd compounds are also reinforced with ﬁber glass and minerals to provide substantial improvements in physical strength, impact resistance, and during the molding process. Advantages • Lower cost resins • Excellent dimensional stability and heat aging resistance • Good electrical and mechanical properties at high temperatures • Excellent insert compression moldability and high melt ﬂow rates • Processable by several molding methods at low pressures • Uniform mold shrinkage and fast cure cycles • Compounds are available in several grades for different processes, melt viscosities, toughness, strength, insulation resistance, and ﬁnishes.

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1.5 Thermoset Polymers Disadvantages and Limitations • Chemical resistance is marginal • Poor solvent resistance Typical Application • Small electrical components, circuit breakers, coil forms, capacitors, and resistor encapsulations

Figure 1-141 Rowing shell

• Boat hulls and other ﬁber reinforced items • Sheet molding laminations • Coatings and ﬁlled molding compounds • Compounds can be modiﬁed with styrene or acrylic monomers • Suitable for ﬁberglass composites

1.5.2

Figure 1-142 Boat hull

Diallyl Phthalate/Isophthalate (DAP, DAIP)

General Properties of DAP Speciﬁc gravity

1.94

Tensile modulus @ 73 °F (psi)

1.40

Tensile strength @ yield (Kpsi)

7.50

HDT (°F) @ 66 psi @ 264 psi

500–670 450–550

Thermal limits service temp. (°F)

430 (short) 390 (long)

Water absorption @ 24 h (%)

0.12–0.20

Shrinkage (%)

0.10–0.90

Notch Izod impact73 °F (ft-lb/in) @

1.00

Dielectric strength (V/Mil)

400–450

Dielectric constant @ 106 Hz

3.8–4.4

Volume resistivity (Ohm-cm)

1010–1016

Dissipation factor @ 106 Hz

0.011- 0.017

Process temp. (°F)

290–350

Diallyl phthalate offers a balance of electrical insulating properties, volume resistivity, dielectric strength, and arc resistance. DAP retains these properties even under long-term exposure to high heat and humidity conditions. Diallyl phthalate resins are products of the reaction of allyl alcohol and an organic acid or anhydride. The monomer diallyl phthalate can be prepared by the direct esteriﬁcation of allyl alcohol and phthalic anhydride, which is then partially polymerized to a fusible resin or prepolymer by being heated with a free radical initiator. This diallyl phthalate prepolymer, combined with a free radical initiator and various ﬁllers, constitutes a diallyl phthalate molding compound. There are two molding compound types of diallyl phthalate resins: the orthoresin diallyl phthalate (DAP) and the metaresin diallyl isophthalate (DAIP). The orthoresin is the most commonly used; it provides excellent electrical properties, while the metaresin provides superior heat resistance characteristics.

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1 Polymeric Materials The most frequently used compounds are short glass ﬁber reinforced, which represent approximately 70% of the market. The remaining 30% is divided fairly evenly between long glass ﬁber reinforced and mineral ﬁlled compounds. DAP molding compounds are available only as a ﬁlled system and are supplied complete with catalyst, pigment, and mold release. Its physical form varies with the type of reinforcement used. Mechanical properties of diallyl phthalate vary widely, depending on ﬁller or reinforcement type and quantity. Because of its extremely stable carbon-to-carbon linkage and its tight knit three-dimensional structures, fully cured DAP is extremely resistant to creep or cold ﬂow, moisture, strong and weak acids, alkalis, and organic solvents. Moisture has very little effect on the dielectric strength of the molded parts. Figure 1-143 Automotive distributor cap

Thermal properties may be classiﬁed into long- and short-term heat resistance. DAIP has better heat resistance than DAP, with use temperatures approximately 20 °F higher than the typical 370 °F of DAP and typically has heat deﬂection temperatures from 450–550 °F. DAIP has a UL temperature index of 390 °F. DAP molding compounds process extremely well, using any conventional thermoset molding equipment, such as compression, transfer, or injection molding. Diallyl phthalate molding compounds have a high degree of dimensional stability when molded. Mold temperatures may slightly affect part dimensions with differences in the coefﬁcient of thermal expansion. The degree of cure affects several DAP properties: chemical resistance, heat deﬂection temperature, and most electrical properties will be reduced without proper cure.

Figure 1-144 Cook ware handle

A wide variety of molding compounds is made from these two base resins. Fillers, such as short and long glass ﬁbers, minerals (talcs, clays, aluminum trihydrates), and synthetic ﬁbers are compounded with the resins to obtain speciﬁc properties. Because of the catalyst systems used, DAP has a room temperature shelf life of over a year. Advantages • Excellent electrical insulating properties, volume resistivity, dielectric strength, and arc resistance Figure 1-145 Electrical bobbin

• Excellent creep or cold ﬂow resistance under long-term exposure to high heat and humidity conditions • Good moisture resistance • Low burning and self-extinguishing grades are available • Heat deﬂection temperatures from 450–550 °F at 264 psi • DAIP compound has a UL temperature index of 390 °F

Figure 1-146 Computer chip housing

• Good chemical resistance to strong and weak acids, alkalis, and organic solvents. Disadvantages and Limitations • Low impact resistance • Mold temperatures and curing time are critical process parameters

Figure 1-147 Computer micro processor housings

• Expensive compound • Not recommended for use in contact with phenols and oxidizing acids

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1.5 Thermoset Polymers Typical Applications • Electrical and electronics, connectors, potting cups, switches and electrical bobbins. DAP connectors are used in aerospace or military applications (100% reliability). DAP connectors are used in communications equipment and computers, in wet or humid environments • Appliance handles • Automotive distributor caps

1.5.3

Melamine Formaldehyde (MF)

General Properties of MF Speciﬁc gravity

1.94–2.00

Tensile modulus @ 73 °F (psi)

2.40

Tensile strength @ yield (Kpsi)

5.90–7.00

HDT (°F) @ 264 psi

360–400

Thermal limits service temp. (°F)

300 (short) 170 (long)

Water absorption @ 24 h (%)

0.09–0.30

Shrinkage (%)

0.70–0.90

Notch Izod impact73 °F (ft-lb/in) @

0.25–0.35

Dielectric strength (V/Mil)

170–370

Dielectric constant @ 106 Hz

6.60–7.90

Volume resistivity (Ohm-cm)

1012

Dissipation factor @ 106 Hz

0.013–0.016

Curing temp. (°F)

310–335

Figure 1-148 Electric circuit breaker housing

Melamine formaldehyde molding compounds are produced by the reaction of the NH2 amino group with formaldehyde. Melamines are produced with α-cellulose ﬁllers, which permit an unlimited range of colors. They exhibit excellent surface hardness and resistance to abrasion, excellent compressive strength and resistance to deformation under load. They are also good insulators and have excellent electrical and thermal resistance. Melamine formaldehyde compounds reinforced with ﬁber glass and/or minerals have better electrical, thermal, and impact strength properties; however, they are available in fewer colors.

Figure 1-149 Heated plate for baby food

Because melamine formaldehyde compounds ﬁlled with cellulose have been approved by the FDA for use in food contact applications, dinner ware is the single largest end use of melamine molding compounds. Melamine formaldehyde is listed in Underwriter’s Laboratories index of recognized components. The chemical resistance of melamines is poor; under elevated temperatures they are attacked by strong acids and alkalis. Urea formaldehyde is generally available only with cellulose fillers, while melamine formaldehyde molding compounds are available with cellulose, mineral, or glass ﬁber ﬁllers. Melamine formaldehyde molding compounds can be compression, transfer, or injection molded. Although the materials are available in powder form, a

Figure 1-150 Office accessories

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1 Polymeric Materials granular form is preferred for most molding applications. Injection molding requires the use of equipment designed speciﬁcally for thermosets. Melamines are available in a broad range of ﬂow rates and cure capabilities. Because water is given off during curing, the molding equipment must include a means to remove this moisture. Advantages • Melamines ﬁlled with cellulose have been approved by the Food and Drug Administration for use in food contact applications • Good insulators, both electrically and thermally Figure 1-151 Dinnerware in various colors

• Melamine ﬁber glass and/or mineral reinforced grades have excellent electrical properties • Excellent surface hardness, abrasion resistance, and compressive strength properties • Melamines can be processed by compression, transfer, and injection molding • Alpha cellulose ﬁlled compounds come in an unlimited range of colors and exhibit good physical and electrical properties • Melamines have excellent storage life charactersitics Disadvantages and Limitations • Melamine has poor chemical resistance to strong acids and alkalis. • Melamine ﬁlled with ﬁber glass or mineral are limited in color availability and their speciﬁc gravity is higher. • Moisture removal equipment is required during mold curing. Typical Applications • Electrical Industry: Circuit breaker arc chutes and a great variety of other electrical wiring devices, wall plates, switch toggles, closures, buttons, stove hardware, and small housings. • Common Applications: Dinnerware for household and institutional use, ashtrays, ofﬁce accessories, clothing buttons, and jewelry

1.5.4 Figure 1-152 Extrusion profiles

Cellulosic Ester

Cellulose nitrate, more than 100 years old, is man’s ﬁrst successful effort in modifying a natural polymer to improve its processibility. Celluloid and pyroxylin were two forms of this resin used in many molded components and coated articles. Also known as gun cotton, it is rarely used today in plastic applications because of its high ﬂammability. It has been almost entirely displaced in thermoplastic applications by cellulosic thermoplastic compounds, such as cellulosic acetate (CA), cellulosic butyrate (CAB), and cellulosic proprionate (CAP). Advantages • Good electrical properties

Figure 1-153 Structure platform

• Good impact strength

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1.5 Thermoset Polymers • Good processibility • Excellent surface ﬁnish • Non-petrochemical base • Moderate resin cost Disadvantages and Limitations • Poor chemical resistance to solvents, alkaline materials, and fungus • High moisture absorption and permeability

Figure 1-154 Automotive steering wheel

• Poor long-term weatherability • Poor ﬂammability Typical Applications • Telephone components, eyeglass frames, audio tape cases, tool handles

Eraser

• Automotive steering wheels, structure platforms • Extrusion proﬁles • Pens external strucrures

1.5.5

Figure 1-155 Pencil external body

Cyanate

General Properties of Cyanate Polymers Speciﬁc gravity

1.10–1.35

Tensile modulus @ 73 °F (Mpsi)

0.45–0.50

Tensile strength @ yield (Kpsi)

10.0–13.0

HDT (°F) @ 212 psi

450–500 (dry) 212–390 (wet)

Thermal limits service temp. (°F)

450 (short) 350 (long)

Water absorption long-term @ 212 °F (%)

1.30–2.40

Shrinkage (%)

0.40

Torsion toughness @ 73 °F (ft-lb/in)

0.60–1.20

Coefﬁcient of thermal expansion (10–6 in/in/°F)

33.0–39.0

Dielectric constant @ 106 Hz

2.70–4.90

6

Dissipation factor @ 10 Hz

0.001–0.005

Flammability rating (UL-94)

V1 V0

Curing temp. (°F)

350–480

Cyanate resins, also known as cyanate esters, cyanic esters, or triazine resins, feature the polymerizable functional group –O–C≡N on an aromatic backbone. They are derived from bisphenols or polyphenols and are available as monomers, oligomers, blends, and solutions, providing an alternative to epoxy resins. Cyanate resins are chemically cures when heated. Their functionality undergoes cyclotrimerization to form symmetrically substituted triazine structures. This

Body

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1 Polymeric Materials ring-forming additional polymerization results in a thermoset network of oxygen linked triazine rings (cyanurates) and bisphenol esters/ethers. Cured cyanate resins are classiﬁed as polycyanurates, a type of polyarylate, or wholly aromatic polyester that is cross linked. The unique properties associated with cyanate resins and their ring-forming polycyanurate reaction characteristics are: • Purity. Crystalline monomers are supplied at over 99% pure. Ionic and potentially ionizable impurities are typically less than 10 ppm. • Reactivity. Cyclotrimerization rates are essentially catalyst dependent. A variety of transition metal carboxylates and chelates (latent) are available to provide a cure response ranging from one minute to shelf stable prepreg curing at 350 to 480 °F. • Compounding with Epoxy Resins. Formation of chain extending oxazoline rings and catalysis of excess epoxide permits formulation of hybrids containing up 70 to 80% epoxy resin. • Toughness and Tg. Toughness properties indicated by impact strength, fracture toughness, strain at break and adhesive peel strength are high for 480 °F Tg resins • Volume Expansion after Gelation. Increasing volume noted with conversions of over 65% eliminates stress inducing shrinkage at cure temperature. • Low Dielectric Loss. The dielectric constant (2.7–4.9) and dissipation factors (1–5 × 10–3) are unusually low for high Tg resins. • Low Moisture Absorption. Weight gain in boiling water is (1.3–2.4%). Long term stability in 212 °F water is achieved with epoxy modiﬁcation. Unreinforced polymers such as cyanate homopolymers, toughened bismaleimides, epoxies of the diglycidylether and tetraglycidylamine types are cured with aromatic diamines. Polycyanurates match the toughness and adhesion of difunctional epoxies while providing elevated temperature service mid-range between the more brittle tetrafunctional epoxies and toughened bismaleimides. The epoxy-modiﬁed cyanates systems are used widely in aerospace composite applications and retain a higher percentage of original heat deﬂection temperature. The best performers in long term boiling water/steam environments are formulated with 1.0–1.8 epoxide equivalent. Fiber glass reinforced circuit board laminated with woven E-glass provides the following differences in performance between the unmodiﬁed cyanate resins: Dielectric loss properties are lowest for o-methylated. The ﬂame retardant (100% epoxy) laminates, by comparison, have dielectric constant and dissipation factor values of 4.9 and 0.001, respectively. The laminates are rated self-extinguishing in the UL-94 V1 to V0 categories. The absence of halogen modiﬁcation increases the onset temperature of rapid thermal degradation by approximately 212 °F. The onset temperature of the homopolymer is 765 °F, compared to 545 °F for a ﬂame retardant blend with brominated epoxy resin. Toughness, as indicated by a 90° peel strength for one ounce copper foil cladding test, is superior for unmodiﬁed resins. Epoxy resin blends do not require curing at temperatures exceeding 350 °F and retain an excellent balance of properties at 60% modiﬁcation.

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1.5 Thermoset Polymers Carbon ﬁber reinforced base cyanate composites have upper temperature limits of hot and wet service in the range of 265–320 °F, depending on the test temperature, conditioning, ﬁber form, and post cure temperature. A variety of resin physical forms, including crystalline monomers, amorphous semisolid prepolymers or oligomers, powderable hard prepolymers, solvent solutions and blends, are offered for impregnating woven and unidirectional reinforcements. Advantages • Constant mechanical properties above 85% conversion • No volatiles evolved during cure (epoxy alternative) • Short-term temperature stability to 450 °F • Long-term stability at continuous service temperatures up to 350 °F • Lower dielectric loss characteristics • Excellent adhesive strength at 480 °F • Self-extinguishing ﬂammability rating (V1-V0) • Dicyanates provide greater toughness and increased resistance to strong solvents • Sulfur-linked structures have good solvent resistance • Low cure temperatures, increased steam resistance, and reduced cost • Prepolymers have excellent hot melt processibility characteristics • Dicyanates with Tg in the range from 465–520 °F develop higher fracture toughness and strain at break values • Dicyanates are used as an impact modiﬁer (15–20%) for some thermoplastic polymers • Improved dimensional stability, because denser, finer circuit patterns interconnected between layers are more prone to failure from differential movement caused by dissimilar expansion coefﬁcients and swelling in cleaning solvents, etching agents, and strippers. • Elimination of corrosion caused by resin impurities, epichlorohydrin residues, attacking conductor metals in hot and wet environments. • Lower dielectric loss characteristics reduce dielectric constant and dissipation factor in high-frequency, elevated temperature service. • Improved ﬁeld reparability, needed to prevent blistering and loss of adhesion when mainframe computer boards are soldered manually. Laminate surface temperatures can reach 660 °F, resulting in rapid decomposition of brominated ﬂame retardants. Disadvantages and Limitations • Moisture absorption • High shrinkage during curing. • Limited availability • Comparatively high cost

Figure 1-156 Electronic print circuit board

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1 Polymeric Materials Typical Applications • Adhesives and composites • Thermoplastic impact modiﬁer (15–20%) • Printed wiring boards • Fiber glass reinforced cyanate used for large boat hulls • Composite components for aerospace applications • Replacement for epoxy resin in applications requiring high reliability

1.5.6 Figure 1-157 Boat hull

Epoxy (EP)

General Properties of Epoxy Polymers Speciﬁc gravity

1.84

Tensile modulus @ 73 °F (Mpsi)

3.00

Tensile strength @ yield (Kpsi)

18.0

HDT (°F) @ 212 psi

300–460 (dry) 212–300 (wet)

Thermal limits service temp. (°F)

450 (short) 350 (long)

Water absorption long-term @ 212 °F (%)

2.0–6.0

Shrinkage (%)

0.60

Torsion toughness @ 73 °F (ft-lb/in)

0.40–1.20

Coefﬁcient of thermal expansion (10–6 in/in/°F)

60–70

Dielectric strength (V/Mil)

300

6

Dissipation factor @ 10 Hz

0.030

Flammability rating (UL-94)

HB V0

Curing temp. (°F)

300–430

Epoxy resins are available in a wide variety of thermosetting structures and curing agent variations. The name epoxy is for the products that have in common an epoxy ring consisting of two carbon atoms single bonded to an oxygen atom. There are two types of epoxies. Those polymers made by a reaction with epichlorohydrin are known as glycidyls, while those made by per oxidizing oleﬁns are known as cycloaliphatics. The epoxidized phenols, or phenol glycidyl ethers, are the most commercially important resins, particularly epoxidized bisphenol A, known as the diglycidyl ether of bisphenol A (DGEBA). This resin provides an excellent balance of physical, chemical and electrical properties. Epoxidized phenol novolacs are also available in a range of viscosities. Epoxies are used in combination with a coreactant, or curing agent. Therefore, the ﬁnal composite properties are inﬂuenced by the choice of coreactant. The large number of coreactants available include amines, anhydrides, acids, phenolics and amides, thereby providing a broad spectrum of performance possibilities. Figure 1-158 Bus external panels and doors

After aromatic glycidyl ethers, the epoxidized alcohols, glycols and polyols, or aliphatic glycidyl ethers, are considered most valuable. They are typically used in combination with DGEBA resin to allow better processing. These materials

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1.5 Thermoset Polymers are often very low in viscosity, which makes them highly attractive as dilutents for the more viscous phenolic base products; they can also improve ﬂexibility and toughness properties, although this is generally at the expense of thermal and chemical resistance, particularly if the dilutent is mono functional. The most common members of this group are epoxidized butanol, known as butyl glycidyl ether and the epoxidized long chain mono alcohols. Most epoxies are thermoset materials; however, there are some situations in which high molecular weight versions of DGEBA can be used strictly as a thermoplastic, such as for coating applications to provide very ﬂexible coatings. Most applications can be satisﬁed by an epoxy resin because the resin and curing agent selection is so large. For example, the cured physical properties of different epoxy systems can be modiﬁed by substituting a different curing agent (hardener) to the same epoxy. It is therefore possible to upgrade the performance of a composite structure by simply utilizing a different epoxy resin or curing agent.

Figure 1-159 Mono-train external shell structure

Physical properties can thus be varied over a wide range of rigidity and ﬂexibility. This group of resins created the structure for the adhesive industry, as well as being used extensively in the ﬁber reinforced composite industries. Performance properties of the ﬁnal composite depend on the resin system used. Epoxy composites have ratings from good to excellent for electrical resistance, good chemical resistance, and reasonably high glass transition temperatures. The use of ﬁllers can often raise the Tg, reduce shrinkage, and increase the thermal conductivity and thermal resistance of the composite. The amount of ﬁller used depends on the rheology of the system, the particle size, and the oil absorption tendency of the particular ﬁller.

Figure 1-160 Office chair structural shell

In addition, high ﬂexural strength may require the addition of mica or ﬁber glass, while silica might be best for chemical and abrasion resistance. A major weakness of epoxy resins is their poor ultraviolet resistance and weathering. Processing of high performance composites is quite complex; particularly, if a multiple-step cure schedule is required to achieve the highest possible thermal resistance in the composite structure. Epoxies used in composites are cured by an addition cross linking mechanism, which does not generate volatile byproducts. Part dimensions can be inﬂuenced by the cure, particularly if the cure temperature is too high, because a very vigorous exotherm can be generated, causing excessive shrinkage and changing the composite structure.

Figure 1-161 Electrical power line insulators

Epoxy resins are generally liquids that range from low viscosity for the epoxidized alcohols to very high viscosities for the epoxidized phenol novolac (EPN) resins. Low viscosity EPN resins have recently been developed; curing agents can be low viscosity liquids or solids, such as diaminodiphenyl sulfone. Advantages • Excellent balance of physical, thermal, chemical, and electrical properties • Convenient range of cure conditions from 73–350 °F • Thermal stability for continuous service temperatures up to 350 °F • No volatiles formed during cure • Excellent high-temperature adhesion • Suitable for all thermosetting processing methods

Figure 1-162 Bathroom solid counter top

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1 Polymeric Materials Disadvantages and Limitations • Processing high performance composites is complex • High cure temperatures cause excessive shrinkage and change the structure of the part • Poor ultraviolet resistance and weathering • Poor oxidative stability and some moisture sensitivity Figure 1-163 Epoxy coated corrosion resistant structural steel

• Specialty grades are comparatively expensive Typical Applications • High Performance Aerospace: These resins provide excellent elevated temperature resistance as well as good mechanical properties. The most commonly used epoxy resin is tetraglycidyl methylene dianiline (TGMDA). When cured with agent diaminodiphenyl sulfone, it provides excellent high performance composite structures.

Figure 1-164 Racing bicycle external shell

• General Purpose: DGEBA is the preferred material. The choice of curing agents includes aromatic and aliphatic amines or anhydrides, because they provide the best balance of cost and performance. Applications include pipes, automotive accessories, adhesives, composites, coatings, and sporting goods. • Electrical: Epoxies are used to prepare printed wiring boards and electronic encapsulations, which represent the most important applications. Because ﬂame retardancy is a requirement in these applications, a specialty halogenated epoxy is often used. The curing agent is often an amine curing agent known as dicyandiamide, which needs a cure accelerator, such as N,N-dimethylbenzylamine.

1.5.7

Phenol Formaldehyde (Phenolic, PF)

General Properties of PF Thermosets Speciﬁc gravity

1.74–1.88

Tensile modulus @ 73 °F (Mpsi)

1.90–2.28

Tensile strength @ yield (Kpsi)

6.0–10.0

Elongation at break @ 73 °F (%)

0.20

HDT (°F) @ 264 psi

410–525

Thermal limits service temp. (°F)

450 (short) 350 (long)

Notch Izod impact @ 73 °F (ft-lb/in)

0.75–0.90

Shrinkage (%)

0.20–0.40

Dielectric constant @ 50-100 Hz

5.0–7.0

Dielectric strength (V/Mil)

300

Dissipation factor @ 106 Hz

0.030

Water absorption @ 24 h (%)

0.10–1.00

Curing temp. (°F)

330–390

1.5 Thermoset Polymers Phenolics, discovered in 1907 by Dr. Leo Baekeland, are one of the oldest types of thermosetting compounds. Phenolics are now considered the work horse of the plastics industry. Phenolic molding materials are high-performance thermoset compounds. With the heat and pressure of the molding process, PFs react to form a threedimensionally cross-linked molecular structure. This structure yields excellent dimensional and thermal stability with high load-bearing capability at elevated temperatures. Phenolics are used for close tolerance precision molded components that must function in hostile environments. Phenolic resins are products of the condensation reaction of phenol and formaldehyde. Water is the by-product of this reaction. Substituted phenols and higher aldehydes may be incorporated to achieve speciﬁc properties, such as reactivity and ﬂexibility. A variety of phenolic resins can be produced by adjusting the formaldehyde/phenol ratio, the process temperature, and the catalyst. Two distinct types of resin are produced for use in phenolic molding materials: • Single stage (resole) resin is produced with an alkaline catalyst and a molar excess of formaldehyde. The reaction is carefully controlled to allow the production of a low-molecular weight, noncross-linked resin. Single-stage resins complete the curing reaction in a heated mold with no additional catalyst to form a three-dimensionally cross linked, insoluble, infusible polymer. • Two stage-(novolac) resin is produced by the acid catalyzed reaction of phenol and a portion of the required formaldehyde. The resulting resin product is a brittle thermoplastic at room temperature. It can be melted, but it will not cross link. Novolac can only be cured by the addition of a hardener, which is almost always formaldehyde supplied as hexamethylenetetramine (hexa). Upon heating, hexa decomposes to yield ammonia and the formaldehyde needed to complete the cross linking reaction. The unreinforced phenolic polymer is a very brittle material. However, a wide range of properties can be obtained by using a variety of ﬁllers. These ﬁllers and reinforcements (45–65%) of the formulations impart the mold processing properties. Lubricants, colorants, and other modiﬁers are also used. Wood ﬂour or cellulosic ﬁller yields a molding material with a balance of properties and cost effectiveness. The typical Underwriter’s Laboratories end use temperature index is 300 °F. Improved impact and toughness properties can be achieved with cellulosic ﬁbers while maintaining a substantial creep modulus. Modiﬁcation with mineral ﬁllers up to 45% yields rigidity with improved dimensional and thermal stability. Reduced water absorption and lower coefﬁcient of thermal expansion can also be obtained. Glass ﬁber reinforcement yields substantial improvement in dimensional stability, rigidity, and mechanical properties. Glass ﬁber reinforcement can be tailored to equal the thermal expansion of metals. The UL end use temperature index is 355 °F. Specialty grades can be formulated with graphite, Teﬂon®, and elastomer particles for improved toughness and/or self lubricating capabilities. Phenolic materials are available in black and brown colors. Phenolics are not stable under ultraviolet radiation. They are inert to most common solvents and weak acids and they have excellent resistance to natural oils, fats, greases, petroleum products, and automotive ﬂuids. Resistance to strong acids and alkaline reagents is poor.

95

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1 Polymeric Materials Unﬁlled phenolic resins are available in ﬂake, powder, or liquid form. Filled/ reinforced phenolic molding materials are available in granular, pellet, or ﬂake form. Advantages • Moldability: Phenolics offer both processing and performance advantages; they can be molded by injection, compression, or transfer processes. • Dimensional Stability: Phenolics retain dimensional stability for an indeﬁnite period of time under normal atmospheric conditions. • Creep Resistance: Phenolics have a high degree of resistance to deformation under load, especially at elevated temperatures. • Thermal Stability: Phenolics can withstand 300 °F continuously; certain grades are resistant to 450 °F for short periods and 350 °F for extended periods of time. • Hardness: The cross linked nature of phenolics makes them one of the hardest plastics available. • Excellent electrical insulation properties • Excellent resistance to solvents and automotive ﬂuids • Comparatively low cost • High tensile modulus (rigid) and compressive strength • Self-extinguishing Disadvantages and Limitations • Poor chemical resistance to alkaline reagents • Requires ﬁllers for molding compounds • Poor resistance to bases and oxidizers • Volatiles released during cure (condensation of the polymer) • Limited to dark colors because of oxidation discoloration Figure 1-165 High capacity mini compact disk (CD)

Typical Applications • Industrial: Plywood and particle board, brake and clutch linings, ﬁberglass, cellulose, foam insulation, grinding wheels and coated abrasives, adhesives and glues, coatings and varnishes, electrical and decorative laminates • Electrical: Wiring devices, switch gear, circuit breaker, commutators, brush holders, and connectors • Appliances: Knobs, handles, and heated components for toasters, broilers, and steam irons, motor housings, and timer cases • Automotive: Disc brake caliper pistons, power assist braking components, accessory drive pulleys, water pump housings, solenoids, ashtrays, and transmission components

Figure 1-166 Coffee maker support base

• Special Purpose: Single stage materials are also used for pump housings, vaporizers, steam irons, closures, and to hermetically seal encapsulation for electrical devices

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1.5 Thermoset Polymers

1.5.8

Polybutadiene (PB)

This thermosetting resin is cross-linked by means of its pendant vinyl group in the prepolymer stage. The hardness and cross-link density are a function of the severity of post cure time and temperature. The cross linkable polybutadiene shows surprisingly good thermal stability and electrical properties; particularly in its low dielectric constant. Some modiﬁcations have reactive terminal hydroxyl or carboxyl groups useful for other polymer hybrids, such as polyurethanes. Advantages • Good electrical properties • Low polarity Disadvantages and Limitations • High shrinkage • Thermal capability is a function of the curing conditions Typical Applications • Electrical components, particularly those that require some degree of thermal stability and low dielectric constant • High-temperature ﬁlm • Tubing and hoses

1.5.9

Bismaleimide (BMI)

The bismaleimide (BMI) families of thermosetting resins are excellent materials for high-temperature aircraft and aerospace applications. The reaction of two molecules of maleic anhydride with one molecule of diamine is the processing step for manufacturing BMI polymers. The intermediate bismaleamic acid can be formed at room temperature in solvents such as methylene chloride, toluene, or dimethylformamide. The dehydration, or imidization, reaction is obtained using acetic anhydride with a catalytic of sodium acetate at 195 °F, then reﬂuxing the intermediate acid in dimethylformamide, or heat treating the bismaleamic acid alone. The common manufacturing process for BMI is the reaction product between methylene dianiline (MDA) and maleic anhydride, which results in bis(4maleimidodiphenyl)methane (MDA BMI). To improve processibility and eliminate some of the brittleness associated with the original MDA BMI formulation, a mixture of three BMIs is used; a chain extended BMI results from the addition of the MDA BMI, MDA, and a BMI based on 2,5-diaminotoluene. Another BMI product receives a low melting point from the mixture of the BMIs; MDA, toluene diamine, and trimethyl examethylene diamine. This mixture has a low melting point range between 158 and 257 °F and is capable of impregnating reinforced cloth and ﬁbers by using a hot air gun. This technique allows the formation of complex shapes during vacuum bagging steps. The curing, or polymerization, reaction of BMIs occurs through the maleimide double bond. This type of functionality gives rise to an addition-type curing,

Figure 1-167 Polybutadiene (PB) – applications

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1 Polymeric Materials that is, curing without the evolution of volatiles. Because BMIs are thermoset resins, ﬁnal cured structures have highly cross linked networks that are infusible and insoluble. Depending on whether coreactants are absent or present, BMIs cure differently. By allowing the diallyl compound (toughening agent) to coreact with the double bond of the maleimide, better toughness is obtained. Processing BMI prepregs into parts for structural applications requires an autoclave cure and a free standing post cure. A typical cure may consist of the following steps: Autoclave cure • 1 hour at 144 °F and 100 psi. • 4 hours at 351 °F and 100 psi. • 4 hours at 399 °F and 100 psi. Post cure (free standing) • 4–24 hours at 428–500 °F with a slow cool down at 3 °F/min. Advantages • Continuous service temperature of 350 °F under moist environment conditions • Excellent high strength and corrosion resistance • Excellent high degree of hardness, lower weight, and rigidity • Excellent chemical resistance Disadvantages and Limitations • Brittleness problems • Micro cracking problems • Autoclave curing method for fabricating aerospace parts, the original BMI materials were unacceptable • Long autoclave and post cure cycles • Special manufacturing equipment and processing techniques are required Typical Applications • Aircraft and aerospace components • High temperature electronic components

1.5.10

Unsaturated Polyester (UP)

The commercial use of ﬁber glass reinforced thermoset unsaturated polyester resins began in 1942, when polyester resin was combined with glass ﬁber reinforcement to produce protective housings for radar equipment. Unsaturated polyester resins are used in a wide variety of markets, including construction, marine, and transportation, industrial, electrical, and sanitary ware. Because they are frequently used in conjunction with styrene, fabrication shops using these unsaturated polyester materials have pronounced styrene odor levels.

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1.5 Thermoset Polymers

General Properties of UP Thermosets Speciﬁc gravity

1.75–1.90

Tensile modulus @ 73 °F (Mpsi)

1.90–2.00

Tensile strength @ yield (Kpsi)

10.0–15.0

Elongation at break @ 73 °F (%)

0.50–5.00

HDT (°F) @ 264 psi

390–400

Thermal limits service temp. (°F)

250 (short) 200 (long)

Notch Izod impact @ 73 °F (ft-lb/in)

0.70

Shrinkage (%)

0.20

Dielectric constant @ 50-100 Hz

3.8–6.0

Dielectric strength (V/Mil)

450–530

6

Dissipation factor @ 10 Hz

0.01–0.04

Water absorption @ 24 h (%)

1.00

Curing temp. (°F)

170–320

Unsaturated polyester resins are manufactured by the condensation polymerization of dibasic acids or anhydrides with dihydric alcohols with the dibasic acid or anhydride being partially or completely composed of a 1,2-ethylenically unsaturated material, such as maleic anhydride or fumaric acid. The resultant polymer can vary from a high viscosity liquid (brittle) to a low melt solid material. The polymer is then dissolved in a liquid reactive vinyl (1,2-ethylenically unsaturated) monomer such as styrene, vinyl toluene, diallyl phthalate, or methyl methacrylate to give a solution with a viscosity in the range of 2.0–20.0 Poise. Unsaturated polyester resins are used as replacements for natural materials, such as wood, concrete, marble, steel, and aluminum. Unsaturated polyester resins are commonly chosen because of their ease of fabrication, lower weight, higher strength, corrosion resistance, and lower cost. General classiﬁcations of unsaturated polyesters are based on one common building block used in the polymer chain. Descriptions for each of these ﬁve unsaturated polyester families are listed below: • Orthophthalic polyester resins are manufactured from combinations of phthalic anhydride and either maleic anhydride or fumaric acid (general purpose resins). • Isophthalic or terephthalic polyester resins are the result of a combination of isophthalic acid or terephthalic acid; they are higher quality resins with better thermal resistance, mechanical properties, chemical resistance, and higher in cost. They are easier to process and have better solubility in reactive monomers such as styrene. • Bisphenol A (BPA) fumarates are manufactured by the reaction of propoxylated or ethoxylated BPA with fumaric acid. The aromatic BPA fumarate resins have a higher degree of hardness, rigidity, good chemical resistance, and improved thermal performance. BPA fumarates are used exclusively in high performance applications.

100

1 Polymeric Materials • Chlorendics use a combination of chlorendic anhydride or HET acid with maleic anhydride or fumaric acid. These polymers show some chemical resistance and ﬂame resistance improvements because of the presence of chlorine. Figure 1-168 Racing boat hull

• Dicyclopentadiene can also be incorporated into unsaturated polyesters. The resultant alicyclic groups enhance resistance to thermal oxidative decomposition at high temperatures; these compounds are used in high temperature electrical applications. The use of a reinforcing ﬁber to produce an unsaturated polyester composite dramatically improves both tensile and ﬂexural characteristics.

Figure 1-169 Surf board structure

Inorganic ﬁllers are used in unsaturated polyester resin compounds. They improve stiffness, increase modulus, and they exhibit little effect on corrosion and are used to reduce cost. Mechanical properties at elevated temperatures show the differences in the glass transition temperature (Tg) of the various resins. The rigid and high Tg for BPA fumarate or chlorendic polyesters retain ﬂexural strength up to 250 °F.

Figure 1-170 Automotive external panels, hoods, and doors

Figure 1-171 Racing car external shell

Polyester resins are used in elevated temperature applications, especially in electrical and corrosion resistant areas. Polyester resins have been used for many years in applications requiring resistance to chemical attack. Chlorendics are preferred for use in strong acid environments, especially at elevated temperatures, while BPA fumarate is better in strong basic solutions. However, isophthalic polyesters are the workhorse of the industry. Almost all organic polymers have good electrical properties. Thermal stability and electrical performance at elevated temperatures are directly related, by comparing the retention of dielectric strength at 390 °F with the retention of ﬂexural strength at 390 °F. Vinyl toluene outperformed a styrene-based system in both instances. At temperatures above 390 °F, BPA fumarate outperformed isophthalic polyester when both were used with vinyl toluene. Unsaturated polyester resins are required to have some degree of resistance to burning, which can be accomplished either by simply using a ﬁller or by incorporating halogen into the unsaturated polyester matrix. Advantages • Ease of fabrication, lower weight, higher strength, corrosion resistance, and lower cost • Isophthalic polyester resins have better thermal resistance, mechanical properties, chemical resistance (solubility in reactive monomers) • BPA fumarate resins have a higher degree of hardness, rigidity, good chemical resistance, and improved thermal performance • Chlorendic resins have excellent chemical resistance and some ﬂame resistance • Dicyclopentadiene unsaturated polyesters are widely used in high temperature electrical applications • Tooling construction is fast, simple, and inexpensive • Non-burning halogenated grades are available

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1.5 Thermoset Polymers Disadvantages and Limitations • Orthophthalic resins have poor chemical resistance and process limitations • Orthophthalic phthalic polyester resins have very poor thermal stability performance at elevated temperatures • Isophthalic polyester resins cost more • Poor solvent resistance

Figure 1-172 Home furniture, chairs, tables, and lamps

• Pronounced styrene odor levels Typical Applications • Construction: Pipe, building panels, portable buildings, swimming pools, ﬂoor grating, and doors • Marine: Powerboats, sailboats, canoes, kayaks, gel coating, docks • Automotive: Rear lifts on vans, station wagons, sports vehicles, body panels, truck hoods, trailer panels, structural components, and seating • Industrial: Corrosion control, tanks, process vessels, pipes, ﬁttings, valves, fans, pollution control equipment, scrubbers, hoods, blowers, ducts, stacks, ladders, linings, chutes, connections, sewer lines, and waste water treatment equipment • Electrical: Appliance covers and housings, circuit boards, insulators, and switches • Sanitary Ware: Bathtubs, shower stalls, hot tubs, spas, cultured marble, and food handling containers • Miscellaneous: Hobby castings, decorative art, buttons, bowling balls, skis, ﬁshing rods, and nonstructural furniture parts

1.5.11

Polyimide (PI)

Polyimides are members of a class referred to as heteroaromatics; they are nylon polymers with excellent thermal capability and resistance to temperatures ranging from –270–740 °F for extended exposure times and up to 900 °F short-term for high-temperature specialty grades. PIs can resist higher temperatures better than any other unﬁlled polymer. Polyimides can be produced with either crosslinked molecular structures as thermosets or in linear forms as thermoplastics. Their high glass transition temperatures (Tg) require special processing methods. At the beginning, the polyimide polymers were condensation materials, but at the present time many variations have been developed and introduced to the market. Several approaches have been used to allow the thermosetting reaction to proceed by addition polymerization. Polyimides have high end use temperatures, excellent ﬁre, chemical, and solvent resistance, low coefﬁcient of friction, low wear, abrasion, thermal, and creep resistance, excellent electrical properties, and high mechanical strength properties. The presence of the imide ring in the structure causes some hydrolytic instability, particularly towards alkali. Polyimides are difﬁcult to process; therefore, Du Pont manufactures the basic polyimide polymer, transforms the PI powder into a ﬁlm form for the electrical industry, and laminates the ﬁlm with Teﬂon® (the ﬁlm has the trade name Kepton®). Du Pont also processes the powder into a ﬁber form known as Kevlar®.

Figure 1-173 Indoor spa whirlpool external shell

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1 Polymeric Materials

General Properties of PI – 40% Graphite Fiber Thermosets

Figure 1-174 Vespel® helical gear (Courtesy: Du Pont)

Figure 1-175 Vespel® ball valve seals (Courtesy: Du Pont)

Speciﬁc gravity

1.65

Flexural modulus @ 73 °F (Kpsi)

700.0

Tensile strength @ yield (Kpsi)

7.00–7.50

Notch Izod impact @ 73 °F (ft-lb/in)

0.70

HDT (°F) @ 264 psi

680

Thermal limits service temp. (°F)

900 (short) 600 (long)

Elongation at break @ 73 °F (%)

3.0

Processing temp. (°F)

690

Coeff. of thermal expansion (10–6 in/in/°F)

15.0

Coeff. of friction – metal

0.15 (stat.) 0.09 (dyn.)

Dissipation factor @ 106 Hz

0.0106 0.0034

Pressure velocity limit (psi-fpm)

300,000

Wear factor 10–10 (in3-min/ft-lb-h)

32.0

PI powder compounded with low coefﬁcient of friction additives (Vespel®) is transformed into solid bars, plates, balls, rods, which are provided to molding manufacturers to produce custom parts. Advantages • Continuous use in air at 740 °F; intermediate exposure to temperatures as high as 900 °F for high-temperature specialty grades • Creep is almost nonexistent • Excellent electrical and temperature barriers • High stiffness and strength • Low coefﬁcient of friction • Low wear resistance

Figure 1-176 Vespel® gear box thrust washer (Courtesy: Du Pont)

• Excellent solvent resistance • Excellent adhesion • Fire resistance • Especially suited for composite fabrication Disadvantages and Limitations • Difﬁcult to process • Material handling difﬁculties • Attacked by diluted alkalis an concentrated inorganic acids • Comparatively high cost

Figure 1-177 Kepton® automotive flex print circuit (Courtesy: Du Pont)

• Brown color • Contains volatiles or solvents that must be removed during curing

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1.5 Thermoset Polymers Typical Applications • Fiber for bullet proof clothing and industrial gloves • Fiber reinforcement for automotive tires • Fiber reinforced composite laminates for speed boats, race cars, surf boards, aerospace and airplane structural components • Molding components, such as aircraft brake pads, compression seals, piston rings for transmissions, valves, thrust washers, and bearings • Refrigeration compressors, turbine engine parts, hot glas handling equipment, plasma cutting torchees, and power tools • High temperature films, flexible print circuits, capacitors, automotive electronic sensors, and high performance insulation tapes • Coatings and adhesives • High temperature wire insulation

1.5.12

Polyxylene

These polymers are unique in that they are deposited in the vapor phase by thermal decomposition of a solid dimmer into gaseous diradicals that combine on a surface to form a high molecular weight polymer ﬁlm. Three variations are available: Unsubstituted, monochloro-substituted, and dichloro-substituted. Extremely thin pinhole free coatings and ﬁlms may be deposited. Advantages • Ultra-thin continuous coatings • Complete coverage around and under densely spaced components • Excellent electrical properties • Dielectric strength of 5000 V/mil in thin layers • Highly resistant to organic solvents • Excellent moisture barrier for electronic assemblies Disadvantages and Limitations • Requires special vacuum deposition equipment for application • Air oxidation at temperatures above 257 °F • Difﬁcult to remove the polymer for rework in electronic assemblies • Extra care needed in masking Typical Applications • Coatings for printed wiring assemblies • Insulation of ferrite toroids, sleeves, and bobbins • Insulation of high voltage components to prevent corona and arcing • Insulation for ﬁne gauge magnetic wire • Protective coating for biomedical devices

Figure 1-178 Kevlar® Lotus 81 Formula One racing car (Courtesy: Du Pont)

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1 Polymeric Materials

1.5.13

Polyurethane (PUR)

General Properties of Generic PUR Thermosets Speciﬁc gravity

1.47

Tensile modulus @ 73 °F (Mpsi)

1.00

Tensile strength @ yield (Kpsi)

14.00

Notch Izod impact @ 73 °F (ft-lb/in)

2.20

Elongation at break @ 73 °F (%)

4.0

HDT (°F) @ 66 psi @264 psi

260 150

Thermal limits service temp. (°F)

120–250

Shrinkage (%)

0.30

Water absorption @ 24 h (%)

0.40

Process. Temp (°F)

450–500

Dielectric constant @ 106 Hz

3.2–4.3

Dielectric strength (V/Mil)

400–510

6

Dissipation factor @ 10 Hz

0.003 0.16

Polyurethane (PUR) was developed from polyisocyanate in Germany during Word War II. The general structure of polyurethanes is R–(NCO)n, where n typically equals 2 to 4, or even higher, and R is an aromatic or aliphatic group. The isocyanate group (R–N=C=O) reacts easily with hydroxyl groups (HO-R′). The resulting link between the two residues R and R′ are the urethane group. O N R–N = C = O + HO–R′ → R– O–C–N– R′ | H Polyurethane thermosetting resins are used in coatings, adhesives, binders, sizings, ﬂexible foams, ﬁbers, sealants, biomedical, and external automotive structural applications. Polyurethane is produced in the following forms: Isocyanate: The aromatic isocyanates are the most important members of this family. They are more reactive than the aliphatic types and less expensive. Toluene diisocyanate (TDI): It is the largest volume product of all isocyanates. It is a diisocyanate that is used as a mixture of 80% 2,4-isomer and 20% 2,6isomer. Diphenyl-methane-diisocyanate (MDI): It is the second largest volume aromatic diisocyanate. The commercial products contain mainly three isomers. In addition, MDI can contain components having more than two benzene rings. Polyol, Cross Linker, and Chain Extender: The counterparts for the reaction with the diisocyanates are hydroxyl or amino terminated components. Low molecular weight components are used as chain extenders or cross linkers. In combination with diisocyanate, the high temperature properties are

105

1.5 Thermoset Polymers especially inﬂuenced. The ﬂexible parts of PURs are higher molecular weight compounds. They are not only responsible for the elastic properties, but they also determine the low temperature performance of the polymer. However, they can be terminated with hydroxyl or amino groups, with a functionality of two or more. This means that chemical cross linking can also be introduced through the ﬂexible element. The main classes of polyols are polyethers and polyesters. The polyethers are mostly derived from propylene oxide, while the polyesters are derived from adipic acid. All formulations contain additives, such as catalysts, stabilizers, blowing agents, ﬂame retardants, mold release agents, and other nonreactive additives, such as ﬁllers and pigments. Reaction Product PUR: The combination of the above components in different ratios, with and without water or external blowing agents, leads to a great variety of polymeric products. Because of different porosity levels, the speciﬁc gravity can range from 0.03–1.15. With different chemical compositions, the modulus of elasticity can be varied from values typical for rubber-like materials to values characteristic for engineering thermoplastics. Thermoplastic PUR: Bifunctional raw materials lead to linear PURs when secondary reactions are avoided. Certain combinations of diisocyanates and chain extenders generate PURs that are processed like thermoplastics. Flexible foam is mainly used for cushioning, mattresses, and packaging. Rigid foam is primarily used for thermal insulation; however, in a great percentage of applications, rigid foam also serves as a structural element.

Figure 1-179 Various extrusion profiles used in the home

Semirigid PUR: Semirigid PURs have a good combination of high tensile strength, high elongation, high tear resistance, and excellent ﬂexural fatigue behavior in a broad temperature range. Low-Density Flexible RIM Systems: PURs integral skin foams with low densities are primarily used for bicycle seats and interior car parts. They are predominantly open cell foams with a tough elastic skin, which protects the foam core from mechanical damage. A void-free, smooth, soft surface is an additional advantage. Medium-Density Flexible RIM Systems: With speciﬁc gravities between 0.300– 0.700, these are typical shoe sole materials, but they have also found applications in automobile interiors. High-Density RIM Systems: Semirigid PURs, including solid RIM materials, have speciﬁc gravities from 0.70–1.15; they are mainly used for exterior automobile body parts required to withstand large elastic deformations.

Figure 1-180 Office furniture external structure

Rigid PUR: This material can be rigid block foams or rigid integral skin foams. Rigid Block Foams and Similar Systems: They have cellular structures with closed cells. Because of the cross-linked structure of the matrix material, these foams can be used at temperatures up to 250 °F, depending on the load. The most important property for rigid PUR applications is low thermal conductivity (K factor). Because the foam has a solid material content of only 3–6%, the K factor is determined by the composition of the cell gas. The normal blowing agent is CFCl3, but water is part of the formulation, leading to the production of carbon dioxide (CO2). A considerable increase in the K factor occurs if the foam is soaked with water. A water absorption of only 1% nearly doubles the thermal conductivity, and even a closed cell foam can absorb 2–5% of water.

Figure 1-181 Automotive door energy absorption

106

1 Polymeric Materials Rigid Integral Skin Foams: As with semirigid integral skin foams, the density of the rigid systems can vary widely. Because they usually serve a structural purpose, their speciﬁc gravities are relatively high (0.200–0.800). Reinforced Integral Skin Foam: This polymer was introduced in the late 1970s. The main advantages of a reinforced system are higher stiffness, lower heat sag, and a smaller coefﬁcient of thermal expansion, which allows the production of articles with tight dimensional tolerances. The ﬁllers used are glass ﬁbers, glass ﬂakes, mica, and minerals. Figure 1-182 Bicycle seat

The production of ﬁlled PURs by the RIM process requires premixing the ﬁller with the liquid components (polyol). Special tanks equipped with stirrers are used to disperse the ﬁller. Processing ﬁlled systems requires special precautions because of their abrasiveness and high viscosities. The metering units must be equipped with mono piston pumps. Granular ﬁllers change viscosity only slightly. Milled glass ﬁbers are very common and can be used in concentrations up to 75% in the polyol. The effect of ﬁllers on mechanical properties of the ﬁnal foams follows the same order as observed in their inﬂuence on the viscosity of the raw materials. Granular ﬁllers have a comparatively small effect on stiffness and the coefﬁcient of thermal expansion.

Figure 1-183 Sandwich sailboat hull

Flake materials have a more favorable effect on these properties and improve the dimensional stability at higher temperatures. With 20% mica, for example, the Young’s modulus can be increased by a factor of 3, but tensile strength and elongation at break are reduced. An example of reinforced PUR is the product of the SRIM process (structural reaction injection molding). In fact, SRIMs are composites that contain random, unidirectional or multi-directional glass ﬁber mats. These mats are put into the mold, which is then closed and ﬁlled with the reactive mixture of raw materials. Because of its extremely low viscosity, the resin penetrates the reinforcing mat and solidiﬁes the structure. This procedure allows the use of glass ﬁber contents of up to 70%.

Figure 1-184 Automotive fender panel

The resulting composites combine extremely high stiffness, which is typical for epoxy and polyester laminates, with the advantages of RIM (reaction injection molding) process, i.e., parts can be molded in less than one minute cycles. The manufacturer receives raw materials, such as isocyanates, polyols, blowing agents, and additives. The only exception is TPU, which is delivered as a polymer resin. The isocyanates and polyols are usually liquid. They are simultaneously metered and mixed and the reacting liquid is delivered into a mold where the reaction proceeds until the solid demolding state is reached. Additional components and additives can either be metered directly into the mixing head or prebatched into the polyol. Advantages • Integral skin foams have a porous core and a solid skin; they are produced in a wide range of stiffnesses and densities by RIM processes. • Semirigid molded foams have an open cell structure and considerable elastic hysteresis, providing good mechanical damping properties.

Figure 1-185 Acoustical foam

• Semirigid cast elastomers are solid or micro cellular, and offer high stiffness.

107

1.5 Thermoset Polymers • Rigid integral skin foam has a sandwich-like structure, a porous core and a solid skin, with a transition zone between them (high stiffness and low weight). • Rigid foam has excellent thermal insulation. • The PUR family has a huge variety of properties that can be obtained by combining a relatively small number of basic components in different compositions. • PUR materials can be processed, ranging from simple hand casting to the highly sophisticated RIM processes. Both allow the production of large, complicated parts and design freedom, at relatively low costs.

Figure 1-186 Horseshoes

Disadvantages and Limitations • Diffusion barriers are required to control water absorption. Typical Applications • Integral Skin Foams: They are used in the production of interior parts that have inserts of steel or other materials, such as steering wheels, gear shift knobs, brake handles, head rests, arm rests, instrument panels, and door panels. Typical exterior body parts include spoilers and encapsulated glass, as well as body panels and fascias. • Semirigid Molded Foams: They are used in parts, such as head rests and crash pads, for which energy absorption is important. • Semirigid Cast Elastomers: Typical applications include exterior automotive body parts, elastic load transmission elements, rollers, shock absorbers, and cyclone separators.

Figure 1-187 Furniture finish panels

• Rigid Integral Skin Foams: Typical applications are skylight frames and window frame proﬁles, housings for personal computers, loudspeakers, televisions, radios, wind deﬂectors for trucks, chair shells, tables, book shelves, and furniture. • Rigid Foams: Thermal insulation for the transportation industry, such as frozen food containers for trains, trucks, aircraft, ships, warehouses, portable coolers, display cases, refrigerators, and deep freezers. The automotive industry uses rigid foam as ﬁlling to provide stiffening, sound absorption, and corrosion protection in motor and trunk hoods. Sandwich constructions are for boats, surfboards, skis, and bathroom components. Additional insulation applications for storage vessels, pipelines, doors, roofs, tanks, lightweight concrete material for wall panels, window sills, to ﬁll the gaps between door casings and walls, and insulation for building blocks. • Others: Coatings, adhesives, binders, sizings, ﬂexible foams, ﬁbers, sealants, and biomedical applications.

1.5.14

Silicone (SI)

Silicones (SIs) are suitable as engineering materials for an extensive range of applications. Since their commercial introduction in the 1940s, they have become a universally useful class of materials. Their structure resembles that of silicon dioxide (SiO2) glass, which provides high temperature properties and resistance to radiation, ozone, and chemicals. The structure of SI is very different from that of organic polymers. Speciﬁcally, the polymer chain backbone consists of

Figure 1-188 Electrical insulator

108

1 Polymeric Materials repeat units, whereas the polymer chain backbone of organic polymers is some conﬁguration of repeat units. A Si–O2 bond is stronger and more ﬂexible than a C–C or C=C bond, which makes silicone polymers chemically stronger and more ﬂexible than organic polymers. Silicone polymers resemble the three-dimensional network structures of SiO2 (sand). Organic functionality of SI allows light cross linking between polymer chains, creating rubber elasticity.

Figure 1-189 Electrical plug weather seals

Outstanding characteristics of SI include its ability to maintain properties at both low (down to –150 °F) and high (up to 600 °F) temperatures, and its weathering and chemical resistance, physiological inertness, lubricity, excellent electrical properties, low surface tension, and compression set resistance. Silicone rubber burns to a white, non conductive ash. Burning a specimen is one method for identifying the material. Silicone Fluids The most common product is methyl silicone ﬂuid, best known for its relatively constant viscosity over the range of its temperature resistance.

Figure 1-190 Shaft rotating seals

SIs are an important class of lubricants for industrial as well as household applications. SI greases are used in industry wherever a low temperature (–100 °F) or high temperature (445 °F) resistance is needed. At extreme temperatures, SI greases retain their characteristics. They are used in the automotive and aerospace industries in small motors, ball bearings, gear assemblies, and instrumentation. Room temperature vulcanized materials are also used in mold making because they can provide intricate part reproduction and slight undercuts. It is possible to cast polyester, polyurethane, epoxy, vinyl, and other polymers, as well as low melting metals. SI molds produce excellent prototypes and artistic castings.

Figure 1-191 Prototype mold cavity

Heat curable rubber (HCR) refers to compounded SI rubber with catalyst, color, additives mixed and milled into siloxane polymers. HCR can be extruded or molded and requires a cure time of 10 minutes at temperatures between 250 and 350 °F. For applications requiring service temperatures higher than 150 °F and dimensional stability, post curing at the service temperature or above is recommended. Before molding a catalyzed HCR compound, freshening the material through a calender mixer is recommended, depending on the storage shelf life of the compound. Advantages • Wide range of thermal capabilities • Good electrical properties • Wide variation in molecular structure for ﬂexible or rigid systems • Transparent grades are available • Low water absorption • Inherently ﬂame retardant • Good chemical resistance Disadvantages and Limitations • Attacked by halogenated solvents

Figure 1-192 Automotive spark plug wires

• Comparatively high cost

109

1.5 Thermoset Polymers Typical Applications • Medical implants • Electronics: Room temperature vulcanized materials include potting gel, removable and permanent encapsulation, dielectric gels and/or oils, conformal coatings, calculator and computer touch pads, and insulation. Thermally and electrically conductive products are available for electromagnetic interference and radio frequency interference (EMI/RFI) shielding applications. • Molded Parts: SI is used for high temperature service or in applications that utilize their chemical and weathering resistance. It is used for seals, gaskets, and mechanical diaphragms. HCR matrices can be reinforced with fabrics or ﬁbers for sheeting and tubes, covered roller, and conveyor belts. • Cosmetic Industry: Silicones provide the non-greasy, lubricious texture in lotions, hair preparations, antiperspirants, shaving creams, and lipstick.

Figure 1-193 Automotive engine valve cover seal

• Pharmaceutical: SI is used in ointment bases, over-the-counter medicines, and as tablet lubricants. SIs are used as reactive blocking agents for specialty syntheses of many drugs. • Industrial: Silicone ﬂuids as anti-foaming agents include use in detergent manufacturing processes, fermentation, adhesive and sealant manufacturing, waste water treatment, boiler and cooler treatment, metal cutting ﬂuids, agricultural chemicals, cleaning solutions, printing, paints and coatings. • Textile Industry: Yarn lubricants in high speed knitting and milling applications, ingredients in fabric ﬁnish formulations. • Paper Manufacturing Industry: Release coatings for paper to complement pressure sensitive adhesive coated paper. • Polymer Additive: Used as internal lubricants, external mold releases, and processing aids. Silicone ﬂuids and/or silanes are ingredients in automotive polishes, furniture polishes, and masonry water repellents. • Adhesive and Sealant: SI is used for bonding metals, polymers, and ceramic substrates to one another.

1.5.15

Urethane Hybrid

Urethane acrylic polymers were developed for use in fabricating injected composites when traditional urethanes were too high in viscosity or when polyesters either required dilution with styrene or were too slow in cycle times for large volume production. These acrylamates are formed by the reaction of two liquid components, an acrylesterol and a modiﬁed diphenylmethane-4,4-diisocyanate (MDI). The ﬁrst component of this polymer, the acrylesterol, is a hybrid of a urethane and an acrylic. The other component, a liquid modiﬁed MDI, contains two or more isocyanate groups that can react with the hydroxyl portion of the acrylesterol molecule. In addition, the acrylesterol reacts with itself by means of the acrylic group and other reactive groups that can be present in the acrylesterol molecule. The single hydroxyl group on the acrylesterol is very reactive with the isocyanate groups on the MDI because the activation energy necessary for this reaction is

Figure 1-194 Automotive engine gaskets

110

1 Polymeric Materials relatively low. However, the urethane reaction does not form large molecular weight molecules, because the acrylesterol has only one hydroxyl that reacts. This results in a very low viscosity mixture during the initial portion of the reaction. The acrylic reaction has a higher activation energy and it requires higher temperatures to react at an appreciable rate. As the urethane is formed, the exothermic heat of reaction increases the reactant temperature until the free radical acrylic reaction is initiated. The polymer is formed rapidly through a snap cure. Appropriate catalysts are used to accelerate the urethane reaction and the free radical reaction. The resultant polymer is a highly cross linked structure. Acrylamate polymers can be tailored to provide broad processing ratios, toughness, and elongation. Combustion modiﬁed versions are also available to meet standard industry burn rate tests. Acrylamate polymers have better high temperature capabilities. They are particularly suited for mold injection into preplaced reinforcement mat. Directional reinforcement combined with continuous strand random mat provides higher strength than composites with random mat alone. The basic fabrication processes for molding acrylamate polymers include transfer molding, reaction injection molding, and use of prepregs. Reaction Injection Molding (RIM) This process consists of metering and mixing two or more liquids and injecting the mixture into a closed mold. Because the components are low in viscosity, the molding process requires low clamp pressures (less than 100 psi). SRIM Processing One difference between the SRIM and RIM processes is the method of adding a reinforcement. In SRIM, reinforcement can be provided by a random ﬁberglass mat or by preforms combined with bi-directional woven reinforcement placed into the mold before injecting the polymer inside the mold cavity. During injection, the chemical mixture flows through the reinforcement, wetting the ﬁbers. As it forms, the polymer reacts with the treated surface of the reinforcing ﬁbers. Within 1–2 minutes after injection, a ﬁnished composite can be removed from the mold. The two liquids that combine to form acrylamate acrylesterol and MDI have characteristics that are ideal for SRIM processing. The viscosity of the injected mixture of acrylesterol and MDI is only about 25.0 Poise at room temperature and even lower at mold temperature. This allows the reacting mixture to ﬂow through the reinforcement without deforming or distorting it. Most important is that the viscosity remains low long enough for the part to be ﬁlled and the mixture to wet the reinforcing ﬁbers. In addition, the mixture is fast reacting, for maximum productivity. Preforms have been used for many years in transfer molding. The method of making preforms uses chopped ﬁberglass shaped on a special screen. Air drawn through the screen holds the chopped ﬁbers in place while a binder applied to the ﬁbers is curing. The ﬁnal preform holds the general shape of the part. Faster methods of making preforms with random continuous ﬁbers or oriented directional ﬁbers use the technique of thermoforming or thermoset binder molding. Thermoformable mats of continuous or directional ﬁbers contain

111

1.5 Thermoset Polymers a thermoplastic binder, which is heated to soften it. The glass mat can then be molded into shape using an inexpensive preform tool in a press. As the mat cools, the binder hardens and holds the mat in the desired shape. A heat curable binder is applied to precut ﬁberglass mats and the mats with binder are placed into a heated mold. The mold is closed for a short time, curing the binder. After the cure, a ﬁnished preform is removed from the mold. A prepreg is formed by reacting the two components in acetone, resulting in an acrylamate prepolymer. The cross linking reaction is avoided at this point. This solution is used to impregnate glass cloth and specialty ﬁbers. Physical properties depend on the type and amount of reinforcement used.

Figure 1-195 Church, large upper structure

Advantages • Acrylamate resin systems produce strong and high ﬂexural modulus ﬁber reinforced composites • Low density • Good performance over a wide range of temperatures • Heat deﬂection temperatures are in excess of 465 °F • Excellent fatigue performance Disadvantages and Limitations • Low notched Izod impact values for unreinforced resins • Reinforced resins are recommended to avoid brittle failures Typical Applications • Automotive Industry: The use of these composites for prototypes has been extended from components to whole body structures of cars and trucks. • Recreational Products: Reinforced acrylamate resins allow cost effective production of large area parts in low to medium volumes. Using resin transfer molding permits the replacement of complex, labor intensive assembly operations. • Electronics and Communication: These resins are especially advantageous in communications equipment because of the dimensional control of the molded parts and the capability of incorporating desired electronic characteristics into the composites. Electromagnetic interference/radio frequency interference (EMI/RFI) shielding can be accomplished by incorporating conductive carbon mat or metallized cloth in the composites. • Industrial Uses: These resins have been used in the aerospace industry, for agricultural equipment, and materials handling equipment.

1.5.16

Vinyl Ester (BPA)

Vinyl ester resins are unsaturated esters of epoxy compounds. The most common versions are the reaction products of methacrylic acid and Bisphenol A (BPA), an epoxy compound dissolved in styrene monomer. Corrosion resistant reinforced resins based on vinyl esters and other compounds have successfully replaced traditional materials, such as glass, carbon steel, concrete, and brick. Vinyl ester resins have many properties of epoxies combined with the processibility of a polyester. The addition of the methacrylate group allows vinyl esters to

Figure 1-196 Portable cooler insulation

112

1 Polymeric Materials

General Properties of Generic BPA Thermosets Speciﬁc gravity

1.55–1.98

Tensile modulus @ 73 °F (Mpsi)

1.70

Tensile strength @ yield (Kpsi)

21.0

Notch Izod impact @ 73 °F (ft-lb/in)

6.0–20.0

Elongation at break @ 73 °F (%)

2.0

HDT (°F) @264 psi

428–550

Th ermal limits service temp. (°F)

150–210

Shrinkage (%)

8.0–9.0

Water absorption @ 24 h (%)

0.10–0.15

Process. Temp (°F)

73.0–300 6

Dielectric constant @ 10 Hz

4.58

Dielectric strength (V/Mil)

350–470

6

Dissipation factor @ 10 Hz

0.008

be cured in ways similar to the curing of unsaturated polyester compounds. The use of styrene or other reactive monomers allows low viscosities to be obtained at ambient temperature. Because of curing and material handling similarities, vinyl esters are often classiﬁed with unsaturated polyesters. The combination of excellent chemical corrosion resistance and good mechanical and material handling properties similar to those of unsaturated polyesters have made vinyl ester resins a very important product for many applications. One of the advantages of vinyl ester resins reinforced with ﬁber materials is that the reinforcement can be oriented in the direction requiring strength. The properties of vinyl ester resins are more important in other areas, such as fatigue resistance, retention of properties at moderate temperatures, impact resistance, and creep resistance. Chlorendic polyester retains its properties up to 210 °F. The higher residual modulus above the glass transition temperature shows that chlorendic polyester has the highest cross link density. The corrosion resistance of composites made of vinyl ester resins compares favorably with metals, thermoplastics, and unsaturated polyester compounds. Their excellent corrosion resistance is attributed to three basic factors. First, the corrosion susceptible ester linkage is shielded by a methyl group. Second, the vinyl groups are very reactive and a complete cure of the backbone is easily accomplished. Third, the epoxy backbone is very resistant to chemical attacks. The following vinyl ester products are available commercially: • BPA epoxy products are available in different molecular weights and styrene content. Cross link density decreases as molecular weight increases, because the methacrylate cross linking sectors are on the ends of the molecular chains. Vinyl ester made with a low molecular weight epoxy will have a high cross link density. When cured, this material will have a high heat deﬂection temperature and good solvent resistance, but lower impact resistance. Resins made with high molecular weight epoxies will have lower cross link densities, less solvent resistance, but with higher toughness.

113

1.5 Thermoset Polymers • Novolac epoxy polymers have been used to improve the heat and high temperature corrosion resistance of vinyl ester resins. Novolac epoxybased vinyl esters also have improved solvent resistance. The novolac structure provides a higher aromatic content and more cross link sectors in pendant positions along the backbone of the molecule. The glass transition temperatures of these resins are 50–100 °F higher than those of conventional vinyl ester resins. Flame retardant vinyl ester resins use halogenated epoxy resins. These materials retain all the desirable properties of vinyl ester resins plus the ﬂame retardant characteristics. • Modiﬁed vinyl ester resins are tailored to satisfy the needs of the applications. A number of vinyl ester resins, such as aliphatic polyol epoxies and epoxy esters, are available. Aliphatic polyol epoxies provide additional resiliency, impact strength, and thermal and mechanical shock resistance. Modiﬁed vinyl ester resins based on epoxy esters of dibasic acids containing carboxy terminated rubbers provide higher elongation, improved thermal and mechanical shock resistance, and improved adhesion to a variety of substrate materials. • Maleic anhydride modiﬁed vinyl esters. The additional reactive unsaturations raise the heat deﬂection temperature of the resin and improve the retention of properties at elevated temperatures. Processing can be based on either ambient temperature or elevated temperature cure systems. Vinyl ester resins are cured by peroxide initiated free-radical polymerization of the reactive unsaturations of styrene and the methacrylate groups. Using these ambient temperature cure systems and additives gives vinyl ester resins a wide range of gel times. Gel times from less than one minute to over three hours are possible. Some of the processes used to fabricate parts with vinyl ester resins at ambient temperature include open molding techniques, such as hands lay up and spray up, ﬁlament winding, and centrifugal casting. Vinyl ester is also used in press thermo molding. In this process, chopped glass ﬁbers and peroxides are incorporated into the resin. Fillers and thermoplastic additives are used to control the viscosity and shrinkage characteristics during cure. When the mixture is thickened between plastic sheets, it is known as sheet molding compound (SMC). If the material is stored in sealed bags or boxes, it is referred to as bulk molding compound (BMC). The transfer molding process at high temperatures is bridging the gap between press thermo molding and open molding at room temperature. The tooling used is similar to that used in an ambient temperature cure. Because the pressure is typically less than 50 psi and is usually about 5 to 20 psi during the injection step, no pressure is used during cure, and the cure cycle can be as short as 90 seconds.

Figure 1-197 Natural gas composite tanks

Advantages of this Process • Faster cycle times than ambient temperature open molding, with a smooth ﬁnished surface on both sides of the part • Lower tooling costs than with press molding • Retention of the ability to place oriented reinforcement in the part for optimum strength • Low costs for parts of low to medium volume

Figure 1-198 Large composite tanks

114

1 Polymeric Materials Polymer Advantages • Good gel ﬂow characteristics • Toughness because of the high elongation relative to the heat deﬂection temperature • High cohesive strength, wet-out, and bonds well to glass ﬁbers and many other ﬁbrous reinforcements • Excellent processibility when using catalyst systems that have low toxicity and are easily dispersed • Excellent chemical resistance to strong acids and strong alkalis over the whole PH spectrum Figure 1-199 Water transfer composite pipes

• Vinyl ester resins are competitive in applications in which temperature and chemical corrosion resistance is required • Good mechanical strength and material handling properties • Excellent corrosion resistance Disadvantages and Limitations • High shrinkage during curing unless ﬁlled • Some systems have volatile components such as styrene Typical Applications

Figure 1-200 Army tent composite poles

• Corrosion resistant reinforced plastic parts, such as pipes, ducts, tanks, electrical equipment, linings, ﬂooring, pumps, oil ﬁeld sucker rods, process vessels, chemical waste water holding tanks, grating, plating and etching tanks, hoods, scrubbers, washer drums, exhaust stacks, and underground gasoline storage tanks • Industrial equipment for pulp and paper, chemical processing, power systems, pollution control, waste water, and mining

115

2

Engineering Product Design

When designing plastic components, success will depend on one prime factor: how well we use the variety of plastic properties and the processing methods for obtaining optimum results. The designer should select the best resin, realizing that it is essential for the resin’s full potential to be exploited to ensure that the molded part will satisfy both functional and cost requirements. Plastics are governed by the same physical laws and the same rules for good design as other materials. These principles can be applied if the polymer properties are suitable for the operating environment of the product being considered. It is necessary to know and understand what the end product must do and under what circumstances it will operate, before a design analysis can be done.

2.1

Understanding the Properties of Materials

There is a big difference between the properties, processing methods, and applications of materials manufactured by various industries. There is not a single material that can be used for all applications. Each new outstanding property developed in a material opens the door for new applications, technologies, and innovations that will improve the efﬁciency and quality of life of the end users. Product designers should compare the properties of various groups of materials (steels, thermoplastics, aluminum alloys, rubber, etc.), because each material has different properties developed for speciﬁc applications and markets and uses different manufacturing processes. All materials have beneﬁts and deﬁciencies (properties, processes, and quality), making it difﬁcult to compare the cost of ﬁnished products made of different materials and processes. The material properties are directly related to the end use applications whether or not one material is better than another. To illustrate this point, a thermoplastic resin cannot replace a structural steel beam used in building construction; the thermoplastic resins do not have the strength, creep resistance, or melt strength to be extruded into thick walled shapes. Thermoplastic beams would also warp in all directions. However, structural beams can be made of thermoset composites, although this is expensive. In less critical applications, such as the housing industry, wood composite structural beams are replacing steel beams, because of their performance and light weight; they are easy to work with and offer a competitive price. A thermoplastic resin cannot replace the steel in automotive disc/drum brake housings, because the product requires dimensional stability, low thermal expansion, and high strength and rigidity at elevated temperatures. Thermoplastic resins do not meet the requirements. However, brake pads made of thermoset polyimide have been successfully used in airplanes. Metals cannot replace automotive rubber tires, bellows, diaphragms, or compression seals, because metals do not have the elasticity, fatigue endurance, wear resistance, and toughness of rubber. Metals are not used for light-weight and compact cellular phone housings, because metals are electrical conductors, heavy, corrosive, and expensive.

116

2 Engineering Product Design

Ferrous metals Nonferrous metals Thermosets Thermoplastics 0

2

4

6

8

10

Specific gravity

Portable electrical tools and small kitchen appliance housings are no longer made of die cast steel or aluminum but have been replaced with nylon and ABS, improving toughness, electrical insulation, and styling, lowering weight and cost reduction.

Ferrous metals Nonferrous metals Thermosets Thermoplastics -460

0

500 1.000 1.500 2.000

Continuous exposure temperature (˚F.) Ferrous metals Nonferrous metals Thermosets Thermoplastics 0

50

100 150 200

Automotive engine cast iron and aluminum intake manifolds are being replaced by ﬁber glass reinforced nylon to improve efﬁciency, lower weight, creating new manufacturing processes, and cost reduction. Automotive steel bumpers, external side panels, and hoods have been replaced with TPE, thermoset composites, and PC alloys to reduce weight, improve styling, and reduce costs.

250

Tensile strength (kpsi)

Water faucet valves made of die cast steel, brass, or copper are being replaced by new designs, updated styles, and colors, using acetal, which eliminates corrosion, providing cost reduction and opening new markets. High performance, large size irrigation valves (from 1.50 to 3.0 in dia.) and small valves (0.75 and 1.00 in dia.) made of die cast steel and brass were successfully replaced with GR nylon 6/12 for the large valves and with GR nylon 6/6 or acetal for the small valves. This improved performance and reliability, eliminated corrosion, and provided cost reduction. Other low performance commercial valves made of rigid PVC (lower cost) are also produced for the irrigation market. Toilet anti-siphon (ballcock) valves made of several brass and copper components were replaced with a multi-functional design in acetal, improving performance, eliminating corrosion, and providing cost reduction. The acetal valves had excellent performance over a 30 year period.

Ferrous metals Nonferrous metals Thermosets Thermoplastics 0

5

10

15

20

25

30

Modulus of elasticity (Mpsi)

The comparison of properties is an effective tool when applied to materials in the same family. To illustrate the point that properties between different material families cannot be compared, Figure 2-1 shows several graphs using different generic property values of the different material families.

Ferrous metals Nonferrous metals Thermosets Thermoplastics 0

40

80

120 160

200

Coefficient of linear thermal expansion (in/in/˚F) x 10-6

Figure 2-1 Comparison of generic properties of materials

The ferrous metal bars include cast iron, cold rolled steels, structural steels, alloy steels, stainless steels, and tool steels. The nonferrous metal bars include magnesium, aluminum, copper, nickel and brass alloys, and titanium. The rubber bars include acrylic, butadiene, butyl, chloroprene, nitrile, silicone, urethane, EPDM, EPM, ﬂuorocarbon, and natural rubbers. The thermoset bars include phenolics, silicones, alkyds, DAP, polyimides, aminos, unsaturated polyesters, epoxies, and urethanes. The thermoplastic bars include ABS, acrylics, acetals, nylons, LCP, PBT, PET, PS, PE, PP, PC, PPO, PEI, PEKK, PSU, PPS, PTFE, PVC, and SAN. The speciﬁc gravity graph shows the unit weight of a material compared to water and reveals that metals are two to eight times heavier than plastics. On a strength-to-weight basis, plastics have a more favorable position, as indicated by the speciﬁc gravity graph. In general, the cost of metals is much higher than plastics. The continuous exposure temperature graph shows that metals have wider temperature ranges than plastics; metals can be used at colder and at elevated temperatures. This property is used for the classiﬁcation and temperature range of plastics. The tensile strength (kpsi) graph shows that metals are much stronger than plastics; metals resist higher forces when being pulled apart before breaking. The tensile strength of a plastic varies with temperature; it decreases with increasing temperature over a much smaller temperature range.

117

2.1 Understanding the Properties of Materials The modulus of elasticity (Mpsi) graph shows that metals have higher resistance to deﬂection for short-term, intermittent, or continuous loading than plastics. Metals have better dimensional stability at elevated temperatures than plastics. Since plastics deﬂect more than metals under the same loading, it is important that metal and plastic parts be loaded using different techniques. Plastics require that the load be distributed in compression mode. The coefﬁcient of linear thermal expansion graph shows that increasing the temperature causes more dimensional changes for plastics than for metals. When plastics and metals are used together and are exposed to the same temperatures, plastic parts become larger than metals; therefore, design compensations should be provided to compensate dimensional change in plastics. The thermal conductivity graph shows that metals are good conductors of heat while plastics are excellent insulators. Despite their relatively low effective temperature range, plastics may be superior to metals as high temperature heat shields for short exposures. A plastic part exposed to a radiant heat source soon suffers surface degradation. However, this heat is not transmitted to the opposite surface as rapidly as in metals. The electrical volume resistivity graph compares only the insulation materials used in electrical applications, (metals are conductors). The dielectric strength graph shows the voltage gradient at which electrical failure or breakdown occurs as a continuous arc; the higher the value the better the material. Plastics have excellent electrical resistance properties, while metals are conductors.

2.1.1

Plastics Selection Guidelines

More than 20,000 thermoplastic grades and over 5,000 thermoset grades have been developed for the plastics industry. Because of the enormous diversity of plastic materials, the selection of the best plastic material for a given application is relatively difﬁcult and time consuming, especially for inexperienced plastic designers. Table 2-1 provides a comparison of plastics and their properties. The table includes the most widely used unreinforced, 30% GR thermoplastic, and reinforced thermoset materials; basic mechanical, thermal and electrical properties, and process temperatures, indicating the process characteristics of the resins. Table 2-1 should be used as a preliminary plastic selection guide. The material properties listed in Table 2-1 were obtained by the resin producers by testing molded bars using ASTM procedures under laboratory conditions. Because most applications are not ﬂat bars, but complex conﬁgurations, the actual properties will be different from the published ASTM properties. The values given are only approximate guides used to compare the values between resins for material selection and for preliminary product design calculations. To obtain precise properties for the new product design and conﬁguration, a prototype mold is required, molding the selected materials, and testing the performance under actual service conditions. This chapter provides detailed information for all important plastics, their chemistry, characteristics, advantages, limitations, and applications. Several plastic organizations, such as ASTM, Modern Plastics, D.A.T.A., Inc., Engineering Plastics, IDES “Prospector” and all the resin suppliers provide data properties sheets.

Ferrous metals Nonferrous metals Thermosets Thermoplastics 0

0.5

1.0

1.5

2.0

2.5

Thermal conductivity (BTU/hr/ft2/˚F/in) x 103 Rubber Mica laminations Glass laminations Thermosets Thermoplastics 105 107 109 1011 1013 1015 1017 1019

Electrical volume resistivity (Ohm-cm) Rubber Mica laminations Glass laminations Thermosets Thermoplastics 0

1.0

2.0

3.0

4.0

Dielectric strength (Volt/0.001 in) x 103

Figure 2-1 (continued)

118

2 Engineering Product Design Table 2-1 Property Comparison for Selected Plastics

Tensile Modulus @ 73 °F (Mpsi)

Tensile Strength @ Yield (Kpsi)

Notch Izod Impact @ 73 °F (ft-lb/in)

Continue Expose Temperature (°F)

Processing Temperature (°F)

Dielectric Strength (Vol/Mil)

Dissipation Factor @ 1.0 × 106 Hz

ABS Unreinforced

1.05

0.30

5.00

2.50 12.00

167 185

410 518

HB

350 500

0.03 0.04

Acrylic Unreinforced

1.17

0.38

7.50

0.03 0.50

150 190

410 575

HB

450 530

0.09

Acetal Unreinforced

1.42

0.400

10.00

1.30

195 230

375 450

HB

560

0.005

HDPE Polyethylene Unreinforced

0.94

0.20

3.50

No Break

158 176

400 535

HB V2

450 500

0.0005

PP Polypropylene Homo Unﬁlled

0.90

0.17

4.00

0.50 20.00

212

390 525

HB V2

450 600

0.002

PS Polystyrene Unﬁlled

1.05

0.45

6.00

0.25 0.60

122 158

390 480

HB V2

300 600

0.004 0.0020

PVC Polyvinyl Chloride Rigid

1.38

0.35

5.90

0.40 20.00

150 185

365 400

HB V1

600 800

0.115

PC – 30% Fiber Glass

1.40

1.25

19.00

1.70 3.00

220 265

430 620

V1 V2

450

0.001

PPO – 30% Fiber Glass

1.25

1.10

14.50

1.70 2.30

200 240

520 600

HB V0

550 630

…

PBT – 30% Fiber Glass

1.53

1.35

17.50

0.90

200 250

470 530

HB V0

750

0.004

PET – 30% Fiber Glass

1.67

1.50

22.0

1.60

392

510 565

V0 5V

430

0.002

LCP – 30% Fiber Glass

1.62

2.25

23.00

1.30

430 465

660 680

V0 5V

640 1,000

0.0019

HTN – 30% Fiber Glass @ 73 °F – 50% RH

1.44

1.50

32.00

1.80

315

580 620

V2 V0

500

0.004

Nylon 6/6 – 33% GR @ 73 °F & 50% RH

1.38

0.90

18.00

2.50

265

530 580

HB V2

400

0.006

PEI – 30% Fiber Glass

1.50

1.30

24.50

1.90

356 390

640 800

V0

495 630

0.0025

PPS - 30% Fiber Glass

1.38

1.70

22.0

1.10

390 450

600 750

V0 5V

450

0.0014

PSU – 30% Fiber Glass

1.46

1.35

14.50

1.10

350 375

600 715

V0 5V

450

0.002

DAP – (TS) Fiber Glass

1.94

1.40

7.50

1.00

390 430

290 350

V1 V0

400 450

0.011 0.017

(EP) Epoxy – (TS) Fiber Glass

1.84

3.00

18.00

0.50

350 4450

300 430

HB V0

380 400

0.02 0.05

(PF) Phenolic – (TS) Fiber Glass

1.74 1.88

1.90 2.28

6.50 10.00

0.75 0.90

350 450

330 390

V1 V0

300

0.03

(UP) Polyester – (TS) Fiber Glass

1.75 1.90

1.90 2.00

10.50 15.00

0.50 18.00

200 250

170 320

V0 5V

450 530

0.01 0.04

(PI) Polyimide – (TS) Graphite Fiber

1.65

0.70

7.50

0.70

600 740

690

V0 5V

500 560

0.010 0.003

Flammability UL-94

Speciﬁc Gravity

Types of Polymers

2.1 Understanding the Properties of Materials Designer Check List

General Considerations : Performance requirements (structural, loading cycle, aesthetic, etc.) : Multifunction design : Product design for assembly : Structural load (static, dynamic, cyclic, impact, etc.) : Product tolerance speciﬁcations : Life of product : Resin selection based on performance of similar applications and end use : Product design for assembly process : Quality of product vs. process : Secondary operations : Packaging and shipping Environmental Requirements : End use temperature : Time, weather, strain, and stress cracks : Others (chemical, lubricants, water, humidity, pollution, gasoline, etc.) Design Factors : Type, frequency, direction of loads : Working stress selected (tensile, compression, ﬂexural, combination) : Strain percentage selected : Load deformation (tensile, shear, compression, ﬂexural, etc.) : Tensile, ﬂexural, initial, secant, yield modulus used (temperature, creep) : Correlating the test results to end use environment conditions : Safety factor : Design product for efﬁcient molding Economic Factors : Cost estimate of the new product : Resin cost vs. molding performance : Number of mold cavities vs. size of machine and automatic fast cycles : Eliminate secondary operations : Redesign part to simplify production

Quality Control Tests Required : Tension : Compression : Flexural : Impact (drop weight, dynatup, etc.) : Torsion, fatigue : Creep (tension, ﬂex, temperature) : Chemical resistance : Weather (outdoors or accelerated) : UL electrical classiﬁcation : UL continuous service temperature : UL temperature index : Final product UL approvals Resin Processing Characteristics : Viscosity and crystallization : Difﬁculties in molding the resin : Melt and mold temperature : Sensitivity to thermal degradation : Directional layout of reinforcements : Frozen stresses : Mold shrinkage control : Molding problems (ﬂashing, voids, warpage, short shots, brittleness, tolerances, surface ﬁnishing, etc.) : Material handling : Percentage of reground (runners and rejected molded parts) allowed to mix with the virgin material : Drying the virgin resin and reground material. : Prototype molding the product (resin behavior unknown) Appearance of Product : Aesthetic product application : Dimensional control, warpage, etc. : Color matching, discoloration : Surface ﬁnishing : Weld lines, sink marks, ﬂow lines : Parting line ﬂash : Gate type, size, number, location : Decoration

119

120

2 Engineering Product Design If the product information and the quality of data available about a material have not been developed by the resin supplier, the designer should develop a check list by gathering all the facts related to the application. A typical designer’s check list has been included here (Table 2-2). It may be used as a guideline to develop a speciﬁc check list for any application. All aspects of the part are covered, including the product end use requirements, the structural considerations, the operating environment, the economics, and the appearance factors. This information is provided for making a quick analysis of the part requirements, such as temperature, environment, product life expectancy, and cycle and rate of loading. Designing with plastics requires maximizing the performance and efﬁciency of the product and the injection molding process. The following basic principles should be adopted in designing plastic products. • Design freedom is achieved using multifunctional design concepts. • When comparing materials that satisfy the requirements, remember that most metals have greater strength than plastics, and that all plastic material properties are time, temperature, and environment dependent. • Metal design principles are very different from the concepts used in plastic parts design. • Polymers are not substitutes for metals; in most designs the product geometry must be redesigned using plastic principles to be successful. We need to remember that there are no bad thermoplastic materials, only bad plastic applications.

2.2

Structural Design of Thermoplastic Components

This section will present principles for structural design of molded plastic parts. The only data provided are what is necessary to illustrate the type of information needed for analysis of plastic design structures. The mechanical properties described are the properties frequently used by designers of plastic components.

Tensile stress, σ, (psi)

E = Modulus of elasticity

P Stress limit

O

L Elastic range

Strain, ε, (%)

Figure 2-2 Stress-strain curve

Figure 2-2 shows two regions of the stress-strain curve. First, the region of low strain (O – L) will be discussed. This region is known as the elastic range; it is pertinent to applications where minimum deformation of the part under load is of prime concern. The second region of low stress (O – P) is known as the stress limit, which is important when the specimen springs back without deformation. The following discussion of creep and relaxation describes the effect of loading time on strength properties within the stress-strain curve. Speciﬁc attention is paid to creep under a constant load and relaxation from a ﬁxed deformation. The design methods present the recommended methods for using the mechanical properties and concepts for designing with plastics. Illustrations are included to show how the equations, originally developed for metal designs, can be modiﬁed. Designing within the viscoelastic modulus utilizes modiﬁed elastic design equations. This method is normally used when deformation of the part is of prime concern. Yield design uses design principles that originate from the principles of plasticity. In this section, the yield stress is the controlling material

121

2.2 Structural Design of Thermoplastic Components variable. It is emphasized that the major difference between metal and plastic designs is the necessity of allowing for the time dependence of the mechanical properties of polymeric materials over the entire range of temperatures and environmental conditions that the part may encounter in use.

2.2.1

Stress-Strain Behavior

To understand the response of the material, design engineers have been using a set of relationships based on Hooke’s law, which states that for an elastic material, the strain (deformation) is proportional to the stress (the force intensity). Roark and Young, Timoshenko, and others have developed analyses based on elastic behavior of materials that exhibit a good approximation of simple elastic behavior over a wide range of loads and temperatures. For high stress levels and repeated loading and creep, more sophisticated analyses have been developed to deal with these types of applications. Unfortunately, Hooke’s law does not reﬂect accurately enough the stress-strain behavior of plastic parts and it is a poor guide to successful design, because plastics do not exhibit basic elastic behavior. Plastics require that even the simplest analysis take into account the effects of creep and nonlinear stressstrain relationships. Time is introduced as an important variable and, because polymers are strongly inﬂuenced in their physical properties by temperature, that is another important parameter to be considered.

GA1

A

Model "A"

One of the results of the viscoelastic response of polymers is to vary the relationship between the stress and strain, depending on the rate of stress application. The standard test used to determine structural properties for many materials is the analysis of the stress-strain curve. Figure 2-4 shows the slope of the curve, which is the elastic constant called Young’s modulus; the stress at which the slope of the curve deviates from the straight line is referred to as the tensile strength; and the stress at which the material fails by separation is called the ultimate tensile strength. In the case of viscoelastic behavior, the shape of the curve will depend on the rate of loading or on the rate of straining, depending on the way in which the test is performed. The modulus can vary over a range of three or four to one within the usual testing range and the material can exhibit ductile yielding at the lower straining rates. The value of the tensile strength and the ultimate strength can frequently vary by a 3 : 1 ratio.

GB1

GB2

B

B 1

2

Model "B"

Figure 2-3 Plastic resin structural models, elastic and plastic range

E = Young´s modulus or modulus of elasticity

Tensile stress, σ, (psi)

The retarded elastic response which occurs in plastic materials is best represented as a spring and dashpot acting in parallel. The creep or cold ﬂow, which occurs in plastics, is represented by a dashpot. The combination best representing the plastic structure would be a spring and dashpot in parallel combination, in series with a dashpot. The basic elements and the combinations are shown in Figure 2-3.

2

1

In order to analyze these effects, mathematical models exhibiting the same type of response to applied forces as plastics are used. The elements that are used in such an analysis are a spring, which represents elastic response because the deﬂection is proportional to the applied force, and the dashpot, which is an enclosed cylinder and piston combination that allows the ﬂuid ﬁlling the cylinder to move from in front of the piston to behind the piston through a controlled oriﬁce.

A

GA2

Failure Ultimate tensile stress Stress limit

Elastic range

It is apparent that, when tensile tests are done on plastics, the loading rates must be speciﬁed to make sure the data have any meaning at all. It also becomes clear

Strain, ε, (%)

Figure 2-4 Young’s modulus

122

2 Engineering Product Design

Surface at specific temperature

Isochronous stress-strain curves

that such conventional data is essentially useless for the design of plastic parts, unless the end use loading rates happen to be the same as those of the test. In order to be useful, the tensile test would have to be run over a wide range of rates and the form in which the data is best presented is a three dimensional plot of stress-strain and time as illustrated in Figure 2-5.

2.2.2

Tensile Testing of Viscoelastic Materials

Creep curve

Str

ess Lo

gt

im

e

Strain

Figure 2-5 Three-dimensional graph, stress-strain-log time

In this section we will address the internal effects of forces acting on a structure. The thermoplastic components will no longer be considered to be perfectly rigid, such as in the static analysis cases. Structural design is concerned with the analysis of material strength, such as the deformations of various structures under a variety of loads. The simple tensile test is probably the most popular method for characterizing metals and so it is not surprising that it is also widely used for plastics. However, for plastics, the tensile test needs to be very carefully performed, because plastics, being viscoelastic, exhibit deformations that are very sensitive to such things as cross head speed rate in tension testing, moisture, stress level, temperature, and creep time. The stress-strain curves as shown in Figures 2-10 and 2-11 illustrate an interesting phenomenon observed in some ﬂexible plastics, such as thermoplastic elastomers. This behavior is known as the plastic range, cold drawing, or continuous elongation of the specimen beyond the yield point without breaking. It occurs because, at low cross head speed rates, the molecular chains in the plastic have time to align themselves under the inﬂuence of the applied stress. Therefore, the plastic specimen’s molecular chains are able to align at the same rate at which the material it is being strained. The simplest case to consider is the application of a straight tensile load on a test specimen of constant cross section. The specimen is loaded at both ends with an equal force applied in opposite directions along the longitudinal axis and through the centroid cross section of the tensile test specimen. Under the action of the applied tensile forces, internal resisting forces are set up within the tensile test specimen. The tensile test assumes that the forces are applied through an imaginary plane passing along the middle of its length and oriented perpendicular to the longitudinal axis of the tensile test specimen. The magnitude of these forces must be equal and directed away from the test specimen (tension loading) to maintain an equilibrium of these forces. Typical tensile test equipment, including an extensometer, is shown in Figure 2-6. Some assumptions are made regarding the variation of these distributed internal resisting forces within the specimen. Because the applied tensile forces act through the centroid, it is assumed that they are uniform across the specimen’s cross section. The load distribution depends on the tensile test specimen geometry, dimensions, and manufacturing process. It also depends on the crystalline molecular structure of the polymer, the coupling agent used to reinforce the compound, and the ﬂow orientation of the material reinforcement. However, to determine the mechanical properties of a polymer by performing either test in compression or tension, the cross head speed rate, at which loading is applied, has a signiﬁcant inﬂuence on the physical properties obtained when running the tests at different loading rates. Ductile materials exhibit the greatest sensitivity of physical property variations at different cross head speed loading rates, whereas these effects are reduced and sometimes negligible for brittle materials.

123

2.2 Structural Design of Thermoplastic Components

Extensometer

Figure 2-6 Tensile test equipment and temperature chamber

2.2.2.1

Stress or Tensile Strength (σ)

Instead of referring to the internal force acting on some small element of the area, it is easier to use the ratio between the force acting over a unit area of the cross section. The force per unit area is termed as the stress (σ) and is expressed in units of force per unit area, e.g., lb/in2 (psi). If the forces applied to the ends of the tensile test specimen are such that the bar is in tension, then the term stress or tensile strength (σ) condition can be applied to the specimen. It is essential that the forces are applied through an imaginary plane passing through the centroid cross section area of the tensile test specimen. 2.2.2.2

Tensile Test Specimen

The tensile test specimen is held in the grips of either an electrically driven gear or hydraulic testing equipment. The electrically driven gear testing equipment is commonly used in testing laboratories for applying axial tension or compression loads. To standardize material testing procedures, the American Society for Testing Materials (ASTM) has issued standard specifications and procedures for testing various metallic, non-metallic, and thermoplastic resins in tension and compression tests. The ASTM test procedures for thermoplastic materials can be found in Chapter 11. Figure 2-7 shows a tensile test specimen speciﬁed for plastic materials. The dimensions shown are those speciﬁed by ASTM for tensile test specimens to ﬁt the grips of the tensile test equipment. The elongations of the tensile test specimen are measured by a mechanical extensometer (see Figure 2-6), an internal gauge (micro-processor tester), or by cementing an electric resistance type strain gauge to the surface of the tensile test specimen. This resistance strain gauge consists of a number of very ﬁne wires oriented in the axial direction of the tensile test specimen. As the test specimen elongates, the electrical resistance of the wire changes and this change of resistance is detected on a Wheatstone bridge and interpreted as elongation.

8.50 inch

0.75 inch

0.50 inch

0.125inch

Figure 2-7 Thermoplastic tensile test specimen

124

2 Engineering Product Design Stress-Strain Curves for Various Materials

Non ferrous alloys

PL

00

Med. carbon steel

Strain, ε , (%)

Figure 2-14 Stress/strain curve for reinforced resins

B Stress, σ , (psi)

Stress, σ , (psi)

0

Figure 2-11 Stress/strain curve for nonferrous alloys and cast iron materials

PL

Rubber / elastomers

Alloy steel Strain, ε , (%)

Figure 2-9 Stress/strain curve for alloy steel

0

0

Strain, ε , (%)

Figure 2-12 Stress/strain curve for rubber or elastomeric materials

Stress, σ , (psi)

PL Stress, σ , (psi)

Y

High carbon steel Strain, ε , (%)

Figure 2-10 Stress/strain curve for high carbon steel

B

PL

Unreinforced resins 0

B PL

Brittle resins

B

0

B PL

Reinforced resins

Strain, ε , (%)

ε0

Strain, ε , (%)

Figure 2-8 Stress/strain curve for medium carbon steel

0

01

Stress, σ , (psi)

0

Y Stress, σ , (psi)

Y

B

Stress, σ , (psi)

Stress, σ , (psi)

Y B

Strain, ε , (%)

Figure 2-13 Stress/strain curve for unreinforced resins

Strain, ε , (%)

Figure 2-15 Stress/strain curve for brittle resins

2.2 Structural Design of Thermoplastic Components 2.2.2.3

Strain (ε)

The elongation over the tensile test specimen gauge length is measured for any predetermined increment caused by the tensile load. From these values the elongation per unit length, called strain and denoted by ε, may be found by dividing the total elongation ∆L by the original gauge length L, i.e., ε = ∆L / L. The strain is usually expressed in units of inch per inch and consequently is dimensionless. 2.2.2.4

Stress-Strain Curve

As the tensile load is gradually increased at a cross head speed rate, the total elongation over the gauge length and the load are measured and recorded continuously at each increment of the load until fracture of the specimen takes place. Knowing the original cross sectional area of the tensile specimen, the stress (σ), may be obtained for any value of the tensile load by applying the following formula: Tensile Stress = σ = W / A where W denotes the tensile load in pounds, and A the original cross sectional area in square inches. Having obtained the numerous values of stress (σ) and strain (ε), the test results are plotted with these quantities considered as ordinate and abscissa, respectively. This is the tensile stress-strain curve or diagram of the material in tension. The stress-strain curve represents the mechanical characteristics or behavior for each type of material, therefore the stress-strain curves assume widely differing geometries for various materials. Figure 2-8 represents the stress-strain curve for a medium carbon steel, Figure 2-9 the curve for an alloy steel, Figure 2-10 the curve for a high carbon steel, Figure 2-11 the curve for nonferrous alloys and cast iron materials, and Figure 2-12 the curve for rubber or elastomeric materials. Tests conducted at room temperature using ASTM recommended proportional limits showed that polyethylene resin, PP copolymer resin, TPE resins, acetal resin, and unreinforced nylon resin (at 50% relative humidity) are materials that yield gradually until break as shown in Figure 2-13. Reinforced nylon resin (at 50% relative humidity), PC glass reinforced resin, and other compounded polymers that have limited elongation characteristics yield a curve as shown in Figure 2-14. Acrylic resin, PET glass reinforced resin, PBT glass reinforced resin, LCP, PF, PAI, PEI, PEAK, dry as molded nylon glass reinforced resins and most brittle compounded resins usually break before yielding occurs, as shown in Figure 2-15. 2.2.2.5

Hooke’s Law

For any material having a stress-strain curve of the form shown in Figure 2-16, the relation between stress and strain is linear for comparatively small values of the strain. This linear relation between elongation and tensile stress was ﬁrst noticed by Sir Robert Hooke in 1678 and is called Hooke’s law. This initial linear range of action of the material is described by the following formula: Stress (σ) = Modulus of Elasticity (E) × Strain (ε) or

Strain (ε) = σ / E

where E (Modulus of Elasticity) denotes the slope or the straight line 0-PL (origin to the proportional limit) as shown in the stress-strain curve Figure 2-16.

125

126

2 Engineering Product Design E S = secant or apparent modulus 4.460 = 120.540 psi 0.037

E O = initial modulus or E = modulus of elasticity 2.875 = = 141.750 psi 0.02

σ

Y E Y = yield modulus = ε Y 5.200 = = 86.666 psi 0.06

Tensile stregth ( σ Y)

6.000 Y

Break

4.460 4.000

B

S Yield point PL Proportional limit

2.835 2.000

Elastic range

0.16

0.12

0.06

0.037

0

Strain at break ( ε B)

Yield strain ( ε Y)

1.000

0.02

Tensile stress, σ , (psi)

5.200

Plastic range Strain, ε , (%)

Figure 2-16 Stress-strain curve of a thermoplastic material

An element is subject to three mutually perpendicular stresses σx, σy, σz, which are accompanied by the strains εx, εy, εz, respectively. By superimposing the strain components arising from lateral contraction due to Poisson’s effect upon the direct strains, we obtain the general statement of Hooke’s law: εx = (1/ E )[σ x − υ (σ y + σ z )] εy = (1/ E )[σ y − υ (σ x + σ z )]

σ z = (1/ E )[σ z − υ (σ x + σ y )]

2.3

Mechanical Properties of Materials

The stress-strain curve shown in Figure 2-13 may be used to characterize several strength characteristics of thermoplastic materials. Proportional Limit (PL) The ordinate of the point PL is known as the proportional limit, i.e., the maximum stress that may be developed during a simple tension test such that the stress is a linear function of strain. For materials having stress-strain curves as shown in Figures 2-11 and 2-12, there is no deﬁned proportional limit. Elastic Limit The ordinate of the point S (just above the proportional limit PL) is known as the elastic limit, i.e., the maximum stress that may be developed during a simple tension test such that there is no permanent or residual deformation when the load is entirely removed (see Figure 2-16). For many materials, the numerical

2.3 Mechanical Properties of Materials values of the elastic limit and the proportional limit are almost identical and the terms are sometimes used synonymously. In those cases, where the distinction between the two values is evident, the elastic limit is usually greater than the proportional limit. Elastic and Plastic Ranges The region of the stress-strain curve extending from the origin (0) to the proportional limit (PL) is called the elastic range; the region of the stress-strain curve extending from the proportional limit (PL) to the point of break (B) is called the plastic range (see Figure 2-16). Factors Affecting Elastic Properties We assumed that the elastic properties of most structural materials, when stressed below a nominal proportional limit, are constant regarding stress, unaffected by ordinary atmospheric variations of temperature, unaffected by prior applications of moderate stress, and independent of the rate of loading. When precise relations between stress and strain are important, as in product design, these assumptions cannot always be made. Thermoplastic materials exhibit a higher modulus of elasticity and a much higher proportional limit when tested rapidly (cross head speed) than when tested slowly. Absorption of moisture greatly lowers both the modulus of elasticity and the tensile strength. Rubber and thermoplastic elastomers have stress-strain curves approximately curved throughout, and neither has a deﬁnite proportional limit as shown in Figures 2-11 and 2-12. For these materials it is customary to deﬁne the modulus of elasticity (E) as the ratio of some deﬁnite stress to the corresponding strain percentage. The quantity so determined is called the secant modulus, because it represents the slope of the secant of the stress-strain diagram drawn from the origin to the point representing the stress chosen. Plasticity Plastic deformation represents an actual change in the distance between atoms in the molecular chain; plastic deformation represents a permanent change in their relative dimensions. In semi-crystalline thermoplastic materials, this permanent rearrangement consists largely of group displacements of the atoms in the crystal lattice caused by slip on planes of least resistance, thus suffering angular displacement. In amorphous thermoplastic materials, the rearrangement appears to take place through the individual shifting from positions of equilibrium of many atoms or molecules, caused by thermal agitation due to the processing energy. Yield Point (Y) The ordinate of the point Y, denoted by the yield stress σY, at which there is an increase in strain with no increase in stress, is known as the yield point of the material. After loading has progressed to the point Y, yielding is said to take place (see Figure 2-16). To deﬁne the yield point in plastics, the true stress-strain logarithmic curve must ﬁrst be calculated from the conventional stress-strain curve. In the conventional stress-strain curve, strain is calculated in terms of the original length of the test specimen and stress is calculated on the basis of the original cross sectional area. As the cross sectional area of the test specimen changes during a tensile test, the true stress-strain logarithmic curve based on

127

128

2 Engineering Product Design the instantaneous dimensional changes is more meaningful. The equations which convert conventional strain, ε, to logarithmic strain, ε and stress, σ, to true stress, σ are: σ = σ (1 + ε) and ε = ln (1 + ε) The following examples illustrate how these equations may be used:

True stress, (psi)

Example 2-1 12.000

Yield point Y

10.000

B

A specimen is stressed to 6,000 psi at a strain of 1.25%. What is the corresponding true stress and logarithmic strain?

B 73˚ F.

8.000

Y

6.000

True Stress = σ = 6,000 × (1 + 0.0125) = 6,075 psi

Logarithmic Strain = ε = 2.303 Log10 (1 + 0.0125) = 0.0125 or 1.25%

113˚ F.

4.000 2.000

Example 2-2 0 0

5

10

15

22

25

30

Logarithmic strain, (%)

True stress, (psi)

Figure 2-17 Unﬁlled nylon 6/6 tensile stress-strain curve (Courtesy: Du Pont)

12.000

True Stress = σ = 12,000 × (1 + 0.15) = 13,800 psi

Logarithmic Strain = ε = 2.303 Log10 (1 + 0.15) = 0.139 or 13.96%

Yield point Y

10.000

B 73˚ F.

8.000

By applying both equations, the true stress-logarithmic strain curves are constructed from the conventional stress-strain curve. Examples 2-1 and 2-2 show that the differences between the conventional and the true stresslogarithmic strain curves are substantial for relatively large strains and negligible for small strains within the modulus accuracy limit.

B Y

6.000

176˚ F.

4.000 2.000 0 0

2

4.5

6

8

10

12

Logarithmic strain, (%)

True stress, (psi)

Figure 2-18 Acetal homopolymer tensile stress-strain curve (Courtesy: Du Pont)

Yield point

2.000

B

Y

1.600

73˚ F.

1.200

B Y

800

176˚ F.

400 0 0

4

8

Based on the original dimensions, a specimen is stressed to 12,000 psi at a strain of 15%. What is the corresponding true stress and logarithmic strain?

12.5

16

20

24

Logarithmic strain, (%)

Figure 2-19 Polyethylene tensile stressstrain curve

In design terminology, there are many different deﬁnitions of the yield point, and therefore the danger of confusion always exists. For plastic materials that yield gradually, the yield point is determined from the true stress-logarithmic strain curve as shown in Figures 2-17, 2-18, and 2-19. The straight line between the point “Y” and point “B” is the plastic range of the tensile stress-strain curve, and the point “Y” that marks the beginning of the plastic range, is deﬁned as the yield point. For plastic materials that yield abruptly, the yield point is deﬁned as the maximum value (deformation or fracture) of the conventional stress-strain curve. The stress corresponding to the yield point is called the yield stress. This deﬁnition for plastic materials which yield gradually differs from the ASTM deﬁnition. Figures 2-17, 2-18 and 2-19 are the true stress-logarithmic strain curves for unﬁlled nylon 6/6 (50% R.H.), acetal homopolymer, and polyethylene at standard ASTM loading rates. These curves illustrate how to determine the yield point. The logarithmic yield strains are shown to be 22.00% for unﬁlled nylon 6/6, 4.50% for acetal homopolymer, and 12.50% for polyethylene. If a test specimen of a ductile plastic material is elongated to relatively large strains, the test specimen will neck down or elongate and the cross sectional area of the specimen becomes smaller with its structure reoriented. The stress-strain curve beyond the yield point is of little value to the designer, except that it adds conﬁdence that the product can resist the stress loads without fracture but with deformation of the product.

129

Tension and Compression Curves

In some structural calculations it is important for plastic designers to know the differences between the stress-strain curves for tension, compression, and ﬂexure. For example, in ﬂexural design analysis, both types of stress-strain curves are needed, because tensile and compressive stresses are present in any plastic part structure. In addition, various materials are considerably stronger in compression than in tension and ﬂexure. The classic example of a nonmetallic material that illustrates this point is concrete. Tensile and compressive stress-strain curves are shown in Figures 2-20, 2-21, and 2-22 for dry molded and moisture conditioned (50% R.H.) unﬁlled nylon 6/6, acetal homopolymer, and acrylic. From these curves, obtained using the ASTM strain rates, the following conclusions can be drawn: • Tension and compression stress-strain curves at small strains are alike. • The modulus in compression is equal to the modulus in tension. • The ﬂexural modulus is lower than the tensile modulus. • For large strains, the compressive stress is higher than the tensile stress. For example, the tensile stress at 10% strain for acetal homopolymer is 10,000 psi, while the compressive stress is 18,000 psi. Therefore, the yield stress in compression is greater than the yield stress in tension.

15.000

TENSION

10.000

Dry as molded

5.000 0.0

COMPRESSION

5.000

50% R.H. 10.000 15.000 10 8

6

4

2

0

2

4

6

8 10

Strain, (% at 73˚ F.)

Figure 2-20 Unﬁlled nylon 6/6 stressstrain curves, tension and compression (Courtesy: Du Pont)

Stress, (psi)

2.4

Stress, (psi)

2.5 Modulus of Elasticity (E)

10.000 5.000 0

TENSION COMPRESSION

5.000 10.000

Yield or Tensile Strength (σY ) The ordinate to the stress-strain curve such that the material has a predetermined permanent deformation or set when the load is removed is called the yield or tensile strength of the material. For the stress-strain curve shown in Figure 2-16, εY is denoted on the strain axis. The ordinate of Y represents the yield or tensile strength (σY) of the material.

15.000 18.000 10 8

6

4

2

0

2

4

6

8 10

Strain, (% at 73˚ F.)

Figure 2-21 Acetal homopolymer stressstrain curves, tension and compression (Courtesy: Du Pont)

The point B is called the breaking strength of the material (see Figure 2-16). Tangent Modulus (Et) The rate of change of stress concerning strain is known as the tangent modulus of the material. It is essentially an instantaneous modulus given by the following equation:

Stress, (psi)

Breaking Strength (B) 15.000

TENSION

10.000 5.000 0

COMPRESSION

5.000

Et = dσ / dε

10.000 15.000 5

2.5

Modulus of Elasticity (E)

The modulus of elasticity (E) is the ratio of the unit stress to the unit strain of the material in tension, or, as it is often called, Young’s modulus. Values of E for various metallic engineering materials are tabulated in machine handbooks, for engineering thermoplastic materials, the modulus of elasticity (E) is the initial modulus (Eo) shown in the stress-strain curve (Figure 2-16) and it is the highest value at room temperature and cross head loading speed rate (i.e., 0.2 in per minute), typically published and provided by the resin manufacturer’s product

4

3

2

1

0

1

2

3

4

5

Strain, (% at 73˚ F.)

Figure 2-22 Acrylic stress-strain curves, tension and compression

130

2 Engineering Product Design properties information. Because the unit strain (ε) is a pure number (percentage of a ratio of two lengths), the modulus of elasticity (E) has the same units as stress, that is lb/in2 (psi). Secant Modulus (ES) or Apparent Modulus (EA) The secant modulus (ES) is the ratio of the unit stress to the unit strain of the material. When a simple tensile force is gradually increased from the initial point “0” up to the point “S” (see Figure 2-16), this is deﬁned as secant or apparent modulus at a predetermined working strain. The secant modulus (ES) or the tangent of the stress-strain curve from the origin “0” up to the point “S” may be calculated by dividing the secant stress (σS) by the predetermined strain (εS). The secant modulus (ES) is equivalent to 85% of the initial modulus (EO). The secant modulus (ES) is the maximum limit where the structural design principles may be applied. It is used for analyzing thermoplastic materials that have well deﬁned elastic limits without suffering a permanent or residual deformation when the applied force returns to its original loading condition. Yield Modulus (EY ) The yield modulus (EY) is the ratio of the unit stress to the unit strain of the material. When a simple tensile force is gradually increased from the initial point “0” up to the yield point “Y” this is deﬁned as the yield modulus. The yield modulus (EY) or the tangent of the stress-strain curve from the origin “0” up to the yield point “Y” may be calculated by dividing the yield or ultimate tensile stress (σY) by the yield strain (εY) of the curve (see Figure 2-16). It is measured in lb/in2 (psi). The yield modulus (EY) is used for determining the maximum load conditions a thermoplastic material can resist before large deformation, failure, or fracture occurs in the plastic product. Percentage Elongation at Break The increase in length (at the gauge length) after fracture divided by the initial length and multiplied by 100 is the percentage of elongation. The percentage of elongation test is performed at various temperatures and at same loading speed rates. This property is considered to be a measure of the ductility of a material.

2.6

Stress and Strain Analysis

In determining stress by mathematical analysis, it is customary to assume that the material is elastic, isotropic, homogeneous, and inﬁnitely divisible without change in properties and that it conforms to Hooke’s law, which states that strain is proportional to stress. In fact, none of these assumptions are strictly true. A structural material is usually an aggregate of semi-crystalline or amorphous resins, ﬁbers, minerals, and additive particles, the arrangement of which may be either random or systematic. When the arrangement is random, the material is essentially isotropic, when the arrangement is systematic, the elastic properties and strength are not equal, the properties are different in the other directions and the material is anisotropic. Finally, very careful experiments show that for all materials there is probably some set and some small deviation from Hooke’s law for any stress. These facts impose certain limitations on the conventional methods of stress analysis and must often be taken into account, but formulas for stress and strain,

2.7 Thermoplastics Elastic Design Method mathematically derived and based on the assumptions stated, give satisfactory results for nearly all problems of engineering design. If Hooke’s law holds true for the material, of which a part or structure is composed, the part or structure usually will conform proportionally to a similar law of load and deformation. The deﬂection of a beam or truss, the twisting of a shaft, the internal pressure in a container, etc., may in most instances be assumed proportional to the magnitude of the applied load or loads. There are two important exceptions to this rule. One is encountered when the stresses due to the loading are appreciably affected by the deformation. An example is a beam subjected to axial and transverse loads or a helical spring under severe extension. The second exception is represented by any case in which failure occurs through elastic instability, as in a slender column. Here, for loads less than the critical, elastic instability plays no part and the load deformation is linear. At the critical load the type of deformation changes, the column bending instead of merely shortening axially and the deformation becomes indeterminate. For any load beyond the critical load, failure occurs through excessive deﬂection.

2.7

Thermoplastics Elastic Design Method

The viscoelastic behavior of thermoplastic materials results in deformations being dependent on the time under load, the temperature, and the environmental conditions. Therefore, when structural components are to be designed using thermoplastic materials, it must be remembered that the classical equations available for the design of columns, springs, beams, plates, cylinders, etc., have been derived under the following assumptions: • The strains are small • The modulus of elasticity is constant • The strains are independent of cross head loading rate and are immediately reversible • The material is isotropic • The material behaves in the same way under tension and compression Because these assumptions are not always justiﬁed for thermoplastic materials, the equations can not be used indiscriminately. For each case, factors such as mode of deformation, service temperature, creep, injection molding conditions, environment, etc., must be considered. In particular it should be noted that the classical equations are derived using the relation: Stress (σ) = Modulus of Elasticity (E) × Strain (ε) The tensile stress-strain curves of thermoplastic materials show that the modulus of elasticity is not a constant. Several approaches have been used to allow for this and some provide very accurate results. The drawback is that the methods can be quite complex for designers. However, one method that has been widely accepted is the “Elastic Design Method”. With this method, using appropriate values of time-dependent properties, such as modulus of elasticity, stresses are selected and substituted into the classical equations. It has been found that this

131

132

2 Engineering Product Design approach gives sufﬁcient accuracy in most cases. Provided that the values chosen for the modulus of elasticity and the stresses take into account temperature, service life of the component, and the limiting strain of the thermoplastic materials. This, of course, assumes that the limiting strain for the thermoplastic material is known. Unfortunately this is not just a value that applies for all thermoplastic materials or even for one plastic in all applications. One method is to use the tensile stress-strain curve and plot a secant modulus that is 85% of the initial tangent modulus (see Figure 2-16). However, for many thermoplastic materials this method is too restrictive and in most practical situations the limiting strain is determined in consultations between the product designer and the thermoplastic material’s manufacturer.

2.7.1

Working Stress

The above mentioned strength characteristics may be used to select the working stress. The working stress is determined merely by dividing either the stress at yield or the ultimate stress by a safety factor. Selection of the safety factor is based on the product application requirements or by the designer’s judgment and experience. Example 2-3

15.00 in. W=200 lb.

The two PET (30% glass reinforced) bars AB and CB are pinned at each end and support a load of 200 pounds as shown in Figure 2-23. The bar AB in tension should have a safety factor of 2.5, while the bar CB in compression should have a safety factor of 2.0. Determine the cross section area of these bars and also the horizontal and vertical components of displacement of point B. A free body diagram of the joint at B assumes that the unknown forces are in tension.

A

B CB =

60˚ .3 17

2i

n.

W (Load) 200.0 = = 400 lb. sin30° sin30°

AB = CB × cos30° = 400 × cos30° = 346.41 lb.

C W=200 lb.

The working stresses are 23,000 / 2.5 = 9,200 psi in tension and 23,000 / 2 = 11,500 psi in compression. The required areas are found by dividing the axial force in each bar by the allowable working stress.

B

A B 30˚

W W σ = → A= A σ

346.41 = 0.0376 in 2 9,200 400 = = 0.0347 in 2 11,500

AAB = ACB

C B Figure 2-23 Structure diagram and load components reaction, where: modulus of elasticity (E) = 1,300,000 psi, tensile strength (σ) = 23,000 psi

To analyze the displacement of point B, it is ﬁrst necessary to calculate the axial deformation of each of the bars. The variation in length in tension is determined by applying the following equation:

133

2.7 Thermoplastics Elastic Design Method

∆=

W ×L A×E

346.41 × 15.00 = 0.106 in 0.0376 × 1,300,000 400.00 × 17.32 = = 0.153 in 0.0347 × 1,300,000

∆ CB

The location of the point B after deformation has occurred and the increment in bar AB or horizontal displacement is 0.106 in and also rotates as a rigid body about the pin at A. In addition, the bar CB shortens 0.153 in and also rotates about the pin at C. Figure 2-24 shows the movement of point B to its deﬂected location B1. The elongation of bar AB is shown by the straight line BB1. The same reasoning applies to the rotation of bar CB. From the geometry of the displacement diagram we can calculate the total vertical component straight line B1B2. ∆V =

H

∆ AB =

0.1

53

0.106 in.

in.

B1 B 30˚

V 30˚

0.106 + 0.153 × cos30° + 0.153 × sin30° = 0.489 in tan30°

∆ H = 0.106 in B2 Figure 2-24 Point “B” deﬂected diagram

2.7.2

Compressive Stress

Compressive properties describe the behavior of a material when it is subjected to a compressive load at a relatively low and uniform rate of loading, which tends to crush the specimen. By choosing the proper test conditions, it is possible to determine the compressive strength of the material. Compression tests for plastics are of limited design value, because plastic products (except foam) seldom deform from compressive loading alone. Compressive properties include modulus of elasticity, yield stress, deformation beyond yield point, compressive stress, and compressive strain. However, compressive stress and compressive modulus are the two properties that are used in part design, resin selection, and thermoplastic material speciﬁcations. In the case of a polymer that fails in compression by a shattering fracture, the compressive stress has a deﬁnite value. For those polymers that do not fail by a shattering fracture, the compressive stress is an arbitrary factor in determining the degree of distortion as the type of failure of the thermoplastic material. The universal testing equipment used for tensile and ﬂexural testing can also be used for testing the compressive stress of various materials. A deﬂectometer or a compressometer is used to measure any change in distance between two ﬁxed points on the test specimen at any time during the test. Figure 2-25 shows a typical specimen set up in compressive testing equipment. Compressive stress is calculated by dividing the maximum compressive load carried by the specimen during the test by the original minimum cross sectional area of the specimen. The result is expressed in psi, either at the rupture of the specimen, or at a given percentage of deformation. Modulus of elasticity or compressive modulus, like tensile and ﬂexural modulus, is also represented by the slope of the initial straight line portion of the stress-strain curve and is calculated by dividing the change in stress by the corresponding change in strain. The method to calculate compressive modulus is the same as the tensile testing procedure.

134

2 Engineering Product Design Movable crosshead speed 0.05 inch/min.

Load direction

Specimen

Fixed head

Figure 2-25 Specimen in compressive testing equipment

It is common practice to report compressive properties with the “Compressive stress at 1% deformation”. It is believed that such values give a more accurate picture of the behavior of plastics under conditions of compression. A 1% deformation is very small and a part could recover from such a deformation and still be usable. Compressive stress at 1% deformation is calculated by dividing the force on the specimen by its original cross sectional area when it has been compressed from 1.00 in to 0.99 in (1%) in length.

2.7.3

Flexural Stress

Flexural stress is the ability of the material to withstand bending forces applied perpendicularly to its longitudinal axis. The stresses induced by the ﬂexural load are a combination of compressive and tensile stresses. Flexural properties are reported and calculated with the maximum stress and strain that occur on the outside edge of the ﬂexural test bar. For thermoplastic materials that break under ﬂexural load during this test, the specimen is deﬂected until rupture occurs in the outer ﬁbers. Many thermoplastic materials do not break, even after large deﬂection of the ﬂexural test bar. The ﬂexural characteristics of these resins make it hard to determine the ultimate ﬂexural stress. In such cases, common practice is to report ﬂexural yield stress when the maximum strain in the outer ﬁber of the specimen has reached 5%. Compression load Specimen Support span

6

5

4

5 3 4 1 2 0 1 3 2

Load cell

Figure 2-26 Procedure “A” specimen in the ﬂexural testing apparatus

The ﬂexural stress test has several advantages over the tensile stress test. If the geometry of the application is like a structural beam and the plastic component service failure occurs in the bending mode, then a ﬂexural test is more practical for design or speciﬁcation purposes than a tensile test. The tensile modulus may be different from the ﬂexural modulus calculated from the outer ﬁber of the bent beam. The tensile specimen alignment is more difﬁcult in the tensile test. Also, the tight clamping of the tensile test specimens creates stress concentration points. The small strains or deformations produced by the ﬂexural test are sufﬁciently large to be measured accurately. There are two basic methods used to determine the ﬂexural properties of thermoplastic materials. Procedure “A” is a three-point loading system utilizing center loading on a simple supported beam structure. A ﬂexural test bar with a rectangular cross section rests on two supports 2.00 in apart and is loaded by means of a loading nose midway between the supports. The maximum axial ﬁber

2.7 Thermoplastics Elastic Design Method

Specimen

W 2

W 2

W 2

W 2 L 3

Load span L

L 3

Support span

Figure 2-27 Flexural stress, procedure “B”

stresses occur on a line under the loading nose. This procedure “A” is especially useful in determining ﬂexural properties for quality control and speciﬁcation purposes. A specimen in the testing apparatus is shown in Figure 2-26. Procedure “B” is a four-point loading system utilizing two load points equally spaced from their adjacent support points, with a distance between load points of one third of the support span. In this procedure “B”, the ﬂexural test bar rests on two supports and is loaded at two points, each an equal distance from the adjacent support point. Procedure “B” is very useful in testing materials that do not fail at the point of maximum stress under a three-point loading system. The maximum axial ﬁber stress occurs over the area between the loading noses. This arrangement is shown in Figure 2-27. Procedure “A” is designed principally for materials that break at comparatively small deﬂections. Procedure “B” is designed particularly for those materials that undergo large deﬂections during testing. The differences between these two procedures are the strain rates (loading speed) used for these tests. The strain rate for procedure “A” is 0.01 in per minute and the strain rate for procedure “B” is 0.10 in per minute. Flexural modulus is a measure of the stiffness during the ﬁrst or initial part of the bending process. The ﬂexural modulus should be called “modulus of elasticity in bending”, but other names are also used, such as modulus of elasticity, elastic modulus, ﬂex modulus, or simply modulus. The ﬂexural modulus is represented by the slope of the initial straight line portion of the stress-strain curve and is calculated by dividing the change in stress by the corresponding change in strain. The procedure to calculate ﬂexural modulus uses the three- or four-point loading deﬂection equations, where deﬂection and force are measured and recorded as with the tensile modulus calculations. When calculating the ﬂexural modulus, several errors can be made. These are largely associated with the fact that stress-strain curves seldom have a truly straight line initial portion and considerable judgment must be used in deciding what line to draw through the stress-strain ﬂexural curve.

2.7.4

Coefficient of Linear Thermal Expansion (α)

The coefﬁcient of linear thermal expansion (α) is deﬁned as the change per unit length of a straight bar subjected to a temperature change of one degree. The value of this coefﬁcient is independent of the unit of length but does depend on the temperature scale used. Temperature changes in a structure create internal stresses just as applied loads do.

135

136

2 Engineering Product Design

2.7.5

Poisson’s Ratio (υ)

When a bar is subjected to simple tensile loading, there is an increase in the length of the bar in the direction of the load, but a decrease in the lateral dimensions perpendicular to the load. The ratio of the strain in the lateral direction to that in the axial direction is deﬁned as Poisson’s ratio (υ). Example 2-4 A rectangular acetal bar 0.125 in thick, 1.00 in wide, and 10.00 in long is subjected to an axial tensile force of 100 lbs. Determine the decreases in the lateral cross section area of the bar due to this load. Where: Modulus of elasticity (E) = 400,000 psi Poisson’s ratio (υ) = 0.35 The loading is axial, hence the stress in the direction of the load Stress (σ ) =

W (Load) 100.00 = = 8,000 psi A (Area) 0.125 × 1.00

The simple form of Hooke’s law for axial loading states that E =

σ σ 8,000 → Axial Strain (ε) = = = 0.02 ε E 400,000

The ratio of the lateral strain to the axial strain is Poisson’s ratio υ=

Lateral Strain → Lateral Strain = υ × Axial Strain Axial Strain

Lateral strain = 0.35 × 0.02 = 0.007 Decrease in thickness = 0.125 × 0.007 = 0.000875 in Decrease in width = 1.00 ×0.007 = 0.007 in

Tensile stress, (psi)

2.7.6

Tensile testing of materials at room temperature (73 °F) has shown that most metals are characterized by a steep linear slope, a small strain, and a well deﬁned proportional limit in the stress-strain curve. This is not the case for thermoplastic materials, they are, in general, more sensitive to moisture and temperature over long periods of time (creep), the crosshead (load) speed rate of testing, and to environmental conditions. Nylon is a good example of this behavior when it is exposed to moisture. The dry-as-molded properties of nylon should never be used to design a product. The 50% relative humidity properties of nylon are recommended for use with the product design calculations.

13.000 11.000 Dry as molded 9.000

hu mi dit y 50 % Re lat ive

7.000 5.000

e la ti ve 100% R

3.000 1.000 0

0 0.2

0.6

1.0

1.4

h u m id it

1.8

y

2.2

Moisture Effects on Nylon

2.6

3.0

Strain, (% at 73˚ F)

Figure 2-28 Unﬁlled nylon 6/6, stress-strain curves, moisture effects (Courtesy: Du Pont)

Figure 2-28 shows how the moisture content in the material can change the results of the stress-strain curve of unﬁlled nylon 6/6, thereby changing the properties such as yield stress, elongation, initial modulus, and toughness. It is very important that the designers of plastic products know under what conditions the material changes its behavior and adapt the design procedure accordingly.

137

2.7.7

Effects of Temperature on the Behavior of Thermoplastics

Thermoplastic materials are used over a wide range of temperatures and the effects of temperature on physical properties must be established before a thermoplastic material can be used for designing a particular product.

Tensile stress, (psi)

2.7 Thermoplastics Elastic Design Method

The service life of a thermoplastic molded part at a given end use temperature will be largely dependent on the requirements of the application and the material selection should be based on heat aging test data of the thermoplastic materials and on actual or simulated end use testing. The effects of temperature on the tensile properties of several thermoplastic materials are given in Figures 2-29, 2-30, 2-31, and 2-32. These stress-strain curves were established by heating the tensile test specimens in an air oven at various temperatures and using a temperature chamber mounted on the universal testing equipment (see Figure 2-3). The changes in tensile properties at various temperatures were measured, recorded, and plotted.

Tensile stress, (psi)

Nylon 6/6 toughened, 33% G. R. 35.000

. ˚F

7.000

12

2˚

6.000

15

5.000

8˚

212

F. F.

˚ F.

3.000 2.000 1.000 0 0

0.5

1.0

1.5

2.0

2.5

3.0

Tensile stress, (psi)

Strain, (%)

73˚ F.

10.000

12 2˚ F. 8.000

6.000

212˚ F . 4.000

2.000

0 0

5.0

10.0

15.0

20.0

Strain, (%)

Figure 2-29 Acetal homopolymer stressstrain curves for two strain ranges at various temperatures (Courtesy: Du Pont)

Unfilled nylon 6/6 at 50% R. H. Tensile stress, (psi)

Exposure of natural- and light-colored thermoplastic resins to elevated temperatures may result in discoloration of the molded part, depending on injection molding conditions. Thermal stabilized additives are compounded with the polymer to minimize discoloration at elevated temperatures. When the thermoplastic materials are subjected to chemicals, oils, grease, water, etc., this could also affect the tensile properties of the thermoplastic materials at elevated temperatures.

73

4.000

The nature of viscoelasticity is such that temperature is of fundamental importance. In thermoplastic materials, the primary bonds are strong covalent bonds along the molecular chains and these bonds are not affected by temperatures, unless high temperatures cause degradation. However, the secondary forces that cause the bonding chains to stay together, restricting their relative movement, may be overcome by increasing temperature. This increases the thermal motion of the molecules, changing the physical properties of the thermoplastic materials. When thermoplastic materials are subjected to elevated temperatures for prolonged periods of time in the presence of air, oxidative degradation will occur. The rate and extent of degradation depends on the type and composition of the thermoplastic material, the temperature, and the time of exposure. This effect reduces the tensile strength and toughness and can eventually lead to surface cracking and brittleness.

8.000

18.000

Yield point

16.000

- 40˚ F.

14.000 12.000 10.000

Yield point 73˚ F. dry as molded

8.000

20.000

73˚ F .

Yield point

6.000

73˚ F. 50% R.H. 200˚ F dry as molded

122˚ F.

Yield point

4.000

15.000

25 0˚ F. 2.000

10.000 0

300˚ F. dry as molded

5.000

Figure 2-31 Nylon 6/6 (toughened, 33% G.R.) stress-strain curves at various temperatures (Courtesy: Du Pont)

0 0

2.0

Strain, (%)

4.0

6.0

8.0

10.0

0

10

20

30

40

50

60

70

Strain, (%)

Figure 2-30 Unﬁlled nylon 6/6 (at 50% R.H.) stress-strain curves at various temperatures (Courtesy: Du Pont)

138 36.000 32.000

0 -4

28.000

Tensile stress, (psi)

Tensile stress, (psi)

2 Engineering Product Design

16.000

. ˚F

14.000

24.000

200˚ F.

12.000

. 73˚ F

20.000

10.000

16.000

8.000

12.000

6.000

8.000

4.000

4.000

2.000

30 0˚ F.

0

0 0

0.4

0.8

Strain, (%)

1.2

1.6

2.0

2.4

2.8

0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

Strain, (%)

Stress, (psi)

Figure 2-32 PET 30% glass reinforced, stress-strain curves for two strain ranges at various temperatures (Courtesy: Du Pont)

8.000

2.8

Stress-Strain Recovery (Hysteresis)

7.000

Thermoplastic components are required to recover after being loaded intermittently in use. The stress-strain hysteresis curves are shown in Figure 2-33. Tensile specimens of acetal homopolymer were initially loaded using the ASTM strain rates of 0.7%, 1.5%, and 2.7%, respectively. The load was released and the strain recovery was observed. This hysteresis testing procedure was repeated ﬁve times and each tensile bar was tested to produce the same initial stress. The tensile specimen tested at 0.7% strain recovered completely. However, the other two specimens tested at higher strain rates elongated without recovering.

6.000 5.000 4.000 3.000 2.000 1.000 0 0

1

0

1

2

0

1

2

Strain, (% at 73˚ F.)

Figure 2-33 Acetal homopolymer, three stress-strain recovery tests (hysteresis) at different strain rates (Courtesy: Du Pont)

3

2.8.1

Creep Behavior of Thermoplastics

The tensile test does not provide information on the creep behavior characteristics of thermoplastic materials. Determining this property requires performing creep tests, which have been widely used to describe the deformation behavior of polymeric materials. In these tests, a constant load is applied to the material and the variation of strain with time is recorded. Normally, a logarithmic time scale is used in the secant or apparent modulus vs. time so that the time dependence after long periods can be included and used as an aid to extrapolation. This curve shows that there is typically an almost instantaneous strain followed by a gradual increase. If a material is linearly viscoelastic, then at any selected time, each line in a family of creep curves should be offset along the strain axis by the same amount. Although this type of behavior may be observed for plastics at low strains and short times, in the majority of cases, the response is nonlinear. The mechanism of creep is not completely understood, but some aspects have been interpreted based on the structures of the polymers. With thermoplastic materials, a particular atom is restricted from changing its position as a result of attractions and repulsions between atoms in the same chain or atoms in adjacent chains. It is generally considered that for an atom to change its position it must overcome an energy barrier and the probability of achieving the necessary energy is improved when a stress is applied. In a semi-crystalline material, the crystalline regions are set in an amorphous matrix. The movement of atoms can

139

2.8 Stress-Strain Recovery (Hysteresis) occur in both regions, but in the majority of cases, atom mobility is favored in the amorphous regions between the cross link chains. Thermoplastic materials have the ability to recover when the applied stress is removed and at a ﬁrst approximation this can often be considered as a reversal of creep. Creep and recovery of plastics can be simulated by the use of simple structural models for materials with elasticity and plasticity. Although there are no discrete molecular structures that behave like the individual elements of the models, they help in understanding the response of these thermoplastic materials.

2.8.2

Creep and Rupture Under Long-Term Load

Most materials will creep or elongate to some extent and eventually fail under a sustained stress less than the ultimate strength (short-time). After a short time at load, the initial creep or elongation (strain) related to stress causes a redistribution in the molecular structure of the specimen; the elastic strain loses its ability to resist or recover and the viscous creep behavior takes over, resisting the load until failure. The viscous creep condition will continue until a break occurs in the specimen, unless the applied load is reduced sufﬁciently. Creep, caused by constant load for a long time, temperature, weather conditions, and moisture must be taken into account in product design and selecting the working stresses for the materials.

2.8.3

Creep and Relaxation of Thermoplastics

The phenomena of creep and relaxation in thermoplastic materials are of prime concern to the designers of plastic products which, in use, may be only slightly deformed yet carry a load for long periods of time. Creep will take place even at low loads and low temperature and the designer must estimate its amount. For metals, analogous behavior is usually only encountered at elevated temperatures, and therefore this situation is of less concern for metal materials. Creep and relaxation can be illustrated by loading tension on the specimen. Figure 2-34 shows a load being applied to the specimen and an initial deformation (∆LO) is observed. The permanent elongation suffered by the specimen with time under a constant load is known as the creep characteristic of the thermoplastic material. Consider a similar test bar being stressed by the same load, but in this experiment, the test bar elongation is conﬁned to its initial deformation (∆LO) as shown in Figure 2-35. The stress level gradually decreases with time and the strain rate also decreases at a constant rate; this characteristic is known as relaxation.

L0

L0

L0

L0 +

L 0+

L 0 (t) Constant strain

L 0 (t) L0 F0

Time = 0 Figure 2-34 Creep effects

C o n st a n t fo rc e

L0

L0+

L0

Creep F0

Time = t

F0

Force decay

Time = 0

Figure 2-35 Relaxation effects

F (t)

Time = t

140

2 Engineering Product Design The creep and relaxation data are measured and presented graphically as shown in Figure 2-36. Strain

Stress

Creep

Basic creep and relaxation behaviors are measured by a simple uniaxial tension test as shown in Figure 2-37. Time-dependent creep modulus is deﬁned: E t (Creep) =

Stress σ W /A W × L = = = Strain ε ∆/L ∆×A

E t (Relax) =

Stress σ W /A W × L = = = Strain ε ∆/L ∆×A

Relaxation Time

Figure 2-36 Constant creep and relaxation in tension

Where: W = Constant force (lb.) A = Original cross sectional area (in2) L = Original length (in) Δ = Increase in length (in)

or

∆=

W ×L A×E

W

W L

Figure 2-37

Tests have conﬁrmed that both of the above equations are applicable when the initial strain is within the ﬂexural modulus. This modulus is used as a checkpoint when making creep (deformation) and relaxation (force decay) calculations. It has been shown experimentally that the creep modulus and relaxation modulus are similar in magnitude and for design purposes may be assumed to be the same. That is, Et (Creep) = Et (Relax). A time-dependent modulus called the secant or apparent modulus (85% of initial modulus) is used as a safety factor. Example 2-5 Determine the apparent modulus from a creep experiment by placing a test bar in tension under a load of 150.00 pounds. The bar has an original length of 4.75 in, a width of 0.75 in, and a wall thickness of 0.10 in, or cross sectional area of 0.075 (0.75 × 0.10) in2. After one year, the increase in length is 0.031 in. Stress (σ ) =

W 150.00 = = 2,000 psi A 0.075

Strain (after 1 year) = ε(1 year) = ECreep (after 1 year) =

∆ 0.031 = = 0.0065 or 0.65% L 4.75

Stress σ 2,000 = = = 307,692 psi Strain(1 year) ε(1 year) 0.0065

The use of tensile strength, strain, and secant or apparent modulus in designing injection molded thermoplastic components is very important in the structural analysis of the product. The standard strength of material design equations are used to analyze the structure of the thermoplastic injection molded products. However, the stress, strain, and the modulus of elasticity chosen for the structure analysis of the product are selected on the basis of the service time and end use temperatures that the thermoplastic product is subjected to under load. Several types of curves have been developed by the plastic suppliers to analyze the behavior of the thermoplastic materials, when they are subjected to various temperatures under load for long periods of time (creep).

141

2.8 Stress-Strain Recovery (Hysteresis)

ABS

3.000

2.000

PP copolymer

1.000

HDPE

LDPE 0 0

1

5

4

Figure 2-38 Isochronous stress-strain curves for several resins under load after 1,000 hours at 73 °F air temperature Unfilled nylon 6/6 at 73˚ F. ur

ur

ho

ho

0

10

2.000

0h

ou

rs

5.00

1.500

s.

7

6

Strain, (%)

0.1

1 hr 10 h r s. 100 h r s. 1.00 0h r s. 10 .00 0h r s.

Tensile stress, (psi)

3

2

2.500

3.000

PP -impact copolymer

PP homopolymer

1.

Figure 2-40 shows three isochronous stress-strain curves of nylon 6/6 under load. The ﬁrst set of curves shows the creep behavior of unﬁlled nylon 6/6 conditioned at 50% R.H. and 73 °F. The second set of curves shows the creep behavior of 33% ﬁber glass reinforced nylon 6/6 conditioned at 50% R.H. and 73 °F. The third set of curves shows the creep behavior of 33% ﬁber glass reinforced nylon 6/6 conditioned at 50% R.H. and 140 °F. These curves are used to determine the apparent modulus, tensile stress, and strain of nylon 6/6 at various periods of time. These creep properties are used for calculating the strength of the nylon 6/6 molded products at the end of the speciﬁed service time.

PVC (pipe grade)

Tensile stress, (psi)

Figure 2-39 shows four different isochronous stress-strain curves of acetal homopolymer under load in air at 73, 113, 185, and 212 °F. These curves are used to determine the apparent modulus, tensile stress, and strain of acetal homopolymer at various periods of time. These creep properties are used for calculating the stress of the acetal homopolymer molded products at the end of the service time.

4.000

Tensile stress, (psi)

Figure 2-38 shows isochronous stress-strain curves for several thermoplastic materials under load after 1,000 hours at room temperature (73 °F). These curves are only used to compare the creep behavior among these thermoplastic materials shown in the isochronous stress-strain graph.

0h

our

s

1.500

u ho

0

00 .000 urs 5 ho 00 0 . 10

0

0.5

1.0

1.5

0

3.0

3.5

ur ho

3.000

5.

0

0.1

500

2.5

Nylon 6/6, 33% G. R. at 73˚ F.

4.000

500

2.0

Strain, (%)

5.000

1.000

1.000

500

rs

1.0 10 h o u 0 0h r ho ou ur rs s

0 1.

2.000

rs

00

1.500

u ho

Tensile stress, (psi)

2.000

2.500

1 10 hour 10 hou 0 h rs ou rs

10

0.

2.500

Tensile stress, (psi)

00

0

hr

1.000

2.000

0

0.5

1.0

1.5

2.0

2.5

0

Strain, (%) at 73˚ F. air

0.5

1.0

1.5

2.0

2.5

1.000

Strain, (%) at 113˚ F. air

ou 0h 0 1.0 ours h 00 0 5. urs o h .000

ur ho 2.000

10

1.500

5

ho

ur

s

ou 0h 0 1 s our 00 h

1.000

0

0.5

1.0

1.5

Strain, (%) at 185˚ F. air

2.0

2.5

0.6

0.8

5.000 4.000 0.

3.000

1

ho

5

ur

0 .00

ur rs ho ou 0 0h . 1 10

ho

ur

s

2.000 1.000 0

0

0

0.4

Nylon 6/6, 33% G. R. at 140˚ F.

rs

s hour 1 .0 0 0 2.000 hours 3.000 hours 5.000 hours

500

500

0.2

Strain, (%)

1

Tensile stress, (psi)

ur ho

0 10

rs

0

2.500

Tensile stress, (psi)

10

s

s ur

r

1.500

1.000

ho

ou 1h

2.000

10

Tensile stress, (psi)

0 2.500

0

0.5

1.0

1.5

Strain, (%) at 212˚ F. air

Figure 2-39 Acetal homopolymer isochronous stress-strain curves, under load at four temperatures (Courtesy: Du Pont)

2.0

2.5

0

0.2

0.4

0.6

0.8

Strain, (%)

Figure 2-40 Nylon 6/6 isochronous stressstrain curves, under load, 50% R.H. at two temperatures (Courtesy: Du Pont)

4.0

142 Tensile stress, (psi)

2 Engineering Product Design 2.600 2.000 1.600

1.0

ho

ur

ur 50 ho

Figure 2-41 shows isochronous stress-strain curves of unreinforced polycarbonate under load at 203 °F. Figure 2-42 shows an isochronous strain-time creep in ﬂexure at 4,000 psi stress curves of 30% ﬁber glass reinforced PET at 73, 140, and 257 °F. Both graphs are used to determine the apparent modulus and strain at various time periods.

s

our 100 h

s

1.000 600 0

0

0.5

1.0

1.5

2.0

Strain, (%)

Strain, (%)

Figure 2-41 Unﬁlled polycarbonate isochronous stress-strain curves, under load at 203 °F (Courtesy: General Electric Plastics)

1.8 1.6 257

1.4

Example 2-6 A container made of acetal homopolymer to store gas at 60 psi for one year or 10,000 hours. The outside diameter must not exceed 1.00 in. The maximum end use temperature is 113 °F and the maximum gas pressure is 80 psi. Determine the wall thickness and the inside diameter.

˚ F.

Solution

1.2 140˚

1.0

F.

The radial growth and mean hoop stress of a thin closed-end tube can be calculated using the following equations:

0.8

Change in radial displacement = ∆ T = (r / E T ) × (1 − υ /2) × σ h

0.6 73 ˚ F.

0.4 0.2 0 0.1

1.0

10

100

1.000

10.000

Time, (hours)

Figure 2-42 PET 30% G. R. isochronous creep ﬂexure curves, at 4,000 psi stress at three temperatures (Courtesy: Du Pont)

Where: Mean hoop stress = σh = (P × r) / t (psi) ∆T = Total radial displacement at 10,000 hours (in) r = Tube radius (in) ET = Apparent modulus at 113 °F and 10,000 hours (psi) υ = Poisson’s Ratio (0.35) P = Maximum internal pressure (80 psi) t = Tube thickness (0.031 in) Wall thickness = 0.031 in, I. D. = 0.837 in and O. D. = 0.9 in (at 113 °F) Then: r = (0.45 + 0.418) / 2 = 0.434 in σh = (P × r) / t = (80 × 0.434) / 0.031 = 1,120 psi Apparent modulus at 10,000 hours (E10,000), 113 °F and 1,120 psi can be determined from the acetal homopolymer isochronous stress-strain curve at 113 °F, a corresponding strain value = 1.05% is found in Figure 2-39 (second set of curves). E10,000, 113 °F, 1,120 psi = σh / ε10,000 = (1,120 × 80) / 1.05 = 85,333 psi ∆10,000 = (0.434 / 85,333) × (1–0.35 / 2) × (80 × 0.434) / 0.031 = 0.0047 in or on the diameter, ∆10,000 = 2 × 0.0047 = 0.0094 in O. D.10,000 = 0.900 + 0.0094 = 0.9094 in (within tolerance).

2.8 Stress-Strain Recovery (Hysteresis)

Example 2-7 Obtain a press-ﬁt of 0.625 in diameter steel shaft into an acetal homopolymer hub, which will require a pull-out force greater than 40 lb after 1,000 hours of exposure at 185 °F in air. A) Determine the hub dimensions. B) Determine the initial pull out force. C) Could the press-ﬁt be accomplished at 73 °F? Solution A) Hub dimensions at 185 °F. Assume the steel shaft size variation with temperature is negligible. The apparent modulus at 1,000 hours (E1,000), 185 °F, and 1,500 psi can be determined from the acetal stress-strain creep curve long-term at 185 °F, a corresponding strain value = 1.80% is found in Figure 2-36 (third set of curves). Assume a design stress of 750 psi (safety factor of 2.0). Assume hub wall thickness is 0.125 in, hub outside diameter is 0.875 in. Hub inside diameter is calculated by ﬁnding the interference between shaft and hub which will produce a stress of 750 psi at 185 °F after 1,000 hours. Interference calculations for press-ﬁtting plastic components used these following formulas: Interference = I = [(σd × Ds) / W] × [(W + υh) / Eh + (1 – υs) / Es] Since we assume variation of shaft size with temperature is negligible. Interference = I = [(σd × Ds) / W] × [(W + υh) / Eh] Geometry factor = W = [1 + (Ds / Dh)2] / [1 – (Ds / Dh)2] Where: I = Diameter interference (in) σd = Design stress (750 psi) Dh = Outside diameter of hub (0.875 in) Ds = Diameter of shaft (0.625 in) Eh = Tensile modulus of elasticity of hub (psi) Es = Modulus of elasticity of shaft (psi) υh = Poisson’s ratio of acetal hub (0.35) υs = Poisson’s ratio of steel shaft W = Geometry factor Then: W = [1 + (0.625 / 0.875)2] / [1 – (0.625 / 0.875)2] = 3.08 The apparent modulus at 1,000 hours (E1,000), 185 °F, and 750 psi can be determined from the acetal isochronous stress-strain at 185 °F, a corresponding strain value = 0.80% is found in Figure 2-39 (third set of curves). E500 hrs. = (750 × 100) / 0.80 = 93,750 psi

143

144

2 Engineering Product Design

Then: I = [(750 × 0.625) / 3.08] × [(3.08 + 0.35) / 93,750] = 0.00556 in Inside diameter of hub = Ds – I = 0.625–0.00556 = 0.6194 in The hub length may be calculated using the force and pressure equations to press two parts together. F = π f P Ds L, P = σd / W Where: F = Assembly force (40.0 lb) f = Coefﬁcient of friction (0.35) P = Joint pressure (psi) Ds = Diameter of shaft (0.625 in) L = Length of press-ﬁt surface (in) σd = Design stress (750 psi) W = Geometry factor (3.08) Combining the previous equations and solve for L: L = (F × W) / (π × f × Ds × σd) L = (40 × 3.08) / (3.14 × 0.35 × 0.625 × 750) = 0.239 in Hub dimensions are: L = 0.239 in, t = 0.125 in, I. D. = 0.6194 in. B) Pull-out load at start of 1,000 hours at 185 °F. By using the acetal stressstrain curve in Figure 2-29 at 1.0% strain, 185 °F, cross head speed 0.20 in per minute, tensile secant modulus = 250,000 psi. Now hub inside diameter will be 0.625 in at 185 °F in the press-ﬁt condition. ε = 0.6194/0.625 = 0.991 in/in Design stress initial = σd = E185 °F × ε185 °F = 250,000 × 0.00991 = 2,477 psi Joint pressure = P = σd / W = 2,477/3.08 = 804.22 psi Force = F = π f P Ds L = 3.14 × 0.35 × 804.22 × 0.625 × 0.239 = 132.09 lbs. C) Could the press-ﬁt be accomplished at 73 °F? I. D.185 °F = 0.6194 in before assembly? Thermal contraction from 185 to 73 °F. D73 – D185 = α × (73–185) D185 in

D73 = D185 + α × (73–185) D185 in

α = Coefﬁcient of linear thermal expansion of acetal from 73 to 185 °F.

α = 7.2 × 10–5

D73 = 0.6194 + (7.2 × 10–5) × (–112) × 0.6194 = 0.6194–0.0049 = 0.614 in

145

2.9 Flexural Beam Stress Distribution

Calculate strain for press-ﬁt of 0.614 in inside diameter hub of acetal homopolymer onto 0.625 in shaft of steel at 73 °F. Eh = σd / εh I = 0.625–0.614 = 0.011 in

εh = (I / Ds) × [W / (W + υh)] = (0.011 / 0.625) × [3.08 / (3.08 + 0.35)] εh = 0.0158 in/in = 1.58% From the acetal homopolymer stress-strain curve at 73 °F (see Figure 2-29, ﬁrst set of curves), it is seen that this strain is well below the 4.50% yield strain and the assembly could be accomplished.

2.9

Flexural Beam Stress Distribution

The difference between metal and thermoplastic design methods lies in the selection of mechanical properties to be used in standard elastic design equations. For metal materials, these properties are relatively constant over wide ranges of temperature, time under load, and other conditions such as weather and environment. For thermoplastic materials, these properties are more sensitive to variations compared to metals under the same environmental conditions. To illustrate this behavior, consider the stress distribution of a simple three-point beam in ﬂexure. When the beam is ﬂexed, two symmetrical regions of stresses in opposite directions are formed at each side of the neutral plane. With one side facing the load, the ﬁber surface stress is in compression, while on the opposite side, facing the two supports, the surface stress is in tension. It assumes that the stress-strain curves in tension and compression are symmetrical with the same characteristics; it also assumes that the modulus of elasticity only at relatively small strains is linear in the elastic region. Therefore, the stress distribution in a beam in ﬂexure is also symmetrical about the neutral plane; this linear stress distribution of a beam in ﬂexure is shown in Figure 2-43 (bottom left).

Load point Compressive forces

Tensile forces

Neutral plane Support point

σcompression

σcompression

Neutral plane

σtension Equal tensile and compressive stresses

σtension Greater compressive stress than tensile stress

Figure 2-43 Flexural stress distribution of a three-point beam

146

2 Engineering Product Design When large beam deﬂections are caused by the load, two distinct rectangular shaped regions of stresses in opposite directions are formed at each side of the neutral plane. It assumes that the material is ductile and does not fracture until the ﬁber of the thermoplastic crystalline structure surface has yielded. The neutral plane is no longer in the center of the beam, because the yield stress in compression is greater than the yield stress in tension. This behavior is illustrated in Figure 2-43 top and right bottom. The standard metal beam deﬂection equation is modiﬁed when calculating thermoplastic beam structures. The ﬂexural modulus of elasticity (E) is replaced with the apparent or secant modulus (ES) of the thermoplastic material. The apparent modulus (ES) is determined by using the thermoplastic isochronous stress-strain curve at the required temperature and service time of the product. Example 2-8 A bar molded of nylon 6/6, 33% ﬁber glass reinforced at 50% relative humidity and measuring 6.00 in length × 0.50 in width × 0.125 in thickness is to carry a center load of 5.00 lbs. The bar will rest on 2 supports on a 4.00 in span with ends free to move. Determine the maximum deﬂection at the center of the beam after 5,000 hours at 50% relative humidity and room temperature (73 °F). Solution 1. Calculate maximum ﬁber stress from stress formula. σ = (3 × W × L) / (2 × b × d2) Where: σ = Stress (psi) W = Load (5.00 lbs.) δ = Deﬂection (in) E = Flexural modulus or apparent modulus (psi) I = Moment of inertia (in4) b = Width of the bar (0.50 in) d = Thickness of the bar (0.125 in) L = Length of support span (4.00 in) σ = (3 × 5.00 × 4.0) / (2 × 0.50 × 0.1252) = 3,840 psi I = (b × d3) / 12 = (0.50 × 0.1253) / 12 = 8.138 × 10–5 in4 2. Calculate deﬂection (δ5,000 hrs.). The apparent modulus at 5,000 hours (E5,000 hrs.), 73 °F, 50% R.H., and 3,840 psi can be determined from the nylon 33% G. R. isochronous stress-strain curve (see Figure 2-37, second set of curves) as a strain of 0.68%. E5,000hrs., 3,840 psi, 73 °F = (3,840 × 100) / 0.68 = 564,705 psi. From three-point beam equation, the maximum deﬂection at the center is: δ = (W × L3) / (48 × E × I);

δ5,000hrs. = (5.00 × 43) / (48 × 564,705 × 8.13 × 10–5) = 0.145 in

147

2.10 Viscoelastic Modulus Design Method

2.10

Viscoelastic Modulus Design Method

0.16

Test results

0.14

The assumption made for the ﬂexural beam deﬂection equations is that the deﬂection distribution is initially linear and remains linear with time. The validity of this assumption was checked by running a ﬂexural beam test in the laboratory. The specimen used for the test was an acetal injection molded bar measuring 5.00 in long × 0.50 in wide × 0.125 in thick; the specimen was centrally loaded, both ends were freely supported using a 4.00 in span. The deﬂection caused by the load over a period of time was measured and recorded. Figure 2-44 shows that the deﬂection-time results are in positive agreement with the calculated and experimental values. The deﬂection-time curve was calculated by using the threepoint beam equation for maximum deﬂection; in this equation the modulus of elasticity (E) was replaced by the apparent or secant modulus (ES). Deﬂection = δ =

W × L3 48 × ES × I

The apparent or secant modulus value was obtained from the low strain tensile relaxation measurements. This apparent modulus was substituted in the equation and the deﬂection-time curve was plotted. The equations used in the calculations are applicable only if the maximum stress or corresponding initial strain is within the viscoelastic modulus boundaries. The equation for calculating the maximum stress for the beam is: Maximum stress = σ Max. =

3×W × L 2 × b × d2

The initial strain can be calculated by applying Hooke’s law, that is, by dividing the maximum stress by the initial modulus. Initial strain = εO =

σ EO

The outer ﬁber surface σ (stress) and the ε (strain) were calculated to be 1,500 psi and 0.41% for this experiment. Because 1,500 psi stress and 0.41% strain are less than stress and strain at the viscoelastic modulus for acetal homopolymer (stress = 5,125 psi at the start, decreasing the strength after one year and strain = 1.25%), a positive agreement between the calculated analysis and the test results were obtained. The deﬂection-time curves were similar in both procedures.

0.12

Deflection, (inch)

The elastic design equations that were developed for metal materials can be applied for designing injection molded thermoplastic products. The deﬂection equations are expressed in terms of two material variables, the modulus of elasticity (Young’s modulus) and Poisson’s ratio. The stress equations are only dependent on the load and geometry. These formulas can be converted to the appropriate time-dependent equations by replacing the modulus of elasticity with the apparent or secant time-dependent modulus and assuming that Poisson’s ratio is a constant.

0.10

Calculated values

0.08 0.06 0.04 0.02 0 0.1

1

10

100

Time, (hours at 73˚ F.)

Figure 2-44 Acetal homopolymer ﬂex loaded beam comparison between the calculated values and test results (Courtesy: Du Pont)

1.000

148

2 Engineering Product Design

Example 2-9 Determine the deformation of an acetal homopolymer aerosol container base; the base is a circular ﬂat plate under an internal pressure of 100.00 psi for one year at 73 °F. The slightest bulge of an initially ﬂat bottom will cause the container to rock on its base. Hence, a skirt around the base is necessary if a ﬂat bottom is intended for the container. The height of the skirt must be greater than the base deformation. In a round container the walls add stiffness to the base. The height for the skirt can be calculated by assuming that a simple vertical support from the walls helps the container to stand erect. Calculating the maximum deﬂection (δMax.) at the center of the plate requires the application of the following equation: Maximum deﬂection = δMax. =

3 × P × r 4 (5 − 4 υ − υ2 ) 16 × ES (1 year) × t 3

and the corresponding maximum stress (σ Max.) is: Maximum stress = σ Max. =

3 × P × r 2 (3 + υ) 8 × t2

Where: r = Radius of plate = 0.75 in t = Thickness of plate = 0.20 in Operating condition at 73 °F after one year (5,000 h) P = Internal pressure = 100.00 psi EO = Initial modulus = 410,000 psi ES (1 year) = Apparent modulus = 175,000 psi (see Example 2-10 procedure) υ = Poisson’s ratio = 0.35 δMax. =

3 × 100 × 0.754 (5 − 4 × 0.35 − 0.352 ) = 0.015 in 16 × 175,000 × 0.203

σ Max. =

3 ⋅ 100 × 0.752 (3 + 0.35) = 1,750 psi 8 × 0.202

ε=

1,750 410,000 = 0.0043 in/in = 0.43%

Because the calculated maximum initial stress and strain values are within the viscoelastic modulus of acetal homopolymer at room temperature (stress = 5,125 psi and strain = 1.25%), the calculated maximum deﬂection of 0.0015 in is a realistic estimated value. The height of the skirt that elevates the base should be at least 0.0015 in, so that the container will stand erect and ﬁrm.

2.10 Viscoelastic Modulus Design Method

Example 2-10 To ﬁnd the radial displacement of a thermoplastic pipe under internal pressure, it is sometimes important to calculate the radial displacement of the pipe wall thickness that is under internal pressure for a speciﬁc period of time. The radial displacement (δr) and the mean hoop stress (σH) of a thin-walled closed-end pipe may be calculated by using the following equations: Radial displacement = δR = Mean hoop stress = σ H =

r ES

υ P × r  1 −  2 t

P×r t

A pipe made of unﬁlled nylon 6/6 operates using compressed air of 150 psi pressure; the pipe is conditioned to 50% relative humidity at 73 °F for 5,000 hours. Where: r = Mean radius of pipe = 0.50 in t = Pipe wall thickness = 0.10 in P = Internal pressure = 125 psi Environmental conditions of 50% R.H. at 73 °F after 5,000 hours EO = Initial modulus of unreinforced nylon 6/6 = 175,000 psi ES (5,000 hrs) = Apparent modulus = 76,500 psi υ = Poisson’s ratio = 0.40 To determine the apparent modulus of nylon 6/6 conditioned to 50% R.H. at 73 °F after 5,000 hours, it is necessary to obtain the isochronous curves for unﬁlled nylon 6/6 of 50% R.H. at 73 °F (see Figure 2-37, ﬁrst set of curves). The apparent modulus after 5,000 hours is calculated by using the small original strain (strain = 0.85%) required to maintain the elongation within the elastic range limits. The stress at 0.85% strain is 650 psi after 5,000 hours. The apparent modulus is the ratio between the 650 psi stress divided by the 0.0085 strain, which equals 76,500 psi. 125 × 0.50   = 0.00326 in (1 − 0.20) × 0.10 

δR =

0.50 76,500

σH =

125 × 0.50 = 625 psi 0.10

ε=

σH 625 = = 0.0037 in/in = 0.37% EO 175,000

The calculation shows that the radius of the pipe would increase by 0.00326 in after 5,000 hours. The radial displacement is realistically estimated, because the initial stress-strain did not exceed the viscoelastic modulus for nylon 6/6. The initial stress, the stress, and strain after one year are the following: σO = 1,485 psi σ(5,000 hr.) = 650 psi, ε(5,000 hr.) = 0.85%

149

150

2 Engineering Product Design

2.11

Centroid, Section Area, and Moment of Inertia

Element Moment The ﬁrst moment of an element area about any axis is deﬁned as the product of an element area and the perpendicular distance between element area and axis. Figure 2-45 shows the ﬁrst moment dQx of the element da about the x-axis deﬁned as: dQx = y da, and about the y-axis the ﬁrst moment it is deﬁned as: dQy = x da.

Y x da

Area Moment The ﬁrst moment of a ﬁnite area about any axis is determined by the summation of ﬁrst moments of all elements contained in the ﬁnite area about the same axis. The ﬁrst moment is developed by means of an integral.

y X

O

Area moment = Qx = ∫ dQx

Figure 2-45

Area Centroid The centroid of an area is the point at which the area might be considered to be concentrated and still leave unchanged the ﬁrst moment of the area about any axis. For example, a thin metal plate will balance in a horizontal plane if it is supported at a point directly under its center of gravity. The centroids of some areas are obvious. In a symmetrical ﬁgure such as a circle or square, the centroid coincides with the geometric center of the ﬁgure. Therefore x indicates the x-coordinate of the centroid. The centroid of an area is deﬁned by the following equations, where A denotes the area. x = ∫ x da / A = Q y / A , y =

∫ y da / A = Q x / A

Example 2-11 Locate the centroid of a triangle shown in Figure 2-46. According to the coordinate system shown, the y-coordinate of the centroid is deﬁned by the following equation

Y s

Centroid = y =

∫ y da A

dy d

Select an element such that y is constant for all points in the element. The horizontal shaded area da of the element is y s da.

y X b

Figure 2-46

Centroid = y =

∫ y s da A

151

2.11 Centroid, Section Area, and Moment of Inertia

The product y s da represents the ﬁrst moment of the shaded element about the x-axis. s d−y From similar triangles it follows that = . Substituting s in above b d integral, d

Centroid = y =

d

=

0

bd 2

2 ( y d − y 2 ) dy 2 ∫ d 0

d d  y3   2   y2  2  d3 d3  1 d   −    = −  = Centroid = d 2 2  3  3 d   2 0  3 0  d  2  

Y

Example 2-12

dρ

The polar coordinate system shown will be a logical choice for such a contour.

y=

Locate the centroid of a semi-circle shown in Figure 2-47. dθ r

θ

The area is approximately a rectangle and is given by: A = ρ dθ dρ

Centroid = y =

y =

∫∫ 0 0

∫∫

ρ dθ dθ

0 0

y =

Figure 2-47

∫ y da ∫ da π

( ρ dθ dρ)( ρ sin θ) π r

X

ρ

The y-coordinate of centroid is deﬁned by the following equation:

π r

ρ sinθ

y =

b

∫ y d (d − y) dy

=

 ρ3  ∫  3  sin θ dθ 0 0

π

r

π

∫ 0

 ρ2   2  dθ  0 r

=

r3 sin θ dθ 3 ∫0 π

r2 dθ 2 ∫0

2r 4r (− cos θ)0π = Centroid = = 0.424 r 3π 3π Y

Example 2-13 Locate the centroid, section area, and the moment of inertia of the shaded area remaining after the semi-circle of radius 0.375 in has been removed from the semi-circle area of radius 0.500 in, as shown in Figure 2-48.

y

The y-coordinate of the centroid of the shaded area is deﬁned by the following equation: Centroid = y =

∫ y da A

X 0.375 r1

Figure 2-48

0.500 r2

152

2 Engineering Product Design

The numerator of this fraction may be evaluated by remembering that it represents the ﬁrst moment of the entire 0.500 in semi-circular area minus the ﬁrst moment of the 0.375 in semi-circular area about the x-axis. The ﬁrst moment of the 0.500 in semi-circular area about the x-axis is given by the product of its area and the vertical distance from the x-axis to the centroid of this area. Similarly, for the ﬁrst moment of the 0.375 in semi-cicular area, the location of the centroid, section area and the moment of inertia of each of these areas were found in the table of geometries and equations to have the following relationship: Centroid = Y =

4 (r23 − r13 ) (0.503 − 0.3753 ) = 0.424 = 0.280 in 2 2 3 π (r2 − r1 ) (0.502 − 0.3752 )

Area = A = 1.5708 (r22 + r12 ) = 1.5708 (0.502 + 0.3752 ) = 0.613 in 2 Moment of inertia = I = 0.1098 (r24 − r14 ) − (0.283 r22 × r12 )

I = 0.1098 (0.504 − 0.3754 ) − (0.283 × 0.502 × 0.3752 )

r2 − r1 r2 + r1

0.50 − 0.375 0.50 + 0.375

Moment of Inertia = I = 0.00328 in4

Composite Areas The moment of inertia of a composite area is the summation of the moments of inertia of the component areas making up the entire area. This eliminates the necessity for integration, if the area can be broken down into rectangles, triangles, circles, etc.; for each of these, the moment of inertia can be calculated. Parallel Axis Theorem Y

yG x1

xG G y1

The parallel axis theorem states that the moment of inertia of an area about any axis is equal to the moment of inertia about a parallel axis (through the centroid), plus the product of the area and the square of the perpendicular distance between the two axes. For the area shown in Figure 2-49, the axes xG and yG pass through the centroid of the plane area. The X and Y axes are parallel axes located at distances x1 and y1 from the centroidal axes. Let A denote the area of the selected cross section, I xG and I yG the moments of inertia about the axes through the centroid, and Ix and Iy the moment of inertia about the X and Y axes. The moments of inertia about both axes are represented by these equations:

O

Figure 2-49

X

Ix = I xG + A (y1)2 Iy = I yG + A (x1)2 Element Moment of Inertia The second moment, or moment of inertia of an element of area about any axis, is deﬁned as the product of the area of the element and the square of the perpendicular distance between the element and the axis.

153

2.11 Centroid, Section Area, and Moment of Inertia The moment of inertia dIx of the element about the x-axis is: dIx = y2 da, and about the y-axis, the moment of inertia is: dIy = x2 da. Area Moment of Inertia The second moment, or moment of inertia of a ﬁnite area about any axis, is deﬁned as the summation of moments of inertia (all elements contained in the ﬁnite area) about the same axis. The moment of inertia of the ﬁnite area about the x-axis is denoted by Ix and about the y-axis is denoted by Iy. The units of moment of inertia are the fourth power of a length (in4). Ix = ∫ d Ix =

∫ y da = ∫ x 2 da

Iy = ∫d Iy

2

Example 2-14 Determine the moment of inertia of a rectangle about an axis through the centroid and parallel to the base, as shown in Figure 2-50.

Y b

The moment of inertia I xG about the x-axis passing through the centroid is deﬁned by the following equation: I xG =

∫y

2

da

d 2

Select an element such that y is constant for all points in the element. This side illustration shows a rectangular shaded area with the following characteristics: d /2

 y3  b d3 2 = = y b y b d  3  ∫ 12   − d /2 − d /2 d /2

I xG =

dy X

G

y

d 2

Figure 2-50

Example 2-15 Determine the moment of inertia of a circle about a diameter, as shown in Figure 2-51. Y

The select shaded element of area shown in the illustration uses the polar coordinate system. The radius of the circle is r. Ix is deﬁned as I x =

∫y

2

da . dθ

But y = ρ sin θ , and da = ρ dθ dρ . 2π r

Ix =

∫ ∫ ( ρ dθ dρ) (ρ

2

0

Ix =

r4 4

sin θ) 2

0

2π

∫ 0

sin 2 θ dθ =

dρ

r 2π

Ix =

∫ 0

π r4 4

 ρ4  sin θ dθ    4 0

y

θ

X

r

ρ

2

Figure 2-51

154

2 Engineering Product Design π D4 . 64 The moment of inertia of a semicircular area about an axis coinciding with its base is half of the value of the polar moment of inertia of a solid circular area.

The diameter of the circle is D = 2 r and I x =

Ix =

1 π D4 π D4 × = = 0.0245 D 4 2 64 128

Example 2-16 Determine the moment of inertia of a circle by using the moment of inertia, area, and centroid table of equations for various cross section geometries. Table 2-2 on page 156/157 shows the following: Cross section

d

Neutral axis “y” Section area distance (in) “A” (in2) y

y=

d 2

A=

Moment of inertia “I” (in4)

π d2 π d4 = 0.7854 d 2 I = = 0.049 d 4 4 64

Example 2-17 Determine the moment of inertia of the T-section shown about a horizontal axis passing through the centroid, as shown in Figure 2-52. First, locate the centroid of the area. Then introduce the x-y coordinate system. By deﬁnition, the y-coordinate is given by the following equation: Centroid = y =

∫ y da A

The numerator of this expression represents the ﬁrst moment of the entire area about the x-axis. This may be calculated by multiplying the area of each of the three component rectangles 1, 2, and 3 by the distance from the x-axis to the centroid of the particular rectangle. (0.3) (0.15) (0.075) + (0.75) (0.15) (0.375) + (0.3) (0.15) (0.075) (0.30) (0.15) + (0.75) (0.15) + (0.30) (0.15) = 0.241 in

y = 0.30

0.30 X

0.15

1

3 2 xG

0.60

Y 0.15

Figure 2-52

x1

The centroid is located 0.241 in below the x-axis. The horizontal axis passing through this point is denoted by xG in the illustration. One technique is to calculate the moment of inertia of the entire area about the x-axis, then use the parallel axis theorem to transfer this result to the xG axis. The moment of inertia about the x-axis is found as the sum of the moments of inertia about this same axis of each of the three rectangles. I x = (1/3) (0.3) (0.15)3 + (1/3) (0.15) (0.75)3 + (1/3) (0.3) (0.15)3 = 0.0217 in 4

2.11 Centroid, Section Area, and Moment of Inertia

The parallel axis theorem is used to ﬁnd the moment of inertia of the entire T-section about the xG-axis. A = (0.15) (0.30) + (0.15) (0.75) + (0.15) (0.30) = 0.2025 in2 Ix = I xG + A (y1)2 0.0217 = I xG + 0.2025 (0.241)2 I xG = 0.0099 in4

Example 2-18 Determine the moment of inertia in Example 2-17, but use the table of equations for the T-section. Select the closest cross section from the table. Modify the geometry and equations to meet your requirements. The table below shows how to make these adjustments. Centroid (y) in, Section Area (A) in2, Moment of Inertia (I) in4

Cross section

y =1− n

s

m T

b

l h

+ 3 T l 2 − h (T − t ) (3 l − h)] y

a

A=

t

b T

l

1 [3 s 2 (b − T ) + 2 a m (m + 3 s) 6A

h (T + t ) + n T + a (s + n) 2

I =

1 3 [h (T + 3 t ) + 4 b n3 − 2 a m3 ] − A (l − y − n)2 12

y=

l 2 T + T 2 (b − T ) 2 (b T + h T )

T

A = bT + hT h

y

I =

y =

1 [T y 3 + b (l − y )3 − (b − T ) (l − y − T )3 ] 3

l 2 T + T 2 (b − T ) 0.752 × 0.15 + 0.152 (0.75 − 0.15) = = 0.241 in 2 (b T + h T ) 2 (0.75 × 0.15 + 0.60 × 0.15)

A = b × T + h × T = 0.75 × 0.15 + 0.60 × 0.15 = 0.2025 in2 I =

1 [T y 3 + b (l − y )3 − (b − T ) (l − y − T )3 ] 3

I =

1 [0.15 × 0.2413 + 0.735 (0.75 − 0.241)3 3 − (0.75 − 0.15) (0.75 − 0.241 − 0.15)3 ]

I = 0.0244 in4

155

156

Table 2-2 Moment of Inertia, Cross Section Area, and Neutral Axis Equations

Cross section

Centroid “y” (in)

Section area “A” (in2)

Moment of Inertia “I” (in4)

y a

y=

a 2

A = a2

I =

a4 12

y=

l 2

A = bl

I =

l b3 12

y=

l 2

A = bl

I =

l b3 12

y=

2 l 3

A=

lb 2

I =

b l3 36

y=

l (a + 2 b) 3 (a + b)

A=

l (a + b) 2

I =

l 3 (a2 + 4 a b + b2 ) 36 (a + b)

y=

l 2

A=

π l2 = 0.7854 l 2 4

I =

π l4 = 0.049 l 4 64

y=

D 2

A=

π (D 2 + l 2 ) = 0.7854 (D 2 + l 2 ) 4

I = 0.1098 (R24 − R14 ) −

0.283 R22 R12 (R2 − R1) R2 + R1

y=

4 (R23 − R13 ) (R 3 − R13 ) = 0.424 22 2 2 3 π (R2 − R1 ) (R2 − R12 )

A=

π (R22 + R12 ) = 1.5708 (R22 + R12 ) 2

I = 0.1098 (R24 − R14 ) −

0.283 R22 R12 (R2 − R1) R2 + R1

a

y d

b

y d b

y

d

b a

y

d b

y d

d

D

y y

r2

y

a

d

y=a b

A = π (a b − c l) = 3.1416 (a b − c l)

I =

π 3 (a b − c 3 l) = 0.7854 (a3 b − c 3 l) 4

I =

1 12

c

a n y

L

t

h b

s

d

l y= 2

A = l t + 2 a (s + n)

 3  1 (h 4 − h 4  b l − 4 G  

GFlange Slope =

h−h b−t

2 Engineering Product Design

r1

Centroid “y” (in)

Section area “A” (in2)

d

s

h

t

y

b

a

y=

I =

b 2

A = l t + 2 a (s + n)

s

a

t

I =

y h d

L

y=

n

l 2

A = l t + a (s + n)

a y

n

h b

n

L

d

s

b

  h t2 G y = b −  s b2 + + (b − t )2 (b + 2 t  ÷ A 2 3   GFlange Slope

h−h = 2 (b − t )

I = A = l t + a (s + n)

h−h b−t

 1  3 1 8 h 4 − h 4 ) b l − 12  8G  h−h 2 (b − t )

1 G 2 s b3 + h t 3 + (b 4 − t 4 )  − A (b − y )2 3  2

GFlange Slope =

h−h 2 (b − t )

m

T

y = l − [3 s 2 (b − T ) + 2 a m (m + 3 s)

a

l h

G 4  3 3 4  b (l − h) + h t + 4 (b − t )

GFlange Slope =

b

s

1 12

GFlange Slope =

n

L

Moment of Inertia “I” (in4)

+ 3 T l 2 − h (T − t ) (3 l − h)] ÷ 6 A

y

A=

h (T + t ) + n T + a (s + n) 2

I =

2.11 Centroid, Section Area, and Moment of Inertia

Cross section

1 3 [h (T + 3 t ) + 4 b n3 − 2 a m3 ] − A (l − y − n)2 12

t m s

I =

y t

y=

b T n

h l

b 2

A=

h (T + t ) + n T + a (s + n) 2

a

s b3 + m T 3 + h t 3 12 a m [2 a2 (2 a + 3 T )2 ] + 36 h (T − t )[(T − t )2 + 2 (T + 2 t )2 ] + 144

t

y

a

y=a−

a

a2 + a t − t 2 2 (2 a − t )

A = t (2 a − t )

I =

1 [t y 3 + a (a − y )3 − (a − t )(a − y − t )3 ] 3

t

t d

y=

l 2

A = l t + s (b − t )

I =

t l 3 + s 3 (b − t ) 12

y=

2a − t 2

A = t [b + 2 (a − t )]

I =

b (a + c )3 − 2 l c 3 − 6 c l a2 12

s

y

b t

t

a a

c d b

y

157

158

2 Engineering Product Design

2.12

Radius of Gyration

If the moment of inertia of an area (I) about the x-axis is denoted by Ix, then the radius of gyration Rx is deﬁned by: Rx = (Ix / A)1/2; similarly, the radius of gyration about the y-axis is given by: Ry = (Iy / A)1/2. The radius of gyration is used for comparative purposes but has no physical signiﬁcance. The frequently used geometries for thermoplastic design structures have their own equations for calculating the neutral axial or centroid distance (y), the cross section area (A), and the moment of inertia (I). Table 2-2 of typical geometries and equations was developed to provide engineering assistance for structural design analysis of injection molded thermoplastic components.

2.13

Stress Analysis of Beams

The types of loads applied to beams are either forces or couples that lie in a plane, containing the longitudinal axis of the beams. The forces are understood to act perpendicular to the longitudinal axis and the plane containing the forces is assumed to be a plane of symmetry of the beam. The effects of these forces and couples acting on a beam are: • Deﬂections perpendicular to the longitudinal axis of the beam • Setting up both normal and shearing stresses on any cross section of the beam perpendicular to its axis.

2.13.1

Types of Loads

If couples are applied to the ends of the beam and no forces act on the beam, this is termed pure bending. For example, in Figure 2-53, a portion of the beam between the two forces is subjected to pure bending. Bending produced by forces that do not form couple is called ordinary bending. A beam subject to pure bending has only normal stresses with no shearing stresses. A beam subject to ordinary bending has both normal and shearing stresses acting within the beam. Behavior of Beams

W a

W a

Assume that a beam is composed of an inﬁnite number of thin longitudinal ﬁbers. Each longitudinal ﬁber is assumed to act independently from the other ﬁbers, without lateral pressures or shearing stresses between the ﬁbers. The beam shown in Figure 2-53 will deﬂect downward and the ﬁbers in the lower part of the beam undergo elongation, while those in the upper part are shortened. These changes in the lengths of the ﬁbers cause stresses in the ﬁbers. The elongated ﬁbers have tensile stresses acting on the ﬁbers along the longitudinal axis of the beam, while the ﬁbers that are shortened are subject to compressive stresses. Neutral Surface of Beams

Figure 2-53 Beam in bending by two central forces

There always exists one surface in the beam containing ﬁbers that do not undergo any elongation or compression and is not subject to any tensile or compressive stress. This surface is called the neutral surface of the beam.

159

2.13 Stress Analysis of Beams Neutral Axis The intersection of the neutral surface with any cross section of the beam perpendicular to its longitudinal axis is called the neutral axis. All ﬁbers on one side of the neutral axis are in a state of tension, while those on the opposite side are in compression. Bending Moment The bending moment is the sum of the moments of the external forces to one side of any cross section of the beam about an axis.

2.13.2

Normal Stresses in Beams

The normal stress of a beam is the product of a longitudinal plane of symmetrical ends subject to a bending moment (M) at a certain cross section. The normal stress acting on a longitudinal ﬁber at a distance y from the neutral axis of the beam is deﬁned by the following equation: σ =

Compressive stresses

My I

where I denotes the moment of inertia of the cross sectional area about the neutral axis. Normal stresses for a beam increase from zero at the neutral axis to a maximum at the outer ﬁbers. The stresses are tensile on one side of the neutral axis, compressive on the other. These stresses are also called bending, ﬂexural, or ﬁber stresses. These stresses are shown in Figure 2-54.

Zero stress at the neutral axis Tensile stresses N. A

. y

Figure 2-54 Beam cross section, differential distribution of stresses

Example 2-19 Derive an expression for the relationship between the bending moment acting at any section in a beam and the bending stress at any point in this same section. The beam shown in Figure 2-55 is loaded by the two couples M and consequently is in static equilibrium. Since the bending moment has the same value at all points along the bar, the beam is said to be in a state of pure bending. To determine the distribution of bending stress in the beam, cut the beam by a plane passing through it in a direction perpendicular to the geometric axis of the bar. In this manner, the forces under investigation become external to the new body formed, even though they were internal effects of the original uncut body. The free body diagram of the portion of the beam to the right of this cutting plane now appears as in Figure 2-56. Evidently, a moment M must act over the cross section cut by the plane so that the left portion of the beam will be in static equilibrium. The moment M acting on the cut section represents the effect of the left portion of the beam on the right portion. Since the left portion has been removed, it must be replaced by its effect on the right portion represented by the moment M. This is the result of the moments of forces acting perpendicular to the cross section. It is now necessary to make certain assumptions to determine the nature of the variation of these forces over the cross section.

M

M

Figure 2-55

M

Figure 2-56

M

160

2 Engineering Product Design O

M

A c A

M

B d ef B

y

Figure 2-57 Beam in bending by two couples

Consider the beam to be composed of an inﬁnite number of thin longitudinal ﬁbers. It is assumed that every longitudinal ﬁber acts independently of every other ﬁber, that is, there are no lateral pressures or shearing stresses between adjacent ﬁbers. Thus, each ﬁber is subject only to axial tension or compression. Further, it is assumed that a plane section of the beam normal to its axis before loads are applied remains plane and normal to the axis after loading. It is also assumed that the material follows Hooke’s law and that the moduli of elasticity in tension and compression are equal. Next, consider two adjacent cross sections A-A and B-B on the side of the beam, as shown in Figure 2-57. Before loading, these sections are parallel to each other. After the applied moments have acted on the beam, these sections are still planes but they have rotated about each other to the positions shown, where O represents the center of curvature of the beam. Evidently, the ﬁbers on the upper surface of the beam are in a state of compression, while those on the lower surface have been extended slightly and are in tension. The line c-d is the trace of the surface in which the ﬁbers do not undergo any strain during bending; this surface is called the neutral surface. Its intersection with any cross section is called the neutral axis. The elongation of the longitudinal ﬁber at a distance y (measured positively downward) may be found by drawing line d-e parallel to A-A. If ρ denotes the radius of curvature of the bent beam, then from the similar triangles c-O-d and e-d-f we ﬁnd the strain of this ﬁber to be ε=

ef de y = = cd cO ρ

Therefore, the strains of the longitudinal ﬁbers are proportional to the distance y from the neutral axis. Because Hooke’s law holds, E = σ / ε, or σ = E × ε, it immediately follows that the stresses existing in the longitudinal ﬁbers are proportional to the distance y from the neutral axis or σ =

Ey ρ

Consider a beam of rectangular cross section, although the derivation actually holds for any cross section, that has a longitudinal plane of symmetry. In this case, these longitudinal or bending stresses are shown in Figure 2-58.

y

da

N. A

.

Figure 2-58 Beam cross section bending stresses

Let dA represent an element of area of the cross section at a distance y from the neutral axis. The stress acting on dA is given by the above expression and consequently, the force on this element is the product of the stress and the area dA. dF =

Ey dA ρ

However, the resultant longitudinal force acting over the cross section is zero (for the case of pure bending) and this condition may be expressed by the summation of all forces dF over the cross section. This is done by integration:

2.13 Stress Analysis of Beams

∫

Ey E dA = ∫ y dA = 0 ρ ρ

Evidently, ∫ y dA = 0. However, this integral represents the ﬁrst moment of the area of the cross section about the neutral axis, since y is measured from that axis. But ∫ y dA = y A, where y is the distance from the neutral axis to the centroid of the cross sectional area. From this, y A = 0 and since A is not zero, then y = 0. The neutral axis always passes through the centroid of the cross section. The moment of the elemental force dF about the neutral axis is given by dM = y dF = y dA

Ey ρ

The resultant of the moments of all such elemental forces summed over the entire cross section must be equal to the bending moment M acting at that section. M = But I = M =

E y2 ∫ ρ dA

∫y

2

dA , and the bending moment is:

EI ρ

Carefully note that this moment of inertia of the cross sectional area is computed about the axis through the centroid of the cross section. Previously we had σ =

Ey ρ

Eliminating from these last two equations, we obtain: σ =

My I

This equation gives the ﬂexural stresses in the beam. M is the bending moment at any section, I is the moment of inertia of the cross sectional area about an axis through the centroid, and y is the distance from the neutral axis to the ﬁber on which the stress (σ) acts. The value of y at the outer ﬁbers of the beam is frequently denoted by c and at these ﬁbers, the bending stresses are maximum resulting in: σ =

Mc I

161

162

2 Engineering Product Design 10.00 in

Example 2-20

75 lb

A

B

M =75 x 10 =750 in-lb

Bending moment diagram

y

1.25 in

0.187 in

Beam cross section

Figure 2-59

A 30% ﬁber glass reinforced PET cantilever beam, 10.00 in in length, is subjected to a concentrated load of 75.00 lbs. at the free end of the beam (A). The beam is of rectangular cross section, 0.187 in ×1.25 in. The tensile stress of this polymer is 23,000 psi. Determine the magnitude and location of the maximum tensile and compressive bending stresses. The bending moment diagram for this type of loading is found by using the following equation: It is triangular with a maximum ordinate at the supporting wall, as shown in Figure 2-59. The maximum bending moment (M) is merely the moment of the 75.00 lbs. force about an axis through point B and perpendicular to the plane of the page, multiplied by the 10.00 in in length. The bending stress (σ) at a distance y from the neutral axis, which passes through the centroid of the cross section, is σ = M y / I. In this equation, I is the moment of inertia of the cross sectional area about the neutral axis and is given by the following equation: I =

1 1 b h3 = × 0.187 × (1.25)3 = 0.03 in 4 12 12

In the supporting wall, where the bending moment reaches its maximum, the peak tensile stress which occurs at the upper ﬁbers of the beam is: σ =

10.00 in

M y 750 × 0.625 = = 15,625 psi I 0.03

The stress required is in tension, because the beam deﬂects downward. At the lower ﬁbers adjacent to the wall, the compressive stress is equal to 15,625 psi. The 23,000 psi tensile stress of the material provides an adequate design safety factor.

W = 50 lb/in

Example 2-21 A 45% glass reinforced PET cantilever beam, 10 in long, is subjected to a uniformly distributed load of 50 lb/in length. The allowable working stress in either tension or compression is 28,000 psi. If the cross section is to be rectangular, determine the dimensions if the height is to be four times greater than the width. M = 50 x 10 x 5 = 2.500 in-lb Bending moment diagram

A cantilever beam bending moment with a uniform load is parabolic, varying from zero at the free end to a maximum at the supporting wall. The loaded beam and the accompanying bending moment diagram are shown in Figure 2-60. The maximum moment at the wall is given by:

h = 4b

b Beam cross section

Figure 2-60

M(x = 10) = 50 × 10 × 5 = 2,500 lb-in The only cross section that needs to be considered for design purposes is the one where the bending moment is at maximum at the supporting wall. We wish to design a rectangular beam to a bending moment of 2,500 lb-in, with a maximum bending stress of 28,000 psi.

163

2.13 Stress Analysis of Beams

Because the cross section is to be rectangular, it will have the appearance shown in the diagram, where the width is denoted by b and the height by h = 4 b. The moment of inertia about the neutral axis, which passes through the centroid of the section, is given by the following equation: I =

1 1 b h3 = b (4 b)3 = 5.333 b 4 12 12

At the cross section of the beam adjacent to the supporting wall, the bending stress in the beam is given by σ = M y / I. The maximum bending stress in tension occurs along the upper surface of the beam, because these ﬁbers elongate slightly and at this surface y = 2 b and σ = 28,000 psi. Then σ =

My I

or 28,000 =

2,500 × (2 b) 5,000 = 5.333 b 4 5.333 b3

from which b = 0.322 in, and h = 4 b = 1.289 in

Location of the Neutral Axis The neutral axis always passes through the centroid of the cross section. This is the same moment of inertia (I) for the normal stress equation, or the moment of inertia of the cross sectional area about an axis through the centroid of the beam. Section Modulus At the outer ﬁbers of the beam, the value of the coordinate y is frequently denoted by the symbol c. The maximum normal stresses are deﬁned by the following equations: σ =

Mc I

or σ =

M (I / c )

The ratio (I / c) is called the section modulus and is usually denoted by the symbol Z. The maximum normal bending stresses may then be represented by the following equation: σ =

M Z

This equation is convenient, because values of section modulus (Z) are available in handbooks for a wide range of standard structural shapes. Assumptions In the derivation of the above normal stress equations, it is assumed that a beam normal to its longitudinal axis before loading remains plane after the forces and couples have been applied. It is further assumed that the beam is initially straight and of uniform cross section. Further, that the moduli of elasticity in tension and compression are equal.

164

2 Engineering Product Design

2.13.3

Shearing Force

The algebraic sum of all the vertical forces to one side of any cross section of the beam is called the shearing force at that section. Shearing Stresses in Beams When beams are subject to a shearing force (Fs) at a certain cross section, both vertical and horizontal shearing stresses (σS) are applied. The vertical shearing stresses at any cross section of a beam cause the formation of the shearing force (Fs). In the cross section of the beam shown in Figure 2-61, the vertical plane of symmetry contains the applied forces and the neutral axis passes through the centroid of the section. The coordinate y is measured from the neutral axis. The moment of inertia of the entire cross sectional area about the neutral axis is denoted by I. The shearing stress on all ﬁbers a distance y0 from the neutral axis is deﬁned by the following equation: σS =

Fs y da Ib ∫

Where b denotes the width of the beam at the location where the shearing stress is being calculated.

Example 2-22 In the case of a beam loaded by transverse forces acting perpendicular to the axis of the beam, not only are bending stresses parallel to the axis of the bar produced, but shearing stresses also act over cross sections of the beam perpendicular to the axis. Derive an expression for these shearing stresses with the shearing force at the section and the properties of the cross section. The theory to be developed applies only to a cross section of rectangular shape. However, the results of this analysis are used to give approximate values of the shearing stress in other cross sections having a plane of symmetry.

M + dM

M A c S

B d

B

A dx

σ =

My I

Where I is the moment of inertia of the cross section about the neutral axis.

c

yO

Let us consider an element of length dx cut from a beam as shown in Figure 2-61. We shall denote the bending moment at the left side of the element by M and that at the right side by (M + dM), as the bending moment changes slightly from one section to an adjacent section of the beam. If y is measured upward from the neutral axis, then the bending stress at the left section A-A is given by the following equation:

N. A.

b Figure 2-61 Element of length cut from the beam

This stress distribution is shown in Figure 2-61. Similarly, the bending stress at the right section B-B is: σ =

( M + dM ) y I

2.13 Stress Analysis of Beams

Let us now consider the equilibrium of the shaded element A-c-d-B-A. The force acting on an area dA of the face A-c is merely the product of the intensity of the force and the area; thus, σ da =

My dA I

The sum of all such forces over the left face A-c is found by integration. c

∫

y0

My dA I

Likewise, the sum of all normal forces over the right face d-B is given by: c

∫

y0

( M + dM ) dA I

Evidently, as these two integrals are unequal, some additional horizontal force must act on the shaded element to maintain equilibrium. Because the top face A-B is assumed to be free of any externally applied horizontal forces, the only remaining possibility is that there exists a horizontal shearing force along the lower face c-d. This represents the action of the lower portion of the beam on the shaded element. Let us denote the shearing stress along this face by σs as shown. Also, let b denote the width of the beam at the position where σs acts. Then the horizontal shearing force along the face c-d is σ S b dx . For equilibrium of the element A-c-d-B-A we have: c

∑ Fh =

∫

y0

My da − I

c

∫

y0

( M + dM ) y dA + σ S b dx = 0 I

Resulting in: σS =

1 dM − Ib dx

c

∫ y dA

y0

But, we have Fs = dM / dx, where Fs represents the shearing force (in pounds) at the section A-A (Figure 2-61). Substituting, σS =

Fs Ib

c

∫ y dA

y0

The integral in this last equation represents the ﬁrst moment of the shaded cross sectional area about the neutral axis of the beam. This area is always the portion of the cross section that is above the level at which the desired shear stress acts. This ﬁrst moment of area is sometimes denoted by Q in which case the above formula becomes: σS =

Fs Q Ib

165

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2 Engineering Product Design

1

m

c

dy

2

p

The shearing stress (σs) acting component is only in the horizontal direction as previously shown. However, let us consider the equilibrium of a thin element m-n-o-p of thickness t cut from any body and subject to a shearing stress σ1 on its lower face, as shown in Figure 2-62.

2

dx

n

o 1

Figure 2-62 Beam equilibrium shearing stresses

The total horizontal force on the lower face is σ1 t dx. For equilibrium of forces in the horizontal direction, an equal force, which acts in the opposite direction, must act on the upper face, hence the shear stress intensity there is also the σ1. These two forces give rise to a couple of magnitude σ1 t dx dy. The only way in which equilibrium of the element can be maintained is for another couple to act over the vertical faces. Let these faces be denoted by σ2. The total force on either vertical face moments about the center of the element we have:

∑ MC

= σ 1 t dx dy − σ 2 t dx dy = 0

or σ1 = σ 2

Thus, we have the interesting conclusion that the shearing stresses on any two perpendicular planes through a point on a body are equal. Consequently, not only are there shearing stresses (σs) acting horizontally at any point in the beam, but shearing stresses of an equal intensity also act vertically at that same point. In summary, when a beam is loaded by transverse forces, both horizontal and vertical shearing stresses arise in the beam. The vertical shearing stresses are of such magnitudes that their resultant at any cross section is exactly equal to the shearing force (Fs) at that same section. The integral c

∫ y da

y0

represents the ﬁrst moment of the shaded area of the cross section about the neutral axis. The integral always represents the ﬁrst moment about the neutral axis of that part of the cross sectional area of the beam between the horizontal plane, on which the shearing stress (σS) occurs and the outer face of the beam, the area between y0 and c (Figure 2-63). From the previous equation, the maximum shearing stress always occurs at the neutral axis of the beam, whereas the shearing stress at the outer ﬁbers is always zero. This is in contrast to the distribution of normal stress over the cross section, as that varies from zero at the neutral axis to a maximum at the outer ﬁbers. c yO

h

In a beam of rectangular cross section, the previous equation for shearing stress becomes: σS =

b

Figure 2-63 Beam cross section shearing stresses

Fs y 2 (2 I )(4 − y 02 )

Where σS denotes the shearing stress on a ﬁber at a distance y0 from the neutral axis and h denotes the depth of the beam. The distribution of vertical shearing stress over the rectangular cross section is parabolic, varying from zero at the outer ﬁbers to a maximum at the neutral axis.

167

2.13 Stress Analysis of Beams

Example 2-23 Using the expression for shearing stress derived in Example 2-22, determine the distribution of shearing stress in a beam of rectangular cross section. What is the maximum shearing stress in a rectangular bar? In Example 2-22, the shearing stress (σs) at a distance y0 from the neutral axis of the beam was found to be: σS =

Fs Ib

c

∫ y dA

y0

Where Fs denotes the shearing force at the cross section and b represents the width of the beam at the position where σs is acting. It is necessary to evaluate the above integral for a rectangular cross section. Let h denote the cross section height and b its width, as shown in Figure 2-64. The integral represents the ﬁrst moment of the shaded area about the neutral axis. Note that this area extends from the level at which the desired shearing stress (σs) works on the outer ﬁbers of the beam. In this manner, we ﬁnd the shearing stress (σs) works on all the ﬁbers at a distance y0 from the neutral axis. In fact, it is not necessary to integrate in such a simple case. Because the integral is known to represent the ﬁrst moment of the shaded area about the neutral axis, we may calculate this ﬁrst moment according to the deﬁnition. That is, the ﬁrst moment of the shaded area is simply the product of the area and the perpendicular distance between the centroid of the area and the neutral axis. The area is given by: h  A = b  − y0  2  The distance from the centroid of the shaded region to the neutral axis is: y =

1 h   + y 0  2 2

Consequently, the value of the integral representing the ﬁrst moment area is: c

∫

y dA =

y0

2  1 h  h  1 h b  + y 0   − y 0  =  − y 02  b     2 2 2 2 4 

and the shearing stress (σs) at a distance y0 from the neutral axis becomes: σS =

Fs Ib

 1  h2   Fs  h2 2 − y 02   b  − y0   =    2 I  4   2  4

From this it may be seen that the shearing stress over the cross section varies from a maximum at the neutral axis (yo = 0) to zero at the outer ﬁbers

N.A.

yO h

b

Figure 2-64 Beam integral rectangular cross section

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2 Engineering Product Design

of the beam (yo = h / 2). This variation is shown in Figure 2-65 (bottom illustration).

N. A. h

At the neutral axis, yo = 0, the maximum shearing stress is found by substitution in the above equation to be:

b

(σ S )Max. =

Fs h2 8I

( S) Max. But, for a rectangular cross section, I = (σ S )Max. =

b h3 . Substitution results in 12

Fs h2 3F  =  s  3 b h  2 b h  8   12 

Figure 2-65 Beam cross section variation

The maximum shearing stress in the case of a rectangular cross section is 50% greater than the average shearing stress obtained by dividing the shearing force by the cross sectional area b h.

2.14

Beam Deflection Analysis

In the previous section it was stated that lateral loads applied to a beam not only create bending and shearing stresses in the beam, but also cause the beam to deﬂect in a direction perpendicular to its longitudinal axis. Deformation of a beam is expressed by the deﬂection of the beam from its original unloaded position. The deﬂection is measured to the neutral surface of the deformed beam from the original neutral surface. The conﬁguration assumed by the deformed neutral surface is known as the elastic curve of the beam. Figure 2-66 represents the beam simply supported in its original state without deformation, while Figure 2-67 represents the deformation of the beam caused by the load. The displacement δ is deﬁned as the deﬂection of the beam. It is necessary to determine the deﬂection (δ) for every value of x along the beam, see Figure 2-67. This relation may be written as an equation that is known as the deﬂection curve (or elastic curve) equation of the beam.

W

Speciﬁcations for the design of beams frequently impose limitations upon the deﬂections as well as the stresses. Consequently, besides the calculation of stresses, it is essential that the product designer be able to calculate the beam deﬂection. A well designed beam must be able to carry the loads to which it will be subjected without causing undesirably large deﬂections and over stressing.

Figure 2-66 Beam simply supported

Y

Numerous methods are available for calculating the deﬂection of a beam. The most common methods are the following:

W

x

X

δ Figure 2-67 Deﬂected beam

• Double integration method • Moment area method • Superposition method

169

2.14 Beam Deﬂection Analysis

2.14.1

Beam Deflection by Double Integration Method

The following differential equation is used for calculating the bending moment of a bent beam: M =EI

dy 2 dx 2

(2-1)

Where X and Y are the coordinates of a deformed beam as shown in Figure 2-67, and where δ is the deﬂection of the beam (this expression is derived in Example 2-24). Here, E denotes the modulus of elasticity of the beam and I represents the moment of inertia of the beam cross section about the neutral axis, that passes through the centroid of the cross section. M represents the bending moment at the distance x from one end of the beam. This quantity is the algebraic sum of moments, caused by external forces. These moments are formed on one side of the section at a distance x from the end about an axis through this section. The moment M is a function of x and it is necessary to integrate Equation (2-1) twice to obtain an algebraic equation expressing the deﬂection (δ) as a function of x. Equation (2-1) is the basic differential equation that governs the deﬂection of all beams, irrespective of the type of applied loading. Double Integration Procedure The double integration method for calculating deﬂections of beams consists of integrating Equation (2-1). The ﬁrst integration yields the slope dy / dx at any point in the beam and the second integration gives the deﬂection (δ) for any value of x. The bending moment (M) must be expressed as a function of the coordinate X before the equation can be integrated. Since the differential Equation (2-1) is of the second order, its solution must contain two constants of integration. These two constants must be evaluated from known conditions concerning the slope or deﬂection at certain points in the beam. For example, in the case of a cantilever beam, the constants would be determined from the conditions of zero change of slope as well as zero deﬂection at the wall support end of the beam. Frequently, two or more equations are necessary to describe the bending moment in the various regions along the length of a beam. For such cases, Equation (2-1) must be written for each region of the beam and integration of these equations yields two constants of integration for each region. These constants must then be determined to impose conditions of continuous deformations and slopes at the points common to adjacent regions. Sign Conventions The sign conventions for bending moment adopted previously (stresses in beam’s section) will be retained for calculating the deﬂection of beams. The values E and I appearing in Equation (2-1) are positive. From this equation, if M is positive for a certain value of x, then dy2 / dx2 is also positive. With the above sign convention for bending moments, it is necessary to consider the coordinate x along the length of the beam to be positive to the right and the deﬂection (δ) to be positive upward. This will be explained in detail in Example 2-24. With these algebraic signs, the integration of Equation (2-1) may be carried out to yield the deﬂection (δ) as a function of x, with the understanding that upward beam deﬂections are positive and downward deﬂections negative.

170

2 Engineering Product Design Assumptions and Limitations In the derivation of Equation (2-1) it is assumed that deﬂections caused by shearing action are negligible compared to those caused by bending action. Also, it is assumed that the deﬂections are small compared to the cross sectional dimensions of the beam. Further, the beam is presumed to be straight before the application of the load (Figure 2-66). These conditions are in addition to the assumptions concerning beam theory.

Example 2-24 Obtain the differential equation of the deﬂection curve of a beam loaded by lateral forces. The bending moment is deﬁned by the following equation: M =

EI ρ

(2-2)

In this equation, M denotes the bending moment acting at a particular cross section of the beam, ρ the radius of curvature to the neutral surface of the beam at this same section, E the modulus of elasticity, and I the moment of inertia of the cross sectional area about the neutral axis passing through the centroid of the cross section. We will be concerned only with those beams for which E and I are constant along the entire length of the beam, but in general, both M and ρ will be functions of x. Equation (2-1) may be written in the following form: 1 M = ρ EI

(2-3)

Where 1 / ρ in Equation (2-3) represents the curvature of the neutral surface of the beam. As M will vary along the length of the beam, the deﬂection curve will be of variable curvature. The heavy line shown in Figure 2-68 represents the deformed neutral surface of the bent beam. Originally, the beam coincided with the x-axis before loading and the coordinate system that is found to be most convenient is shown in Figure 2-66. The deﬂection δ is taken to be positive in the upward direction; hence for the particular beam shown, all deﬂections are negative. An expression for the curvature at any point along the deformed beam is readily available from differential calculus. The exact formula for curvature is: Y

ρ X

δ Figure 2-68 Deﬂected beam, simply supported

 dy 2   dx 2   

1 = 2 3/2 ρ   dy   + 1     dx  

(2-4)

171

2.14 Beam Deﬂection Analysis

In this equation, dy / dx represents the slope of the curve at any point and for small beam deﬂections, this quantity, and in particular its squares, are small in comparison to unity and may reasonably be neglected. This assumption of small deﬂections simpliﬁes the expression for curvature to the following form: 1 dy 2 ≈ ρ dx 2

(2-5)

For small deﬂections, Equation (2-3) becomes: dy 2 M = 2 E I dx

or M =EI

dy 2 dx 2

(2-6)

This is the differential equation of the deﬂection curve of a beam loaded by lateral forces. In solving any problem, it is necessary to integrate this equation to obtain an algebraic relationship between the deﬂection δ and the coordinate X along the length of the beam.

Example 2-25 Determine the deﬂection at every point of the cantilever beam subject to the single concentrated load W at the free end, as shown in Figure 2-69. The deformed beam is shown by the heavy line. First, it is necessary to ﬁnd the reactions exerted by the supporting wall upon the beam. These are easily found from statics to be vertical load reaction W and the bending moment M = W × L. The bending moment at any cross section a distance x from the wall is given by this assumption: It is the sum of the moments of these two reactions about an axis through this section. Evidently, the upward load W produces a positive bending moment W x and the couple W L if acting alone would produce curvature of the beam. According to the sign convention, this constitutes negative bending. Hence, the bending moment M at the section x is: M = −W L + W x Y

W

The differential equation of the bent beam is: M =EI

dy 2 dx 2

L

x M=WxL

X W

Figure 2-69 Cantilever beam, free end, concentrated load

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2 Engineering Product Design

Where E denotes the modulus of elasticity of the material and I represents the moment of inertia of the cross section about the neutral axis. Substituting, EI

d2 y dx 2

= −W L + W x

(2-7)

Equation (2-7) is readily integrated once to yield the following: EI

dy W x2 = −W L + + C1 dx 2

(2-8)

Which represents the equation of the slope, where C1 denotes a constant of integration. This constant may be evaluated by use of the condition that the slope dy / dx of the beam at the wall is zero, because the beam is rigidly clamped. Therefore, (dy / dx)x = 0 = 0. Equation (2-8) is true for all values of x and y, and if the conditions at x = 0 is substituted, we obtain: 0 = 0 + 0 + C1, or C1 = 0. Next, integration of Equation (2-8) yields EIδ= −

W L x2 W x3 + + C2 2 6

(2-9)

Where C2 is a second constant of integration. Again, the condition at the supporting wall will determine this constant. There, at x = 0, the deﬂection (δ) is zero, because the bar is rigidly clamped. Substituting (δ) at x = 0, the condition is equal to zero, in Equation (2-9), we ﬁnd that 0 = 0 + 0 + C2, or C2 = 0. Thus, Equations (2-8) and (2-9) with C1 = C2 = 0 give the slope dy / dx and deﬂection (δ) at any point x in the beam. The deﬂection is a maximum at the right end of the beam (x = L) under the load W and from Equation (2-9) is found to be: E I δ Max. = −

W L3 3

Where the negative value denotes that this point on the deﬂection curve lies below the X-axis. If only the magnitude of the maximum deﬂection at x = L is desired, it is usually denoted by deﬂection (δ) and we have the following equation: δMax. = −

W L3 3E I

(2-10)

173

2.14 Beam Deﬂection Analysis 5.00 in

Example 2-26 The cantilever beam as shown in Figure 2-69 is 5.00 in in length and with a load force W of 10.00 lb. The beam is made of acetal homopolymer with a rectangular section having 0.125 in width and 0.75 in height. The resin has a tensile strength of 10,000 psi and a modulus of elasticity of 410,000 psi at 73 °F. Determine the maximum deﬂection of the beam. The maximum deﬂection occurs at the free end of the beam under the concentrated load as shown in Figure 2-70. This deﬂection is downward as shown in Figure 2-70. In the derivation of this deﬂection equation it was assumed that the material of the beam follows Hooke’s law. In fact, from the above calculation alone there is no assurance that the material is not stressed beyond the proportional limit. If it were, then the basic beam bending equation E I (dy2 / dx2) = M would no longer be valid and the above numerical value would be meaningless. Consequently, in every problem involving beam deﬂections it is to be emphasized that it is necessary to determine that the maximum bending stress in the beam is below the proportional limit of the material. The moment of inertia of a rectangular cross section area is calculated by using the following equation: I =

1 1 b h3 = × 0.125 × (0.75)3 = 0.00439 in 4 12 12

Where σ denotes the bending stress, M the bending moment, y the distance from the neutral axis to the outer ﬁbers of the beam, and I the moment of inertia of the beam cross section about the neutral axis. The maximum bending moment in this problem occurs at the supporting wall and is given by the following equation: M(Max.) = 10.00 × 5.00 = 50.00 lb-in Substituting in the deﬂection and bending stress equations, we have: δMax. =

10.00 × (5.00)3 W L3 = = 0.231 in 3E I 3 × 410,000 × 0.00439

σ Max. =

M y 50.00 × 0.375 = = 4,271 psi I 0.00439

Because the maximum stress is below the acetal homopolymer proportional limit (10,000 psi), the use of the beam deﬂection equation was justiﬁable.

M = 50 lb-in

Max.

10 lb

0.375 in 0.75 in

0.125 in

Figure 2-70 Cantilever beam, free end, concentrated load

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2 Engineering Product Design

Example 2-27

W = 50.00 lb/in

A cantilever beam, as shown in Figure 2-71, is made of 43% ﬁber glass reinforced nylon 6/6 at 73 °F with 50% relative humidity. It has a tubular cross section, 2.00 in outside diameter and 1.50 in inside diameter. The tube is 12.00 in long and carries a uniform load of 5.00 lb/in. The material has a tensile strength of 21,000 psi and a modulus of elasticity of 1,200,000 psi.

Max.

12.00 in r = 1.00 in

Calculate the maximum deﬂection of the beam. The maximum deﬂection of the cantilever beam is uniformly distributed, the load occurs at the free end and is calculated by using the following equations:

2.00 in dia

I = 0.049 (D 4 − d 4 ) = 0.049 (2.004 − 1.504 ) = 0.535 in 4

1.50 in dia

It is important to retain consistent units when substituting in such an equation. One manner of doing this is to take W in units of lb/in, L in in, E in psi, and I in in4. The maximum deﬂection is calculated by using the following equations:

Figure 2-71 Cantilever beam, uniformly distributed load

δMax. =

50.00 × (12.00)4 w L4 = = 0.201 in 8 E I 8 × 1,200,000 × 0.535

The maximum bending moment occurs at the supporting wall and is: M Max. = 50.00 × 12.00 × 6.00 = 3,600 lb-in. The maximum bending stress occurs at the outer ﬁbers of the beam at this section adjacent to the wall and is given by the following equation: σ Max. =

M y 3,600 × 1.00 = = 6,728.97 psi I 0.535

As this maximum bending stress is well below the proportional limit of the material (21,000 psi), the use of the above deﬂection equation is valid.

Example 2-28 Obtain an equation for the deﬂection curve of the simply supported beam, subject to the concentrated load W applied at the center of the beam as shown in Figure 2-72. Y

W x X

W 2

L 2

L 2

W 2

The X-Y coordinate system is introduced. The beam deformed by the load is indicated by the heavy line. Because of symmetry, each end reaction force is W / 2. Use the equation for the bending moment at any section of a loaded beam. According to that equation, the bending moment in the left half of the beam is given by the following equation:

L Figure 2-72 Beam, simply supported, concentrated center load

M =

W L x for 0 < x < 2 2

175

2.14 Beam Deﬂection Analysis

The differential equation of the bent beam is M = E I EI

dy 2 . Substituting, dx 2

dy 2 W L = x for 0 < x < 2 dx 2 2

(2-11)

The ﬁrst integration of Equation (2-11) yields the following: EI

dy W ( x 2 ) = + C1 dx 2 (2)

(2-12)

The slope of the beam is represented by dy / dx. Because the beam is loaded at its midpoint, the deﬂections are symmetric about the center of the beam (x = L / 2). This condition of symmetry tells us that the slope must be zero at x = L / 2, the tangent to the deﬂected beam is horizontal there. Substituting this condition: (dy /dx)x = L /2 = 0 , in Equation (2-12), we obtain: 0=

W L2 W (L2 ) + C1 or C1 = − 16 4 (4)

The slope dy / dx at any point in the beam is given by the following equation: EI

dy W (x 2 ) W L2 = − dx 4 16

(2-13)

By integrating the second time, we develop the following equation: EIδ=

W (x 3 ) W L2 − x + C2 4 (3) 16

(2-14)

The second constant of integration C2 is determined by making the value of the deﬂection δ at the left support of the beam to be equal to zero. Thus δx = 0 = 0. Substituting in Equation (2-14), we obtain: 0 = 0 – 0 + C2 or C2 = 0. The deﬂection curve of the left half of the beam is given by the following equation: EIδ=

W (x 3 ) W L2 − x 12 16

(2-15)

At this point it is to be carefully noted that it is not permissible to make use of the condition that the deﬂection δ is zero at the right support, δx = L = 0. This is because the bending moment equation, M = (W / 2) x, is valid only for values of x less than L / 2 (to the left of the applied load W). To the right of W, the bending moment equation contains one additional term. It would

176

2 Engineering Product Design

be necessary to work with the bending moment equation in the right half of the beam if the condition δx = L = 0 were to be used. In fact, there is no need to examine deﬂections to the right of the load, because it is known that the deﬂection curve of the beam is symmetric about x = L / 2. In determining constants of integration it is permissible to use only those conditions on deﬂection or slope that pertain to the interval of the beam for which the bending moment equation was written. Evidently, the maximum deﬂection of the beam occurs at the center because of symmetry. At this center point, the maximum deﬂection multiplied by the modulus of elasticity and moment of inertia is deﬁned by the following equation: E I δMax. = −

W L3 48

(2-16)

The maximum deﬂection at the center of the beam (without considering the algebraic sign), simply supported at each end of the beam, and subject to a centrally applied concentrated load W, the following equation is applied: δMax. = −

W L3 48 E I

(2-17)

Example 2-29 Determine the deﬂection equation for a simply supported beam loaded by a couple M1 at the right end of the beam as shown in Figure 2-73. First, it is necessary to determine the reactions acting on the beam. Because the applied couple M1 can be held in equilibrium only by the action of another couple, the end reactions must be forces of equal magnitude R, but opposite in direction. To ﬁnd their magnitude we may write the statics equation:

∑ Mo

Y x

M1 X

= − M1 + R L = 0 or R =

The heavy line indicates the conﬁguration of the deﬂected beam. The bending moment at any section a distance x from the left reaction is: M = Rx =

Y x

M1 X R

R

M1 x L

This equation is valid for all values of x. The differential equation of the deformed beam is:

L Figure 2-73 Beam, simply supported, couple loaded at one end

M1 L

EI

dy 2 M = 1 x L dx 2

(2-18)

177

2.14 Beam Deﬂection Analysis

Integrating once, we obtain: EI

dy M (x 2 ) = 1 + C1 dx L (2)

(2-19)

There is no information available concerning the slope of the beam, making it impossible to determine C1 at this stage. It is to be noted that if there is no symmetry to the loading, then there is no reason to expect the slope to be zero at the midpoint of the beam. We integrate again and obtain: EIδ=

M1 ( x 3 ) + C1 x + C2 2 L (3)

(2-20)

At this stage, we are able to determine the constants of integration C1 and C2. It is evident that the deﬂection δ is zero at the left support, δx = 0 = 0. Substituting these values of x in Equation (2-20), we obtain 0 = 0 + 0 + C2 or C2 = 0. Also, the deﬂection δ is zero at the right support, Bdx = L = 0. Substituting these values of x and δ in Equation (2-20), we ﬁnd: 0=

M1 (L3 ) + C1 L or C = − M1 L 1 6L 6

(2-21)

The deﬂection curve of the beam becomes the following: EIδ=

M1 ( x 3 ) M1 L − x 6L 6

(2-22)

The maximum deﬂection of the beam occurs at the point where the slope is zero, at that point where the tangent to the deﬂection curve is horizontal. The coordinate X of this point is found by setting the left side of Equation (2-19) equal to zero. Substituting the equation we obtain the following: 0=

M1 ( x 2 ) M1 L − 2L 6

or x =

L 3

(2-23)

The maximum deﬂection of the beam thus occurs at a distance L / 1.732 from the left reaction. The value of this deﬂection is found by substituting x = L / 1.732 in Equation (2-22). This substitution yields the following equations: E I δMax. =

δMax. =

M1 (L3 ) 18 L 3

−

0.0064 M1 L2 EI

M1 L2 6 3

=

M1 L2 3 27

(2-24)

(2-25)

178

2 Engineering Product Design

tan

at B

B

A

tan

2.14.2

Beam Deflection Moment Area Method

The second beam deﬂection method, known as the moment area method, is considered an alternative procedure to the double integration method.

at A

First Moment Area Theorem dx

x

Bending moment diagram

Figure 2-74 Beam deﬂection and bending moment

Figure 2-74 shows that AB represents a portion of the deﬂection curve of a bent beam. The shaded area represents the bending moment diagram. On the deﬂection curve are drawn tangents at each of the points A and B as indicated. The ﬁrst moment area theorem states that: The angle between the tangents at A and B is equal to the area of the bending moment diagram between these two points, divided by the product E I. If θ denotes the angle between the tangents as shown in the previous illustration, then this theorem may be stated in equation form as follows: θ=

B

M dx EI A

∫

In this equation, E represents the modulus of elasticity of the beam and I denotes the moment of inertia of the beam cross section about the neutral axis, which passes through the centroid of the cross section. M represents the bending moment at the distance x from the point B. Second Moment Area Theorem Let us consider the vertical distance between point B on the deﬂection curve shown above and the tangent to this curve drawn at point A. This vertical distance is denoted by δ in the illustration above. The second moment area theorem states that: The vertical distance of point B on a deﬂection curve from the tangent drawn to the curve at A is equal to the moment about the vertical through B (shaded area of the bending moment diagram), between A and B, divided by the product E I. This theorem is represented in an equation form as follows: δ=

B

M x dx EI A

∫

Moment Area Procedure The determination of the deﬂection of a speciﬁed point on a loaded beam is made according to the following procedure: • The load reactions of the beam are determined; however, for cantilever beams this step is omitted. • An approximate deﬂection curve is drawn. This curve must be consistent with the known conditions at the ends of the support, such as zero slope or zero deﬂection. • The bending moment diagram is drawn for the beam. The moment diagram is constructed by parts. The M / (E I) diagram must be used in connection

179

2.14 Beam Deﬂection Analysis with either of the above theorems. However, for beams of constant cross section, the M / (E I) diagram has the same shape as the ordinary bending moment diagram, except that each ordinate is divided by the product E I. For applications of beams with constant cross section, it is possible to work directly with the bending moment diagram and then divide the computed areas or moment areas by E I. Or, equivalently, the angles or deﬂections may be multiplied by E I when areas or moment areas of the ordinary moment diagram are used. • Convenient points A and B are selected and a tangent is drawn to the assumed deﬂection curve at one of these points. • The deﬂection of point B from the tangent at A is then calculated by the second moment area theorem. In certain simple cases, such as cantilever beams, this deﬂection of B from the tangent at A may actually be the desired deﬂection. However, in other cases, it will be necessary to apply the second moment area theorem to another point in the beam and then examine the geometric relationship between these two calculated deﬂections to obtain the desired deﬂection.

2.14.3

Applications of Moment Area and Double Integration Methods

If the deﬂection of only a single point of a cantilever beam is desired, the moment area method is usually more convenient than the double integration method. On the other hand, if the equation of the deﬂection curve of the entire beam is desired, there is no procedure superior to the double integration method. Example 2-30 The cantilever beam shown in Figure 2-75 is subject to the concentrated load W applied at the free end of the beam. Determine the deﬂection under the point of application of the load. In the case of a cantilever beam, the reactions at the wall need not be determined although their determination is simple. It is known that the slope and deﬂection at the clamped end A are each zero by deﬁnition of a cantilever beam. The heavy curved line represents a realistic deﬂection curve. Next, a tangent to the deﬂection curve is drawn at point A. In the case of a cantilever beam, this tangent coincides with the original unbent position of the bar and is represented by the straight dotted line. The deﬂection of point B from the tangent at A is the actual desired deﬂection. The deﬂection of the free end of the beam, B, may now be found by use of the second moment area theorem. By this theorem, the deﬂection of point B from the tangent drawn at A is given. The moment about the vertical line through B of the area under the bending moment diagram between A and B is divided by the product E I. In fact, because the cross section of the beam is constant along the length of the beam, it is easier to work directly with the ordinary bending moment diagram rather than with the M / (E I) diagram. Therefore, the resulting deﬂection must be multiplied by the product E I.

W A

tan at A

B L X

L M Max. = W x L Bending moment diagram

With the second moment area theorem, the product (E I) times the deﬂection of B from the tangent at A, can be calculated. By the moment of the shaded

Figure 2-75 Cantilever beam, free end, concentrated load

180

2 Engineering Product Design

moment diagram about a vertical line through B, this moment (M) of area may be calculated by multiplying the area by the distance of the centroid of the area from the vertical line through B. The area of the triangular moment diagram is (L / 2) (–W L), where the negative sign is used because the bending moment is negative. The centroid of the moment diagram lies at a distance (2 L / 3) from the right end. Therefore, the moment area theorem becomes: EIδ=

W L3 1 2  (L)(−W L)  L  = − 3  2 3

or δ = −

W L3 3E I

Example 2-31 A cantilever beam is subject to the uniformly distributed load acting over the entire length of the beam as shown in Figure 2-76. Determine the deﬂection of the free end of the beam. The reactions at the end of the beam need not be determined. First, it is necessary to sketch an approximate deﬂection curve for the bent beam. Because the beam is clamped at the right end, evidently the slope and also the deﬂections at that end are each zero. The curved beam represents a deﬂection curve that agrees with the known conditions of zero slope and zero deﬂection at the right end of the beam. The maximum bending moment (M) occurs at the supporting wall and has the value (W L2) / 2, where W represents the intensity of the uniform load per unit length of the beam. In this example, the bending moment diagram is a parabola.

B

(W ) lb /u

n it le n g

th

A tan at A

L

X

L WL 2 Bending moment diagram

MMax. =

Figure 2-76 Cantilever beam, uniformly distributed load

2

Next, a tangent to the deﬂection curve is drawn at point A. The free end of the beam is designated as B. This tangent coincides with the original position of the beam and is represented by the straight dotted line. Therefore, the deﬂection of point B from the tangent drawn at A represents the desired deﬂection. The deﬂection of point B may now be found by use of the second moment area theorem. From this theorem, the deﬂection of point B from the tangent drawn at A is given. The moment about the vertical line through B of the area under the bending moment diagram between A and B is divided by E I. It is convenient to work with the bending moment diagram as shown in Figure 2-76 and then multiply the resulting deﬂection by the product E I. Therefore, by the second moment area theorem, E I times the deﬂection of B from the tangent at A, is given. By the moment of the shaded moment diagram about a vertical line through B, this moment of area may be calculated by multiplying the area by the distance between the centroid of the area and the vertical line through B. The area of the parabolic moment diagram is 1/3 the area of the rectangle enclosing the parabola. Therefore, the area under the moment diagram is given by the following equation: M =

1  W L2  L − 3  2 

181

2.14 Beam Deﬂection Analysis

The centroid of the parabolic ﬁgure lies at a distance 3 L / 4 from the left end. Therefore, the second moment area theorem becomes: EIδ=

 W L2   3 1 (L)  −  3 2   4 

W L4  L = −  8

or δ = −

W L4 8EI

The negative sign indicates that the ﬁnal position of point B lies below the tangent drawn at point A.

Example 2-32 The simply supported beam has a concentrated loadapplied at its midpoint as shown in Figure 2-77. Find the maximum deﬂection at the center of the beam. The reaction forces at the ends of the beam are each W / 2 by symmetry. The heavy line represents the deﬂection curve of the beam. Because of this symmetry, the tangent drawn at the midpoint of the deﬂected beam will be horizontal. The midpoint of the beam is denoted as point A and the tangent drawn at A is shown by the horizontal line. The left end of the beam is designated as point B. We want to ﬁnd the deﬂection of the beam at the midpoint, under the concentrated load W. Inspection of the bending moment diagram reveals, that this central deﬂection A, is identical with the deﬂection of point B from the tangent drawn at A. In calculating the deﬂection of B from the tangent at A, it is necessary to evaluate the moment. This is the area under the moment diagram between these two points about a vertical line through B; divide this quantity by the product E I. The area to be considered is the left half of the above shaded moment diagram. Look at the triangle of altitude (W L) / 4 and base L / 2. The distance from the centroid of this triangle to the vertical line through B is (2 / 3) × (L / 2), Therefore, the second moment area theorem applied between the points A and B gives the desired deﬂection:

L 2

W

L 2

B

EIδ=

tan at A

3

WL 1  L  W L   2 L      ×  or δ = 2  2   4  3 2  48 E I

A W 2

W 2

The deﬂection is positive, because point B lies above the tangent at point A. The second moment area theorem always indicates relative deﬂections: the deﬂection of one point on a beam about a tangent drawn at a second point on the beam. In this example, the displacement of B is zero. This is true, when we compare the deﬂection of B relative to the tangent drawn at A. Fortunately, this relative displacement is equal to the maximum displacement because of symmetry.

X L 2

L 2

MMax. =

L Bending moment diagram

Figure 2-77 Beam simply supported, concentrated load in middle

WL 4

182

2 Engineering Product Design

a

b

Example 2-33

W B

D C

Wa L

Wb L

L 2 L

Figure 2-78

H

Wb 2

Wb B

D

A

Figure 2-79

By calculating the reaction of force at each end of the beam, we have the values indicated in Figure 2-75. The approximate form of the deﬂection curve is indicated by the heavy line. Evidently, there is no symmetry of the deﬂection curve about the midpoint of the beam. The right end of the beam is designated as point A, the midpoint as B, the point of application of the load (W) as C, and left end as D. It is convenient to draw the bending moment by parts. Working from right to left along the length of the beam, the moment due to the right reaction alone at any section a distance x from the right end is given by (W b x) / L. This is represented by the triangle shown in Figure 2-79. As we proceed from right to left along the beam, the effect of the downward load W on the moment diagram does not become evident until we have passed to the left of its point of application. At any section a distance x from the point A and lying to the left of load W, the bending moment due to the load W alone is given by –W b. This may be represented by the triangle shown in Figure 2-80. The bending moment diagram drawn by sections is shown in Figure 2-82.

D C

The desired deﬂection at the midpoint of the beam may be found by the following rather indirect use of the moment area theorem. A tangent to the deﬂection curve is drawn at the right end of the beam. This is designated as the tangent at point A, as shown in Figure 2-81. It is now possible to calculate the deﬂection of point D from the tangent at A by use of the second moment area theorem. This deﬂection is designated as D d in Figure 2-81. Remember that the moment area theorem provides only the relative deﬂections. In this example, the deﬂection of point D is relative to the tangent drawn at A. The absolute deﬂection of point D is zero, because the beam is supported at that point.

Wb

G

Figure 2-80

W D

B e tan at

A simply supported beam has a concentrated load as shown in Figure 2-78. Determine the deﬂection at the midpoint of the beam.

A

A

A

f

d Figure 2-81

According to the second moment area theorem, the deﬂection of D from the tangent drawn at A is given by the moment of the area under the moment diagram between A and D about a vertical line through D, divided by E I, where E is the modulus of elasticity and I is the moment of inertia. Therefore, taking the moment of the above triangle ADH and CDG about the vertical line through D, we obtain the following equation:

E I (D d ) = Total bending moment diagram H Wb D

C

f

Wb 2

B

A

As stated previously, B represents the midpoint of the beam. Evidently, from similar triangles, the line segment B f shown in Figure 2-81 must be exactly half the length of D d and the following equation is obtained:

Wb G

Figure 2-82

2 W b3 1 L 1 b  W b L (L)(W b)   + (b)(−W b)   = − 3 2 3 2 6 6

E I (B f ) =

W b L2 W b3 − 12 12

183

2.14 Beam Deﬂection Analysis

Next, it is possible to calculate the deﬂection of the midpoint of the beam from the tangent drawn at A. This deﬂection is represented by the line segment e f in the previous ﬁgure. According to the second moment area theorem, this is given by the moment of the area under the bending moment diagram between A and B about a vertical line through B, divided by E I. This portion of the moment diagram is represented by the triangle ABF. Applying the second moment area theorem, we obtain the following equation: E I (e f ) =

1  L   W b   1 L  W b L2    × = 2  2   2  3 2  48

In the previous equation of the deﬂection curve of the beam, apparently, the midpoint deﬂection is represented by the line segment B e in Figure 2-81. This may be found from the following relationship: Be=Bf–ef Substituting the above values in the right side of the equation, we ﬁnd the desired deﬂection of the midpoint.

E I (B e) =

W b L2 W b3 W b L2 − − 12 12 48

or B e =

W b L2 48 E I

 4 b2  3 − 2  L  

We should recognize that this is not the maximum deﬂection of the beam, except when a = b = L / 2. Also, it is assumed that the load W lies to the left of the midpoint of the beam. Otherwise the triangular bending moment diagram CDG in Figure 2-82 would extend to the right of the midpoint of the beam. Therefore, it would be necessary to take a portion of the deﬂection into account in calculating the deﬂection e f in Figure 2-82.

2.14.4

Beam Deflection Superposition Method

The superposition method is a process for calculating the deﬂections of a beam. This method calculates the deﬂection at any speciﬁc required location of a beam by adding together the partial effects of deﬂection. The speciﬁc equation required to calculate each partial deﬂection of the beam is based on the type of loads acting on the beam structure. The equations used in this process are the typical strength of materials design formulas. They were developed by Stephen Timoshenko, James Gere, Joseph Shigley, Fred Seely, James Smith, Raymond Roark and Warren Young, and William Nash just to name a few. These equations are found in the beam deﬂection equation tables in several engineering publications, such as the Machinery’s Handbook, plastic supplier’s design manuals. The basic principle of the superposition method assumes the following: If a beam is subjected to several loads, the bending moment “M” at any cross section will be equal to the sum of the bending moments M1, M2, M3, …, etc. produced by each load separately. M = M1 + M2 + M3 + …

184

2 Engineering Product Design The bending moment is deﬁned by the following equation: M =EI

dy 2 dx 2

Therefore, the sum of the partial bending moments is obtained by the following equation: EI

dy32 dy 2 dy12 dy22 = E I + E I + E I +… dx 2 dx 2 dx 2 dx 2

After integrating the partial bending moments twice, the sum of the partial deﬂection is deﬁned by the following equation: y = y1 + y 2 + y 3 + … or δ = δ1 + δ2 + δ3 + … This equation represents the total deﬂection at any point of the beam. The deﬂection can be deﬁned as the sum of deﬂections caused by the separate load effects on a beam.

Example 2-34 Determine the free end deﬂection of a cantilever beam subjected to two types of loads using the sum of deﬂections superposition diagrams as shown in Figure 2-83. L

L

L W

w (lb/in)

W

w (lb/in)

4

wL 8EI

Total

Compound loads

3

=

WL 3EI

1

Equally distributed load Total

=

+ 1

2

=

2

Concentrated end load 4

Total deflection =

=

wL 8EI

3

+

WL 3EI

Figure 2-83 Beam compound loads deﬂection superposition diagram

The total end deﬂection of a cantilever beam is subjected to two types of loads. One load is a uniformly distributed load and the other a concentrated load at the end of the cantilever beam. Both of these deﬂection equations are found in Table 2.2

Example 2-35 Determine the deﬂection at point “A” of a structure with two cantilever beams having the same cross section and material but of different lengths. Both beams are connected by a free support at point “B”. The upper beam is loaded at the middle by a concentrated load. The beam structure is shown in Figure 2-84.

185

2.14 Beam Deﬂection Analysis

Top beam

W Bottom beam

B

A

D

L 3

L 3

L 3

L

Figure 2-84 Two cantilever beams, loaded and connected at point “B”

The beam structure must comply with the superposition process requirements. The bending moment at any cross section will be equal to the sum of the bending moments produced by the load separately on both cantilever beams. : Calculate the deﬂection of the top beam at pin contact point “B” by the sum of effects caused by the concentrated load “W”. L 3

W D

B L 3

L 3

W

D

L 3

3

B D

D

R

R

2

1

2L 3

Figure 2-85 Top beam superposition deﬂection diagrams

The top beam deﬂection at the pin contact point “B” is calculated by the sum of effects by the concentrated load “W”. δBTop = δ1 + δ2 + δ3 =

W (L /3)3  W (L /3)2   L  R (2 L /3)3 +   − 3E I 3E I  2E I 3

: The deﬂections of the top and bottom beams at the pin contact point “B” are the same, because the beams at this point are moving together. By balancing the deﬂection in both beams, the reaction force “R” is derived with the concentrated load “W” as shown in Figure 2-86. δBottom = B

δBTop = δBottom B

2L 3 D

R Bottom B

B

Figure 2-86 Top beam reaction force at pin contact point “B”

R (2 L /3)3 3E I

5W W (L /3)3  W (L /3)2   L  R (2 L /3)3 R (2 L /3)3 + = →R=    − 3E I 2 3 3 3 32 E I E I E I   : Calculate the deﬂection of the bottom beam at the pin contact point “B” by the reaction load “R” caused by the top beam as shown in Figure 2-87. 3

2

5W  2 L  5W  2 L         W L3 32 3 32   3   L  δA = + =   3E I 2EI 3 370 E I

R =5 W 32

D 2L 3

B L

A L 3

Figure 2-87 Reaction force (B) deﬂects bottom beam free end

186

2 Engineering Product Design Table 2-3 Beam Deﬂection Equations

Beam Loading

Variables

Maximum Deﬂection

δ=

W (2 L3 − 3 a L2 + a3 ) 6×E×I

Max. δA =

MA

W (L − a)2 = 2L

Cantilever Right End, Left End Free, Single Load a W

W × L3 , when a = 0 3×E ×I

L

A

Right End Fix, Left End Guided, Single Load a W

A

δ=

W (L − a)2 (L + 2 a) 12 × E × I

Max. δA =

L

MA

Right End Fix, Left End Support, Single Load

δ= W (L − a)2 (2 L + a) 2 L3

RA =

a W

L

RA

Cantilever Clamp Both Ends, Single Load W

RB = L B

A

a W L

RA

W × a2 (3 L − 2 a) L3

MB =

Simple Support Both Ends, Single Load

W × a (L2 − a2 )3 3 × E × I (3 L2 − a2 )2

W × L3 E×I when a = L (0.414) Max. δ = 0.0098

RA = MA

a

W (L − a)2 (L + 2 a) L3 W ×a = (L − a)2 L2

W × L3 , when a = 0 12 × E × I

δ=

2 × W (L − a)2 a3 3 × E × I (L − 2 a)2

Max. δ =

W × L3 L , when a = 192 × E × I 2

W × a2 (L − a) L2

W (L − a) L W ×a RB = L RA =

RB

δ=

 L2 − a2  W ×a 3 × L × E × I  3 

Max. δ =

3/2

W × L3 L , at x = a = 48 × E × I 2

Left End Guide Moment, Right End Simple Support RB = W a

A

W

M A = W (L − a)

L

MA

RA = W

RA

W

W

L

W × L3 , when a = 0 3×E ×I

RB

Simple Support both Ends with 2 Symmetrical Loads a

Max. δA =

RB = W

a RB

W × a2 (3 L − 4 a) 6×E×I W ×a Max. δ = (3 L2 − 4 a2 ) 24 × E × I Load δw =

187

2.14 Beam Deﬂection Analysis

Beam Loading Uniform Load, Fixed One End

Maximum Deﬂection

W x2 2×L W ×L Max. M = , when x = L 2

δ=

RA = W , M =

W=wxL x A

Variables

W × x2 [2 × L2 + (2 × L − x)2 ] 24 × L × E × I

Max. δA =

W × L3 , when x = L 8×E×I

L

Uniform Load, Two Simple End Supports W=wxL

L RB

Two Simple End Supports, Uniform Load, 1 Couple Guide W=wxL x MA

L RB

RA

Two Simple End Supports, Uniform Load, 2 Couple Guides W=wxL

MA

MB

x L

RA

3×W 8 5×W RB = 8 3 × x x  M =W  − 8 3 × L   W 2 W RB = 2

RA =

Cantilever Right, Free Left End, Vertical & Transversal Loads J =

W

W 2

 L x  x − 6 − L   

E×I L , G= P J

W (3 L × x 3 + 2 x 4 − L3 x) 48 × L × E × I

Max. δ = 0.0054

δ= 2

L 5 × W × L3 , when x = 384 × E × I 2

W × L3 , when x = 0.4215 L E×I

W × x2 (2 × L × x − L2 − x 2 ) 24 × L × E × I

Max. δ =

W × L3 L , when x = 384 × E × I 2

Max. δA =

W ( J × tan G − L), when x = L P

Max. δA =

  W ×J   G2 − sec G  + L (tan G − G) ,  J 1 + 2 P    

M = W × J × tan G

x

A

δ= 2

M =

W ×x (L3 + 2 × L × x 2 − x 3 ) 24 × L × E × I

Max. δ =

RA =

RB

P

δ=

W  x2  M = x−  2  L 

x RA

W 2 W RB = 2

RA =

L

Cantilever Right, Free Left End, Uniform & Transversal Loads J = W=wxL P A

E×I L , G= P J

M = W × J [ J (1 − sec G) + L × tan G]

x

when x = L

L

Two Simple End Supports, Uniform/2 Transversal Loads

J =

E×I L , G= P J

Max. δA =

W=wxL P

P x RA

L

RB

 G  M = W × J 2  sec − 1   2

when x =

W × J2  G G2  sec − 1 − ,  P 2 8   L 2

188

2 Engineering Product Design

2.15

Column Structural Analysis

A long slender vertical bar subject to axial compression loads is known as a column. Failure of a column occurs by buckling or by lateral deﬂection, when the load has exceeded the maximum yield strength of the material. Buckling or failure of a column may occur even at stresses below the yield point of the material.

2.15.1 P

Long Slender Column Critical Load (PCr)

The critical load of a long slender column subjected to axial compression is the value of the axial force that is just sufﬁcient to keep the column in a slightly deﬂected conﬁguration. Figure 2-88 shows a column pinned at both ends in a buckled conﬁguration caused by the critical load (PCr). If a long slender column of constant cross section is pinned at each end and subjected to axial compression, the critical load (PCr) that will cause buckling is given by the following equation: PCr =

P

Figure 2-88 Column, pinned ends axial loaded

π2 E I L2

Where E denotes the modulus of elasticity, I the minimum moment of inertia of the cross sectional area about an axis through the centroid, and L the length of the column. The derivation of this equation is presented in Example 2-37. This critical load (PCr) equation was ﬁrst derived by a Swiss mathematician, Leonhard Euler (1707 – 1783) and the critical load (PCr) is known as Euler’s buckling load equation. This expression is not valid if the corresponding axial stress (found from the expression σCr = PCr / A, where A represents the cross sectional area of the column) exceeds the proportional limit of the material. The value of PCr represented by this equation is the failure load; consequently, a safety factor must be included to obtain a design load for the column.

2.15.2

Column Slenderness Ratio (L / r)

The slenderness ratio of a column (L / r) is determined by the length L of the column the radius r of gyration of the cross sectional area. It is recommended for use only in the range 30 < (L / r) < 120 for main members and a slenderness ratio (L / r) as high as 150 for secondary members. For the design of compression members having a high slenderness ratio, proceed according to Euler’s equation together with an appropriate safety factor. For the design of compression members having lower values of slenderness ratio, it is customary to employ any one of the many empirical formulas giving a relationship between the critical stress and the slenderness ratio of the bar. In fact, the equations usually present an expression for the working stress including a safety factor as a function of the slenderness ratio.

2.15.3

Eccentrically Loaded Columns

Although there are several methods for the rational analysis and design of an eccentrically loaded column, only one of these will be presented here. For a

189

2.15 Column Structural Analysis column subjected to a compressive force P0 acting through the centroid of the cross section together with an additional pressure P applied with an eccentricity e (measured from the centroid), the maximum stress is given by the following equation: σ =

P + Po P e c + A I

Where A represents the cross sectional area of the column and I denotes the moment of inertia of the cross sectional area about the bending axis. The distance from the neutral axis to the extreme ﬁbers of the bar is represented by c. It is necessary to use either the A.I.S.C. speciﬁcations [σW = 17,000 – 0.485 (L / r)2] or the Chicago Building Code equation [σW = 16,000 – 70.00 (L / r)] to obtain a safe value of the allowable compressive stress for use in conjunction with this equation. Y

Example 2-36

L

Calculate the critical load for a long slender column with pinned ends loaded by an axial compressive force at each end. The line of action of the forces passes through the centroid of the cross section of the column. The critical load is deﬁned to be the axial force that is just sufﬁcient to hold the column in a slightly deformed conﬁguration. Under the action of the load pressure P, the column has the deﬂected shape shown in Figure 2-89. It is necessary that one end of the column be able to move axially about the other end, so that lateral deﬂection may take place. The differential equation for the bending moment is: M =EI

dy 2 dx 2

(2-26)

The bending moment at point B having coordinates (X, Y) is the moment of the pressure P applied at the left end of the column around an axis through the point B and perpendicular to the plane of the page. This pressure produces a curvature of the column that is concave downward (negative bending), because the bending moment is deﬁned as M = –P δ. Therefore, we have the following equation: −P δ = E I

dy 2 dx 2

(2-27)

If we deﬁne a constant in the equation: k2 =

P EI

(2-28)

we obtain: dy 2 + k 2 (δ) = 0 dx 2

(2-29)

P

P

B

X x

Figure 2-89 Column, pinned ends, axial loaded

190

2 Engineering Product Design

This equation is solved by differentiating twice, multiplying by a constant, either sin k (x) or cos k (x), and setting the equation equal to zero. Using the combination of these terms, we obtain the following equation: δ = C sin k (x) + D cos k (x)

(2-30)

By the substitution of δ, as given by Equation (2-30), into Equation (2-29), it is next necessary to determine C and D. At the left end of the column, δ = 0 when x = 0. Substituting these values in Equation (2-30) we obtain the following expression: 0 = 0 + D or D = 0 At the right end of the column, δ = 0 when x = L. Substituting these values in Equation (2-30) with D = 0, we obtain: 0 = C sin k (L) Either C = 0 or (L) sin k = 0. But if C = 0, then δ is always zero and we have the conﬁguration before the occurrence of buckling. Because we are not interested in this solution, we must take (L) sin k = 0

(2-31)

For this to be true, we must have (L) k = n radians (n = 1, 2, 3, …) Substituting k 2 =

(L)

(2-32)

P in Equation (2-32) we ﬁnd: EI

2 2 P = n π or P = n π E I EI L2

(2-33)

The smallest load pressure P occurs when n = 1. The ﬁrst mode of buckling where the critical pressure is (PCr), given by: PCr =

π2 E I L2

(2-34)

Example 2-37 Calculate the critical load for a column clamped at each end and axially loaded at each end and with three central support guides as shown in Figure 2-90.

Y L L 4 MO P

L 4 D

L 4

L 4 F

MO P

Figure 2-90 Column, central-supported, axial loaded, both ends ﬁxed

X

The critical pressure load (PCr) is the axial compressive pressure (P) that is just sufﬁcient to keep the column in a slightly deformed conﬁguration. The moments (M0) at each end of the column cause a coupling action of the supports on the column; these moments prevent any angular rotation of the column at either end.

191

2.15 Column Structural Analysis

The deﬂection curve for the buckled column indicates that the central portion of the bar between points D and F correspond to the deﬂection of the pinnedends column discussed in Example 2-37. For the ﬁxed end column, the length (L / 2) corresponds to the entire length (L) for the pinned-ends column. Euler’s buckling PCr for a column with pinned ends and axially loaded, Equation (2-34) in Example 2-37, is used and the length (L) is replaced by (L / 2). Assuming that the maximum stress in the column does not exceed the proportional limit of the material, PCr =

π2 E I (L /2)2

=

4 π2 E I L2

MO

Example 2-38

P

P

Calculate the critical load (PCr) for a long slender column with its base ﬁx, free at the other end, and loaded by an axial compressive force applied at the free end, as shown in Figure 2-91. PCr is that axial compressive pressure P that is just sufﬁcient to keep the column in a slightly deformed conﬁguration. The moment (M0) represents the effect of the support in preventing any angular rotation of the column’s left end.

L

Figure 2-91 Column base ﬁxed and free end axially loaded

Inspection of the deﬂection curve for the buckled column indicates that the entire column corresponds to one half of the deﬂected pinned end column, as discussed in Example 2-37 and using Equation (2-34). The critical pressure (PCr) is developed by replacing the length (L) with (2 L), we obtain the following equation: PCr =

π2 E I (2 L)2

=

π2 E I 4 L2

Example 2-39 A column made of acetal homopolymer has a cross section and dimensions as shown in Figure 2-92. The column is pinned at each end and subject to axial compression. The column is 10.00 in long. The tensile stress of acetal homopolymer is 10,000 psi and the modulus of elasticity is 410,000 psi. Determine the buckling load (PCr) using Euler’s equation.

d s y b

The minimum moment of inertia (I) of this cross section is calculated by using the following equation: I =

2 s b3 + h s 3 2 × 0.125 × 0.753 + 0.75 × 0.1253 = = 0.009 in 4 12 12

Applying the equation for buckling, the critical pressure (PCr), as given in Equation (2-34), Example 2-37, we obtain the following equation: PCr =

π2 E I 9.87 × 410,000 × 0.009 = = 360.15 lb. 100 L2

h

P Cr

s

P Cr 10.00 in

s = 0.125 in b = 0.75 in h = 0.75 in d = 1.00 in

Figure 2-92 “H” cross section, column axially loaded in pinned ends

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2 Engineering Product Design

The area (A) of the column cross section is: A = b d − h (b − s) = 0.75 × 1.00 − 0.75 × (0.75 − 0.125) = 0.282 in 2 The critical stress (σCr) corresponding to the critical pressure (PCr) is: σ Cr =

PCr 360.15 = = 1,277.12 psi A 0.282

Therefore, the critical stress in this example is lower than the tensile stress of acetal homopolymer (10,000 psi).

Example 2-40 A long slender acetal homopolymer column with a circular cross section of 0.50 in in diameter and clamped at each end, is shown in Figure 2-90. The tensile stress of acetal homopolymer is 10,000 psi and the modulus of elasticity is 410,000 psi. Determine the minimum length for which Euler’s equation may be used to determine the buckling pressure (PCr) using a safety factor of 2.0. PCr =

4 π2 E I L2

The buckling critical pressure is: σ Cr =

PCr 4 π2 E I = A A L2

The axial stress prior to buckling is given by: σ Cr =

4 π2 E ( A r 2 ) A L2

=

4 π2 E (L / r )2

But I = A r2 where r denotes the minimum radius of gyration of the cross section. Substituting the moment of inertia values in function of the radius of gyration in the previous equation obtains: I =

π D4 3.1416 × 0.504 = = 0.003 in 4 64 64

The area is: A=

π D 2 3.1416 × 0.504 = = 0.196 in 2 4 4

The radius of gyration is: r =

I = A

0.003 = 0.124 in 0.196

193

2.15 Column Structural Analysis

Or the radius of gyration for a circular cross section is: I D 0.50 = = = 0.125 in A 4 4

r =

The minimum length for which Euler’s equation is found by setting the working stress (σW) equal to the critical stress (σCr) divided by the safety factor. σ Work =

σ Cr 10,000 4 π2 410,000 or L = 7.11 in = = Safety Factor 2.0 (L /0.125)2

The speciﬁc equations required to calculate the critical stress and critical load of the columns are based on the support and type of loads acting on the column structure. These equations are used in the strength of materials design analysis of the column. These column equations are found based on the column support ﬁxtures, types of loading in several engineering strength of materials publications, and plastic suppliers design manuals. The Column Loading Cases and the equations in Table 2.4 were prepared using typical mounting supports, types of loading, and applications.

Table 2-4 Column Loading Cases and Equations

Column Axial Loaded at Both Pinned Ends

Column Base Fixed and Axial Loaded at Top Free End

P Cr

Column Base Fixed and Axial Loaded at Top Pinned End

P Cr

π2 × E × I L2 P = Cr A(Area)

PCr = L

σ cr

P Cr

π2 × E × I 4 × L2 P = Cr a(Area)

PCr = σ Cr

L

1.22 × π2 × E × I L2 P = Cr a(Area)

PCr = σ Cr

L

P Cr

Column base Fixed, Center Support Guides, Axial Loaded at Top Pinned End

Column Axial Loaded at Both Fixed Ends and Three Central Support Guides P Cr

P Cr

0.5 L

0.5 L

4 × π2 × E × I L2 P = Cr a(Area)

L 4 L 4 L 4 L 4

PCr = σ Cr

L

P Cr

General expressions PCr = Critical load (lb) σCr = Critical stress (psi) E = Modulus of elasticity (psi) I = Moment of inertia (in4)

A L r δ

= Cross section area (in2) = Length (in) = Radius of gyration (in) = Column deﬂection (in)

4 × π2 × E × I L2 P = Cr a(Area)

PCr = σ Cr

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2 Engineering Product Design

2.16

Flat Circular Plates

Flat circular plates, such as the bottom of vessels, panels, bulkheads, covers, etc., when supported around the edge and subjected to loads normal to the surface, will bend. The type of bending in plates is different from that in a beam. The plate bends in all planes normal to the plate, whereas the beam bends only in one plane. The bending of a ﬂat circular plate in one plane is greatly inﬂuenced by the bending in all the other planes. The ﬂat circular plate stress and deﬂection analysis takes into account the multiple bending; therefore, these calculations are more difﬁcult than those for a beam structure. When the ﬂat circular plate deﬂects, the middle surface (halfway between top and bottom surfaces) remains unstressed, while at other points there are biaxial stresses in the plane of the plate. Straight lines in the plate that were originally vertical remain straight but become inclined. Therefore, the intensity of either principal stress at points on any such line is proportional to the distance from the middle surface and the maximum stresses occur at the surfaces of the plate. The maximum stress analysis found by the theory of ﬂexure of ﬂat circular plates is usually safe. Comparing the results between a mathematical analysis and testing, we found that the ﬂat circular plates will resist the maximum static loads in testing without being structurally damaged. These test loads were considerably higher than the analytical loads. This deviation is caused by the assumption made in the mathematical analysis that the maximum load is reached when the maximum stress at any point of the ﬂat circular plate reaches the yield strength of the material. This fact is explained by observing how the maximum bending stresses occur only at the clamped rim during the test. The ﬂat circular plate yields slightly before the inside surface and a redistribution of stress is produced that gives the ﬂat circular plate an added usable strength. The equations for the ﬂat circular plates are based on the following assumptions: • The circular plate is ﬂat, of uniform wall thickness, and made of homogeneous isotropic material. • The wall thickness is not more than 25% of the transverse dimension and the maximum deﬂection is not more than 50% of the wall thickness. • All forces, loads, and reactions are normal to the plane of the ﬂat circular plate. • The ﬂat circular plate is not stressed beyond the elastic limit. • The plane of the ﬂat circular plate is horizontal. There are three stages in the behavioral changes of thin and medium thickness, ﬂat circular plates when resisting loads, particularly in the case of a plate restrained at its edges that is made of ductile material. • The stage of purely elastic strain, when the deﬂection of the ﬂat circular plate is strictly proportional to the load and the deﬂection, because of the bending action alone. • The stage of breakdown of elastic action, when the yielding at the portions of high stress becomes sufﬁcient to permit a large permanent elongation (cold ﬂow deﬂection) of the plate. During this stage the direct tension has become an appreciable factor in the resistance of the ﬂat circular plate.

2.16 Flat Circular Plates • The stage of direct tension, in which the tension carries the greater part of the load. During this stage the plate gradually takes a “dished” shape. If the ﬂat circular plate is thin, the major portion of the load is carried by the direct tension even within the purely elastic stage.

2.16.1

Classification

Flat circular plates may be divided into four groups: • Thick ﬂat circular plates, in which the shearing stresses become important, as is in the case for short and deep beam structures. • Medium thick ﬂat circular plates, in which the bending stress is the main action, on which the useful resistance of the ﬂat circular plate depends. • Thin ﬂat circular plates, in which the useful strength resistance depends in part on the tension stress and the stretching action of the middle plane. • Flat circular membranes, in which the strength resistance depends exclusively on the stretching action of the middle plane. These ﬂat circular membranes are not considered to be plates; therefore, there is no elastic bending stress resistance present in these structures. The elastic bending stress resistance of a thin ﬂat circular plate is small. The usable strength resistance is greatly increased, because the direct stresses help to resist the load, even when “dishing” occurs. This explains why the thin ﬂat plate structure requires properly spaced bolts around the ﬂat circular plate edges for support and to prevent dishing.

2.16.2

Stress Analysis Methods

It is assumed that the resistance of a ﬂat circular plate to loads is limited by the magnitude of the stresses in a ﬂat circular plate rather than the elastic deﬂection. The strength of the ﬂat circular plate rather than its stiffness is assumed to limit the maximum loads that may be applied to the plate. The main problem encountered in analyzing ﬂat circular plates is ﬁnding a reliable relationship between the loads acting upon the ﬂat circular plate and the signiﬁcant stresses caused by the loads. Three methods have been employed in solving this problem for thin and medium thick ﬂat circular plates: • In the ﬁrst or strip method, the ﬂat circular plate is assumed to be divided into two systems of strips at right angles to each other. Each strip acts as a beam structure. This method is useful in a qualitative analysis of the behavior in a ﬂat circular plate, but it is less reliable for obtaining accurate quantitative results. • The second method is known as the theory of ﬂexure of ﬂat circular plates. This method resembles the theory of ﬂexure for structural beams. It is assumed that the bending stress is the dominant action in the ﬂat circular plate. It requires that the deﬂection of the ﬂat circular plate be relatively small (1/2 or less than the thickness of the plate), and that the bending stresses contribute substantially to the load resistance of the ﬂat circular plate. It is also assumed that the ﬂat circular plate is in equilibrium and is made of an

195

196

2 Engineering Product Design ideal elastic material, and the stress and strain comply with the theory of ﬂexure of structural beams, where the ﬂexure stress and moment are deﬁned by the following equations: Stress (σ) = M c / I and the differential equation is the Moment (M) = E I (dy2 / dx2) of the elastic stress-strain curve. The assumption for a ﬂat circular plate is that every straight line drawn through the ﬂat circular plate is normal to its middle surface before the plate is bent by the loads. Also, the ﬂat circular plate remains straight and normal to the deﬂected middle surface after the plate is loaded. This method is not recommended for simple types of loading and circular shaped plates. The best method for ﬁnding the bending moment for these cases is by solving Lagrange’s equation. This method is somewhat complicated and will not be discussed in this section. When using the theory of ﬂexure of ﬂat circular plates, do not make allowances for the adjustments for local yielding at the clamping edges (high stress) causing a redistribution of stress. This effect increases the usable strength and added resistance to the ﬂat circular plate, but does not take into account the design of plates, particularly ﬂat circular plates made of ductile materials. • The third method shows the total bending moment at the dangerous section from the statics equations alone. The dangerous section is determined by observing the mode of failure from experiments. The factors by which the average stress corresponds to the total moment are multiplied to obtain the maximum stress at the dangerous section. The stress is also determined from experimental results. In this method, no attempt is made to include the effect of direct tensile stresses, which give the plate added strength as the deﬂection increases.

2.16.3

Flat Circular Plate Equations

Unless otherwise indicated, the equations given in Table 2.5, Parts I and II are based on very closely approximated mathematical analysis and may be accepted as sufﬁciently accurate so long as the assumptions stated hold true. Certain additional facts of importance in relation to these equations are as follows. Under Concentrated Load Conditions It will be noted that all the equations for the maximum stress caused by a load applied over a small area give very high values when the radius of the loaded area approaches zero. Analysis obtained by a more precise method shows that the actual maximum stress produced by a load concentrated on a very small area of radius r0 can be found by replacing r0 with its equivalent radius N (see the last case Central Circular Load, Edge Simply Supported in Table 2.5, I), which depends largely upon the wall thickness of the plate (t) and to a lesser degree on its least transverse dimension. Westergaard gives an approximate equation for this equivalent radius (N): N =

1.6 r02 + t 2 − 0.675 t

This equation, which applies to a plate of any form, may be used for all values of r0 less than ﬁfty percent of (t), for larger values the actual r0 may be used. Use of the equivalent radius (N) can calculate the maximum stress produced by a point loading, whereas the ordinary equation would indicate that these stresses were inﬁnite.

197

2.16 Flat Circular Plates Edge Support Conditions The equations given in Table 2.5, I and II are for various combinations of edges, simply supported and ﬁxed. No exact edge condition is likely to be realized in ordinary construction and a condition of true edge ﬁxity is especially difﬁcult to obtain. Even a small horizontal force at the line of contact may appreciably reduce the stress and deﬂection in a simply supported plate. However, a very slight yielding at nominally ﬁxed edges will greatly relieve the stresses there, while increasing the deﬂection and center stresses. Therefore, it is usually advisable to design a ﬁxed edged plate that can carry uniform loads for somewhat higher center stresses than are indicated by theory.

2.16.4

Flat Circular Plate Stresses

Stress in ﬂat circular plates, simply supported on a circular rim around its edge with uniformly distributed load, is deﬁned by the following parameters: A circular ﬂat plate of radius (a), a constant ﬂat plate wall thickness (t), and a uniformly distributed load (w) to which the circular plate is subjected. The bending moment, about any diameter, caused by the forces that lie to one side of the vertical diametral plane, is the moment of the reaction of the supporting rim minus the moment of the downward load. The magnitude of the load on the half plate is W = (1 / 2) π a2 w and its action line passes through the centroid of the semicircular area. The centroid (y) is located at (4 a) / (3 π) from the diametral plane. The resultant force (R1) of the reaction of the supporting rim must be equal in magnitude to the load (W), but its action line passes through the centroid of the semi-circumference, therefore it acts at the distance (2 a) / π from the diametral plane as shown in Figure 2-93. The bending moment (M) about the diameter is deﬁned by the following equations: M = R1

M =

2a 4a w π a2  2 a 4 a  −W = − π 3π 2  π 3 π 

w a3 3

(2-35)

(2-36)

The bending moment may be equated to the resisting moment at the diametral section. Since the resisting moment holds the bending moment in equilibrium, this expression in terms of the stress at any point in the section is unknown. If the expression for the resisting moment (M) is assumed to be the same (σ I) / c as that in a structural beam that bends in one plane only, the assumption is made that the stress at any point along a diameter is independent of the distance of the point from the center of the ﬂat circular plate, giving us the following equations: wa σI = 3 c

W =

2a dia. 4a 3π 2a

(2-37)

π

3

a σ = w  t 

and c = t / 2, therefore, this equation becomes:

π a2 2

3

2at But I = 12

w

t

2

(2-38)

The value of stress (σ) in Equation (2-38) is the average bending stress of the ﬂat circular plate surface at the diametral section.

R1 =

w

π a2 2

Figure 2-93 Flat circular plate, simply supported, and uniformly distributed load

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2 Engineering Product Design

2.16.5

Theory of Flexure Comparison

Figure 2-94, the curve C-D-E shows the stress distribution on the upper surface of a diametral section of a ﬂat circular plate as obtained from the theory of ﬂexure. The average stress (σ) as given by Equation (2-38) is shown by the straight line A-B. The maximum bending stress occurs at the center of the plate and can be expressed by the following equation, where υ is Poisson’s Ratio. σ Max. =

 a2  3 (3 + υ) w  2  8 t 

W= Approximate stress distribution

A

w π a2 2

C

σ=

3 (3 +

µ)

8

w

a2 t2 D

4 a 3π

Theoretical stress distribution

σ=

w

a2

2a

t2

π B

σ=

3 (1 4

µ)

E

w

a2 t2

R1 =

w π a2 2

Figure 2-94 Flat circular plate, ﬂexure stress distribution under uniformly distributed load

Analyzing the average stress of a ﬂat circular plate made of an acetal homopolymer material having a Poisson’s ratio of υ = 0.35, the maximum bending stress is greater than the average stress indicated by using Equation (2-38). However, experiments indicate that the readjustment of stress accompanying local yielding is lower than stress values given by Equation (2-38). These higher stress values become signiﬁcant, particularly if the ﬂat circular plate is made of a ductile thermoplastic material.

2.16.6

Circular Plates Simply Supported, Concentrated Center Load

Let it be assumed ﬁrst that the load (W) at the center is distributed over a small concentric circular area of radius (r0), as shown in Figure 2-95. The bending moment about a diametral plane A-B (Figure 2-95) is found from the statics equations; the total rim reaction force (R1) on one half of the plate is (W / 2) and its moment arm is (2 a) / π, the total load on the semicircular area of radius (r0) is (W / 2), and its moment arm is (4 r0) / (3 π). Therefore, the bending moment at the diametral section A-B is given by the following equation:

199

2.16 Flat Circular Plates

M =

W a 2W a W a − = π π 3π

 2 rO  1 − 3 a 

W 2

(2-39)

The average ﬂexural stress (σ) at the top and bottom surfaces of the ﬂat circular plate at the diametral plane is found from the ﬂexure equation, where (t) is the wall thickness of the plate and (a) is the radius of the plate.

2 rO

σ =

Mc = I

Wa π

 2 rO  t 1 − 3 a  2 2at 12

3

2

πa

4 rO 3π A

B

a

t

=

2r  3W  1− O 2  3a  πt 

(2-40) R1 =

W 2

Figure 2-95 Flat circular plate, simply edge supported, with concentrated center load

2.16.7

Flat Circular Plate under Concentrated Center Load

If r0 = 0 and the load is concentrated on a very small area at the center of the plate in Equation (2-40), the stress equation becomes the following: σ =

3W

π t2

(2-41)

The maximum bending stress in the ﬂat circular plate is greater than the stress as given by Equations (2-40 and 2-41), particularly for small values of r0. However, for ﬂat circular plates made of a ductile material and subjected to static loads, the value of stress (σ) in these equations may often be regarded as the signiﬁcant stress. If the ﬂat circular plate is made of a brittle material, or if the ﬂat circular plate is subjected to repeated stress, the value of the signiﬁcant stress is deﬁned by the following Equation (2-42), where (K) may be assumed to be 1.50 for small values of r0 and a lesser value for larger values of r0.  3W   2r  σ = K  2  1 − O  3a   π t  

(2-42)

The maximum theoretical values of stresses in ﬂat circular plates as obtained by the theory of ﬂexure of ﬂat circular plates are shown in the equations presented in Table 2.5, Part I.

2.16.8

Flat Circular Plate with Fixed Edge

If a flat circular plate is rigidly held (fixed) so that no rotation or radial displacement occurs at the edges, the average bending moment and average bending stress at any diametral section are less than the values given by the previous equations. This is because the negative bending moment at the edge decreases the positive moment within the central portion of the ﬂat circular plate in much the same way as in a structural beam that is ﬁxed at its ends. The negative moment at the edge of the ﬂat circular plate is usually greater than the moment at the center.

200

2 Engineering Product Design Under service conditions, however, the edges of ﬂat circular plates are rarely completely “ﬁxed”, although they usually are subjected to some restraint. A slight amount of local yielding at the ﬁxed edge may destroy much of the effect of the restraint and thereby transfer the moment to the central part of the ﬂat circular plate. For these reasons, the restraint at the edges of a ﬂat circular plate is considered of less importance, particularly if the ﬂat circular plate is made of a ductile material. A medium thickness ﬂat circular plate with a restrained (ﬁxed) edge will be intermediate in strength between the ﬂat circular plate with a simply supported edge and the ﬂat circular plate with an ideally ﬁxed edge.

2.16.9

Flat Circular Plate Compensation Factor for Deflection

In the equations given in Table 2.5, Part I, the maximum deﬂection of ﬂat circular plates made of an ideal elastic material were obtained by the use of the theory of ﬂexure of ﬂat circular plates. Experiments have veriﬁed these equations for uniformly distributed loads and a simply supported edge. The experiments with ﬁxed edged under uniformly distributed loads showed that the equation for the deﬂection was correct for thin and medium thickness (having a thickness / radius ratio of t / a < 0.1), for deﬂections not larger than approx. 1/2 the ﬂat circular plate thickness. However, for thicker ﬂat circular plates, the measured values of deﬂection were much larger than those computed by the equations. Two reasons are given for this discrepancy. • First, the lack of an ideal ﬁxed edge. • Second, the additional deﬂection in the thicker ﬂat circular plates caused by the shearing stresses. Modiﬁcations are suggested for thicker ﬂat circular plates that have a wall thickness/radius ratio of t / a > 0.1, with ﬁxed edges subjected to uniformly distributed loads. The deﬂection equation marked (*) in Table 2.5, Part I should be multiplied by a compensation factor that depends on the ratio of the wall thickness (t) to the radius (a): (*) H = 1 + 5.72 (t / a)2.

2.16.10 Flat Circular Plate Bending under Edge Boundaries Table 2.5, Part II considers several applications of ﬂat circular plates of constant wall thickness and under axisymmetric loading for several combinations of boundary conditions. Besides the equations, tabulated values of deformation and moment coefﬁcients are given for several common loading cases, including concentrated loading and ﬂat circular plates with circular boundaries. The cases presented in Table 2.5, Part II are expressions given for deformations and reactions at the edges of the ﬂat circular plates, as well as general equations that allow the evaluation of deﬂections, slopes, moments, and shear forces at any point in the ﬂat circular plates. Table 2.5, Part II includes several axis-symmetric loadings, such as uniformly distributed load, central circular load, load at center, and parabolically increasing normal pressure over a portion of the ﬂat circular plates. This method of presentation permits the approximation of any reasonable axis-symmetric distributed loading by ﬁtting an approximate second order curve to the variation in loading and solving the problem by superposition.

201

2.16 Flat Circular Plates Table 2-5 Flat Circular Plate Equations, Part I

W = Concentrated load (lb); w = Unit load (psi); M = Moment (in-lb/in); δ = Deﬂection (in); θ = Change in slope (radians); E = Modulus of elasticity (psi); H = Deﬂection factor (in.); υ = Poisson’s ratio; σ = Stress (psi); t = Wall thickness (in); a = Outer radius (in); b = Inner radius (in); d = Shaft radius; r0 = Radius of load (in); K = Plate constant Case Type

Stress and Deﬂection Equations (Constant Thickness)

Concentrate Center Load Edge Simply Supported

Uniform Distribute Load Edge Simply Supported

σ=

y

W 2a

edge simply supported y w

2a

3 W (1 + υ)  1 a 1 − υ r02   υ + 1 + log r −  2 υ 4 a2  2 πt  0

δMax. =

3 W a2 (1 − υ) (3 + υ) 4 π E t3

σ Max. =

3 w (3 + υ)  a2   2 8 t 

δMax. =

3 w a 4 (1 − υ) (5 + υ) 16 E t 3

For a > r0 y

W

Concentrated Center Load Outer Edge Fixed

2a

r02  3 W (1 + υ)  a log +  r0 4 a2  2 π t2 

δMax. =

3 W a2 (1 − υ2 ) 4 π E t3

σ Max. =

3 w a2 4 t2

δMax. =

3 w a 4 (1 − υ2 ) 16 E t 3

y

w

Uniformly Distributed Load Outer Edge Fixed

σ Max. =

2a

δMax. = H

3 w a 4 (1 − υ2 ) 16 E t 3

For thicker ﬂat circular plates having (t / a = 0.1), multiply the deﬂection equation by the constant (H), where *H = 1 + 5.72 (t / a)2. Central Couple Outer Edge Simply Supported

K =

2a

σ Max. =

2d

M

K =

2a

Central Couple Outer Edge Fixed

M

2b

 2 (a − d)  1 + (υ + 1) log  Ka  

3M 4 π d t2

 2 (0.45 a − d)  1 + (υ + 1) log  0.45 K a  

σ Max.

 2 1   2 a  + 1 υ   3W υ  a 1  log +  − 1  = b μ  2 π t 2  (a2 + b2 )  

δMax.

 1  3 W υ2  2 − 1  (a2 υ   = 4 π E t3  

y

W

3M 4 π d t2

0.10 a2 (d + 0.28 a)2

σ Max. =

2d

2a

Radial Center Load Edge Simply Supported

0.49 a2 (d + 0.7 a)2

 3  1  4 a 2 b 2  + 1 − b 2 )  + 1 2 υ  υ   a  +  log   b  1   1  2 2  1 ( ) + − a b      υ υ − 1

202

2 Engineering Product Design Table 2-5 Flat Circular Plate Equations, Part II

W = Concentrated load (lb); w = Unit load (psi); M = Moment (in-lb); δ = Deﬂection (in); θ = Angular change (rad.); Q = shear (lb/in); E = Modulus (psi); υ = Poisson’s ratio; σ = 6M/t2 (psi); t = Wall thickness (in); a = Outer radius (in); r0 = Radius of load (in); D = E t 3 / 12 (1 – ν2); N = Equivalent radius (in); K, C, L, G = Constants (ratio-dependent) Case Type Outer and Inner Edge Simply Supported; Central Radial Load rO

W

Boundary Values δb = 0, M rb = 0, δa = 0, M ra = 0 θb =

− K θb W a

2

D Qb = K Qb W θ a = θ b C 4 + Qb

b W ro − a a δb = 0, θ b = 0, δa = 0, θa = 0

Qa = Qb

2a

Outer & Inner Edges Fixed; Change in Slope rO

O

W a2 a2 C6 − L6 D D

2a

Special Cases b/a ro / a K δMax. K θa K θb KMrb KMro

K θ D b M rb = Mrb o ; Qa = Qb a a K Qb θ o D Qb = ; δMax. = K δo θ a a2 D M ra = M rb C8 + Qb a C9 + θ o L7 a δa = 0, M ra = 0; M C = w a G17 2

Uniform Distributed Load; Edge Simply Supported g

y rO w

w (a2 − ro2 )2 8 D a (1 + υ)

2.4779

0.8114

0.70

0.70

0.10 0.50

0.3376

0.4145

0.90

0.90

0.50

0.70

K δo

–0.1071 –0.0795 –0.0586 –0.0240 –0.0290

KMrb

–2.0540

KMra

–0.6751 –1.7429 –0.8988 –5.0320 –6.3013

KQb

–0.0915 –17.0670

ro / a

0.00

1.1868 –3.5685

0.20

2.4702

0.3122

4.8176 –23.8910 –29.6041 0.40

0.60

0.80

–0.0637 –0.0576 –0.0422 –0.0230 –0.0067

K θa

0.0961

0.0886

0.0678

0.0393

0.0124

KMC

0.2062

0.1754

0.1197

0.0621

0.0177

0.40

0.60

0.80

−w 2 (3 + υ) (a − ro2 ); If ro = 0, G11 = 0.015, G14 = 0.062, G17 = 2a 16

−w a4 −w a3 − w a 4 (5 + υ) w a2 (3 + υ) w a3 G11 ; L Tθ = G14 ; δC = ; MC = ; θa = D D 64 D (1 + υ) 16 8 D (1 + υ)

Linear Increase Load; Edge Simply Supported rO

w

2a

L Tδ =

θa = Qa =

2a

L Tδ =

−W a  G17  − 2 G11    2 D 1 + υ

2.9405

b/a ro / a

K δC

4

δC =

KQb

0.10 0.50 0.70 0.50 0.70 0.70 0.90 0.90 –0.0102 –0.0113 –0.0023 –0.0017 –0.0005 0.0278 0.0388 0.0120 0.0122 0.0055 –0.0444 –0.0420 –0.0165 –0.0098 –0.0048 –0.4043 –0.3819 –0.0301 –0.0178 –0.0063 0.1629 0.1689 0.1161 0.0788 0.0662

δa = 0, M ra = 0; M C = w a2 G18

ro / a K δC

4

0.00

0.20

δC =

− w a  G18  − 2 G12    2 D 1 + υ

Qa =

−w (2 a2 − ro a − ro2 ) 6a

θa =

w a3  G18 (4 + υ)  − 2 G15  ; If ro = 0, G12 = 0.004, G15 = 0.022, G18 =   D 1 + υ 45

–0.0323 –0.0249 –0.0164 –0.0083 –0.0023

K θa

0.0512

0.0407

0.0278

0.0148

0.0043

KMC

0.0955

0.0708

0.0449

0.0222

0.0061

− w a5 − ro − w a 4 (6 + υ) w a2 (4 + υ) w a3 G12 ; δC = ; MC = ; θa = D a − ro 15 D (1 + υ) 45 15 D (1 + υ)

Central Circular Load; Edge Simply Supported 2a

For r > ro ; δ = Mr =

r

W rO

at r = a; δMax. =

Wr  1 W  (3 + υ) 2 a a + ln  (a − r 2 ) − 2 r 2 ln  ; θ = 16 π D  (1 + υ) r 4 π D  (1 + υ) r

W  a (a2 − r 2 ) N 2  4 (1 + υ) ln + (1 − υ) ; N = 16 π  r a2 r2 

or N = ro , If ro > 0.5 t ; M t =

1.6 ro2 + t 2 − 0.67 t ; If ro < 0.5 t

 W  a N 2  4 (1 + υ) ln + (1 − υ)  4 − 2   16 π  r r   

−W a2 (3 + υ) Wa W ; θ Max. = ; M Max. = 16 π D (1 + υ) 4 π D (1 + υ) 4π

a   (1 + υ) ln N + 1

203

2.16 Flat Circular Plates Besides the usual loadings, Table 2.5 Part II also includes several loading cases that may be described best as externally applied conditions that force a lack of ﬂatness into the ﬂat circular plates. The ﬁrst time we look at Table 2.5, Part II it appears to be a formidable task to calculate the strength of these structures. However, when we consider the number of cases it is possible to present in a limited space, the reason for this method of presentation becomes clear. With careful inspection, we ﬁnd that the constants and functions with subscripts are the same except for the change in variables. In Table 2.5, Part II, the tabulated values in the Special Cases are listed for the preceding functions for the most frequently used denominator values of the variable ratios, such as b / a and r0 / a. Example 2-41 A ﬂat circular plate is made of nylon 6/6 with 33% ﬁber glass reinforcement at 73 °F and 50% relative humidity. The radius is 3.00 in with a wall thickness of 0.25 in. The plate is simply supported around its edge and it is loaded with 500.00 lb at the center. The load is distributed through a round area of 0.125 in radius. Determine the maximum bending stress at the surface of the plate and the maximum deﬂection at the center of the plate. Solution

2a

This ﬂat circular plate and loading are covered in Table 2.5, Part I, case load at center with the outer edge simply supported. The diagram and equations in Figure 2-96 were obtained from this table:

W

t = 0.250 in, w = 500 lb, a = 3.00 in, r0 = 0.125 in,

E = 900,00 psi, υ = 0.39, σ = 18,000 psi σ Max. = =

1 − υ r02  3 W (1 + υ)  1 a + − log  r0 1 + υ 4 a2  2 π t2  υ + 1 3 × 500 (1 + 0.39)  1 3.00 1 − 0.39 × 0.1252  + − log 0.125 1 + 0.39 × 4 × 3.02  2 × 3.1416 × 0.252  0.39 + 1

= 10,794 psi

3 W a2 (1 − υ) (3 + υ) 3 × 500 × 3.02 (1 − 0.39) (3 + 0.39) = 4 π E t3 4 × 3.1416 × 900,00 × 0.253 = 0.158 in

δMax. =

Example 2-42 A thick ﬂat circular plate is made of nylon 6/6 with 33% ﬁber glass reinforcement at 73 °F and 50% relative humidity with a radius of 4.00 in and a uniform wall thickness of 0.50 in. The plate’s outer edge is ﬁxed and it is uniformly loaded along the round area of plate with 200.00 lb/in. Determine the maximum bending stress at the surface of the plate and the maximum deﬂection at the center of the plate.

Figure 2-96 Flat circular plate, concentrated center load, and simply supported edge

δ

204

2 Engineering Product Design δ

w

Solution This thick ﬂat circular plate and type of loading is presented in Table 2.5, Part I, case Uniformly Distributed Load with the Outer Edge Fixed. The diagram and equations in Figure 2-97 were obtained from the table.

2a

Figure 2-97 Flat circular plate, uniformly distributed load, and ﬁxed edge

Because this example case deals with a thick plate, we need to investigate if the thickness / radius ratio is greater than 0.1 to modify the maximum deﬂection by multiplying the value by the constant (H). t = 0.500 in, w = 200 psi, a = 4.00 in, E = 900,00 psi, υ = 0.39, σ = 18,000 psi For thicker ﬂat circular plates with a ratio t / a > 0.1, multiply the deﬂection equation by the constant (H), where H = 1 + 5.72 (t / a)2. t 0.50 = = 0.125 > 0.1 a 4.00 H = 1 + 5.72 (t / a)2 = 1 + 5.72 (0.50/4.00)2 = 1.089 σ Max. = δ=

3 w a2 3 × 200 × 4.02 = = 12,223.00 psi π t2 3.1416 × 0.502

3 w a 4 (1 − υ2 ) 3 × 200 × 4.04 (1 − 0.392 ) = = 0.072 in 16 E t 3 16 × 900,00 × 0.503

δ = H x y = 1.089 × 0.072 = 0.079 in

Max.

Example 2-43 A ﬂat circular plate, made of acetal homopolymer, has a wall thickness of 0.187 in and a 5.00 in outside diameter, and is simply supported with a uniformly distributed load of 6.0 psi. Calculate the maximum deﬂection in the center, the maximum stress, and the deﬂection equation for Figure 2-98. This ﬂat circular plate and type of loading is presented in Table 2.5, Part II, case Uniformly Distributed Load Edge Simply Supported. First, we need to determine the maximum moment, the bending stress, the plate constant, and the deﬂection caused by the load. Second, we need to calculate the total deﬂection of the plate caused by the load, the moment, and the loading constant. Finally, we need to check the deﬂection at the outer edge. t = 0.187 in, w = 6.0 psi, a = 2.50 in, r0 = 0,

rO w

E = 410,000 psi, υ = 0.35, σ = 10,000 psi M Max. = M Center =

w a2 (3 + υ) 6.0 × 2.502 (3 + 0.35) = = 7.85 lb-in. 16 16

2a

Figure 2-98 Flat circular plate, uniformly distributed load with simply supported edge

σ Max. =

6 M 6 × 7.85 = = 1,339.97 psi t2 0.1872

205

2.16 Flat Circular Plates

D= δC =

E t3 410,000 × 0.1873 = = 256.66 2 12 (1 − υ ) 12 (1 − 0.352 ) −w a 4 (5 + υ) −6.0 × 2.504 (5 + 0.35) = = −0.0565 in 64 D (1 + υ) 64 × 256.66 (1 + 0.35)

The total deﬂection equation for the ﬂat circular plate is: δa = δC +

MC y2 −w a4 G11 + L Ty , where for this case L Ty = D 2 D (1 + υ)

Where the constant G11 = 0.015, when r0 = 0. 7.85 × a2 6.0 × 0.015 × a 4 − 2 × 256.66 × 1.35 256.66 2 = −0.0565 + 0.01132 × a − 0.000365 × a 4

δa = −0.0565 +

Checking the deﬂection at the outer edge, when a = 2.50 in δa = −0.0565 + 0.01132 ⋅ 2.502 − 0.000365 ⋅ 2.504 = −0.0565 + 0.07075 − 0.01425 = 0.0

Example 2-44 A ﬂat circular plate is made of acetal homopolymer with a wall thickness of 0.125 in and 4.00 in outside diameter. It is mounted in a ﬁxture to produce a sudden change in slope in the radial direction of 0.05 radiant at a radius of 0.75 in. It is then clamped between two ﬂat ﬁxtures as shown in Figure 2-99. Calculate the maximum bending stress. This is an example of forcing a known change in slope into a ﬂat circular plate, clamped (ﬁxed) at both inner and outer edges. This ﬂat circular plate and type of loading is presented in Table 2.5, Part II, case Outer and Inner Edge Fixed and Change in Slope, where: θ0 = 0.05, b / a = 0.10, r0 / a = 0.50 and Poisson’s ratio of υ = 0.35. t = 0.125 in, a = 1.50 in, b = 0.15 in, r0 = 0.75 in,

θ0 = 0.05 rad., θb = 0.0 rad., δb = 0.0 in, E = 410,000 psi,

υ = 0.35, σ = 10,000 psi D=

Qb =

3

Et 410,000 × 0.125 = = 76.04 12 (1 − υ2 ) 12 (1 − 0.352 )

K Mrb × θ × D −2.054 × 0.05 × 76.04 = = −5.20 lb-in. 1.50 a

K Qb × θ × D a2

θO 2a 4.00 dia.

3

M rb =

rO

=

−0.0915 × 0.05 × 76.04 = −0.154 lb/in. 1.502

0.125

0.15 r.

0.05 rad.

1.50 r.

0.75 r.

Figure 2-99 Flat circular plate having a change in slope with both outer and inner edges ﬁxed

206

2 Engineering Product Design

M ra = M rb C8 + Qb a C9 + θ0

D L7 a

= −5.20 × C8 + (−0.154) × a × C9 + σ Max. =

0.05 × 76.04 L7 a

6 × M rb 6 × 5.20 = = 1,996.80 psi 2 t 0.1252

δMax. = K y0 θ r0 = −0.1071 × 0.05 × 1.50 = 0.008 in

Example 2-45 A ﬂat circular plate, made of acetal homopolymer, has a wall thickness of 0.250 in and 5.00 in outside diameter, it is simply supported at the outer edge and subjected to two types of loads. One center load provides a uniform pressure over a diameter of 0.0625 in. The other is axis-symmetrically loaded with a distributed load that increases linearly from the center to the outside radius rO = 1.00 in;, this load has a value of 10.00 psi at the outer edge. Calculate the maximum bending stress. This example requires analyzing two different cases and to superposition the results. The ﬁrst case is the linear increase of the distributed load with simply outer edge supported (Figure 2-100), the second case is the central circular uniform load with simply supported outer edge (Figure 2-101). Both cases are presented in Table 2.5, Part II. t = 0.250 in, a = 2.50 in, r01 = 1.00 in, r02 = 0.031 in, E = 410,000 psi, υ = 0.35, σ = 10,000 psi

From the special case data, the following variable ratios are obtained: rO1 w

r01 / a = 1 / 2.5 = 0.40, K yC = –0.0164, K θa = 0.0278, KMC = 0.0449 D=

2a

Figure 2-100 First case: Linear decreasing distributed load and edge simply supported

E t3 410,000 × 0.2503 = = 608.38 12 (1 − υ2 ) 12 (1 − 0.352 )

δ = K yC

w a4 −0.0164 × 10 × 2.54 = = 0.0105 in D 608.38

M = K MC w a2 = 0.0449 × 10 × 2.52 = 2.80 lb-in. 2a r P

δMax. =

− P a2 (3 + υ) − P × 2.502 (3 + 0.35) = = 0.0105 in 16 π D (1 + υ) 16 × π × 608.38 (1 + 0.35)

P = −20.76 lb. rO2

Figure 2-101 Second case: Center uniformly circular load and edge simply supported

The second moment component is calculated by using the equations provided in Table 2.5, Part II, case Central Circular Loading and Simply Outer Edge Supported.

207

2.17 Torsion Structural Analysis

N =

2 1.6 r02 + t 2 − 0.675 t =

1.6 × 0.032 + 0.252 − 0.675 × 0.25

= 0.085 in P  a  −20.76 (1 + υ) ln + 1 =  4π N 4π  = −9.19 lb-in.

M Max. =

σ Max. =

2.17

2.50   (1 + 0.35) ln 0.085 + 1

6 M 6 (−9.19 + 2.80) = = −613.44 psi t2 0.2502

Torsion Structural Analysis

A bar is rigidly clamped at one end and twisted at the other end by a torque T = F × d, applied in a plane perpendicular to the axis. Plane sections remain plane and radii remain straight. There is at any point a shear stress (τ) on the plane of the section; the magnitude of this stress is proportional to the distance from the center of the section and its direction is perpendicular to the radius drawn through the point. The deformation and stresses are shown in Figure 2-102. In addition to these deformations and shear stresses, there are the longitudinal strain and stress. The longitudinal strain is reduced while the stress is in tension on the outside and a balancing compression stress is exerted on the inside. Assumptions The torsion equations are based on the following assumptions: • The bar is straight, of uniform circular cross section (solid or tubing), and of homogeneous isotropic material. • The bar is loaded only by equal and opposite twisting couples, which are applied at its ends in a normal direction to its axis. • The bar is not stressed beyond the elastic limit of the material. Angle of Twist (θ) If a shaft of length (L) is subjected to a constant twisting moment (T) along its length, then θ is the angle through which only one end of the bar will be twisted. Twisting Moment (T) The twisting moment T for any section along the bar is deﬁned to be the algebraic sum of the moments of the applied couples that lie to one side of the section in question. Shearing Strain If a bar is marked on the surface (unloaded), then after the twisting moment (T) has been applied, this line moves as shown in Figure 2-102. The angle (θ) is

d T=Fxd

F

L

Figure 2-102 Deformation and stress under torque

208

2 Engineering Product Design measured in radians; the ﬁnal and original position of the generator is deﬁned as the shearing strain at the surface of the bar. Shearing Stress (τ) For a solid circular cross section bar, let T = Twisting moment; L = Length of the bar; r0 = Radius; J = Polar moment of inertia; τ = Shear stress; θ = Angle of twist (radians); G = Modulus of rigidity. Then: θ = (T L)/(G J ) ; τ Max. = (T r0 )/ J By substituting for J = (π r04 )/2 in the equation above for a solid circular cross section with radius r0, the following equations are obtained: θ = (2 T L)/(π r04 G) ; τ Max. = (2 T )/(π r03 ) For a circular tube cross section with outer radius r0 and inner radius ri: θ = (2 T L)/ π (r04 − ri4 ) G] ; τ Max. = (2 T r0 )/[π (r04 − ri4 )] The torsional stiffness of the bar can be expressed by the general equation: θ = (T L) / (G K), where K is a factor dependent on the bar cross section. For cross section bars, the factor K is equivalent to the polar moment of inertia J. In Table 2-6, the equations for the factor K and for the maximum shear stress (τMax.) for a variety of cross section bars are given. Example 2-46 Compare the strength and stiffness of a circular injection molded tube made of a plastic material, 1.00 in outside diameter and 0.187 in wall thickness, versus an extruded solid circular bar of the same material with the same diameter. The strengths of both cross sections will be compared by using the twisting moments (T) required to produce the same shear stress. The stiffness will be compared by using the values of factor (K) for both cross sections. For the circular tube bar: K = π (r04 − ri4 )/2 = 3.1416 (0.504 − 0.3134 )/2 = 0.083 in 4 T = τ π (r04 − ri4 )/(2 r0 ) = τ × 3.1416 (0.504 − 0.3134 )/(2 × 0.50) = τ × 0.166 lb-in.

For the solid circular bar: K = π r 4 /2 = 3.1416 × 0.504 /2 = 0.098 in 4 T = (τ π r 3 )/2 = τ × 3.1416 × 0.503 /2 = τ × 0.196 lb-in. The solid circular cross section bar is therefore 1.182 times as stiff as the circular tube cross section bar and 1.18 times as strong.

209

2.17 Torsion Structural Analysis Table 2-6 Torsion Equations

Constant K in θ =

Cross section

TL KG

Shear stress max.

Solid circle 2rO

K =

π rO4 2

τ Max. =

K =

π (rO4 − ri4 ) 2

τ Max. =

2T π rO3

Circular tube

ri

rO

2 T rO

π (rO4 − ri4 )

Solid ellipse 2rS

K =

π rL3 rS3 rL2 + rS2

τ Max. =

2T π rL rS2

τ Max. =

T 0.208 a3

2rL

Solid square a

K = 0.1406 a 4 a

Solid rectangle a

K =

b a3  a 5.33 − 3.36 16  b

 T (3 b + 1.8 a) a4   τ Max. = − 1  4  b2 a2 12 b   

b

θ = Angle of twist (radians); T = Twisting moment (lb-in); τ = Shear stress (psi); G = Modulus of rigidity (psi); J = Polar moment of inertia (in4); K = Constant equivalent to J (in4); ro = Outer radius (in); ri = Inner radius (in); rS = Elliptical short radius (in); rL = Elliptical large radius (in); a = Height (in); b = Width (in).

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211

3

Structural Designs for Thermoplastics

3.1

Uniform and Symmetrical Wall Thickness

Poor design

The ultimate design rule for injection molding thermoplastic products is to ensure that the wall thickness is uniform and symmetrical.

Molding problems

Non-uniform and/or heavy wall thicknesses can cause serious warpage and dimensional control problems in the injection molded products. Heavy wall sections cause not only internal shrinkage, voids, and surface sink marks, but also nonuniform shrinkage resulting in poor dimensional control and warpage problems. Figure 3-1 shows a poor cross section design of perpendicular corner walls that causes molding problems, such as differential shrinkage, warpage (concave) of both walls, and internal voids in the corner of the thicker wall. The last two designs are recommended to avoid these molding problems.

Sharp corner

Warpage

Voids

Good design

Good design

Figure 3-2 shows a heavy wall cross section design that could cause molding problems and the recommended design using a thin wall and proportional ribs.

r.

r.

Figure 3-3 shows a nonuniform wall section that should be replaced with a thin uniform wall having the same strength of the original heavy wall section. Figure 3-4 shows another poor and the recommended uniform wall design. Figures 3-5 and 3-6 show cross sections of two nonuniform wall designs and the recommended designs with a uniform wall thickness to avoid warpage, internal voids, long molding cycles, and surface sink marks.

Poor design

Seat

R.

Figure 3-1 Perpendicular walls, end corner designs

Good design

Figure 3-2 Heavy wall vs. thin uniform ribbed wall designs

Poor design

Good design

Figure 3-5 Nonuniform wall vs. thin uniform wall designs

Poor design

Good design

Figure 3-3 Nonuniform wall vs. thin uniform wall designs

Poor design

Good design

Figure 3-4 Nonuniform wall vs. thin uniform wall designs

Poor design

Good design

Figure 3-6 Nonuniform wall vs. thin uniform wall designs

212

3 Structural Designs for Thermoplastics Poor design

3.1.1

Part Geometries Difficult to Mold

The most serious defects caused by part geometry during the injection molding process are warpage, internal voids, surface ﬁnishing, dimensional control, and sink marks. Molding problems

Good design

These problematic part geometries are illustrated in Figure 3-7, which shows a poor design of a uniform wall thickness rectangular tray. The tray top surface corners are warped upward, while the vertical side walls are warped inward. The recommended design calls for a small crown on the top surface and the side walls to compensate for the warpage, with radii in all corners and a uniformly tapered wall starting from the center (thicker wall for gating) until the side wall ends. Figure 3-8 shows a poor design of a square box. The vertical side walls of the molded box are warped inwards. The recommended design calls for a small crown on the side walls to compensate for the warpage and radii in all corners. Figures 3-9, 3-10 and 3-11 show poor designs. The molded vertical walls are warped. The proper designs have tapered walls and radii in all corners.

Figure 3-7 Problematic rectangular tray design

Poor design

Molding problems

Poor design

Good design

Figure 3-9 Problematic electrical bobbin (spool) design

Poor design

Molding problems

Good design

Molding problems

Figure 3-10 Problematic “U” beam design

Good design Poor design

Good design

Figure 3-8 Problematic square box design

Figure 3-11 Problematic structural beam design

Molding problems

Good design

213

3.2 Structural Rib Design

3.1.2

Wall Draft Angle per Side

Draft angles for internal and external walls are essential to the ejection of the molded parts from the mold. External walls require smaller draft angles than the internal walls. Thermoplastic material expands in volume inside the plastifying unit and the solid material is transformed into a ﬂowing melt. Then, the hot melt is injected inside the mold. The cold temperature inside the cavities initiates cooling-off and shrinking the hot melt. The amount of shrinkage of a molded part is a product of the mold shrinkage characteristics of the polymer, the part wall thickness, injection/packing time, mold temperature, and cooling time. During the shrinkage process, the molded part’s external walls shrink away from the cavity’s external walls, while the internal walls shrink in around the core surface or walls. Semi-crystalline thermoplastic materials have higher mold shrinkage characteristics than amorphous materials. Parts made of semi-crystalline materials require higher draft angles for their internal walls, while amorphous materials that have lower mold shrinkage characteristics, require higher draft angles for their external walls, lower mold temperatures, and longer cooling times. When a molded part requires an internal wall with a minimum draft angle, it is recommended to have an efﬁcient mold temperature control, with mold cavities and cores made of hardening tool steel, well polished in the direction of ejection and low coefﬁcient of friction surface coating on the cores. For external walls without texture made of either unreinforced or reinforced resins, a minimum draft angle of 0° 15′ to 0° 30′ is recommended. For internal walls, a minimum draft angle per wall of 0° 30′ to 1.0° is recommended. For parts made of mineral/ﬁber glass reinforced resins having internal walls without texture less than 1.00 in deep, a minimum draft angle per wall of 1.0° to 1.0° 30′ is recommended. For internal walls without texture deeper than 1.00 in, a minimum draft angle per wall of 1.0° 30′ to 3.0° is recommended. Figure 3-12 shows the equation to calculate the mold cavity external wall draft dimension per side, based on the draft angle and the cavity depth.

3.2

Structural Rib Design

When designing injection molding products with thermoplastic materials, it is critical to maintain uniform thin and symmetrical wall thicknesses. Nonuniform wall thickness can cause serious warpage, sink marks, and dimensional control problems. If greater strength or stiffness is required for a product design, it is more economical to use proportional ribs and thin base walls rather than thickwalled sections. For products requiring good surface appearance, the proper selection of thermoplastic materials is very important. For many resins, the use of ribs should be avoided, because they produce sink marks on the external surface and this defect becomes very noticeable on the molded product. If ribbing is necessary in a surface appearance application, the base wall thickness of the rib should be 40 to 50% of the base wall thickness, with a draft angle per wall of 0° 45′. There are several resins on the market that have good surface appearance behind the rib area, for example, PVC, ABS, PC, LCP, PBT, PET, among others. The sink mark problem could be improved by hiding the sink marks behind an external texture, letter, or surface undulation. For applications requiring uniform base

Draft angle ( ˚) Draft = L x tan ( ˚)

L

Figure 3-12 Mold cavity wall draft angle per side

214

3 Structural Designs for Thermoplastics Base wall thickness

R.

Sink mark

Differential thermal warpage

Heavy cross section area

Voids

Molding problems

Rib thickness

Poor design

Rib wall thinner than the base wall Two thinner ribs better than one

Good design

Good design

Neck down at rib intersection

Base wall Rib

Base wall Rib

Neck down at rib intersection

Good design

Base wall

Partially connected rib

Base wall Rib

Rib bridge connector Rib bridge connector

Good design

The upper sides of base wall and the rib are not connected, just a small lower area of the base wall and the rib are attached with a reduce neck down wall section

Figure 3-13 Poor rib design, molding problems, and recommendations

215

3.2 Structural Rib Design walls and ribs, the use of heavy wall sections must be avoided, which can not only cause sink marks, but also internal voids, molded-in stresses, and nonuniform mold shrinkage. Figure 3-13 shows some typical injection molding problems caused by a poor rib design.When the thermoplastic material and the product application allow design changes, the design recommendations will help to overcome this problem. The upper sides of base wall and the rib are not connected, just a small lower area of the base wall and the rib are attached with a reduced neck down wall section. The ribs should be 40% of the base structure wall and a draft angle per wall of 0° 45′ should be used, if a good surface appearance is required for the application. When using reinforced thermoplastic materials and the application does not require a good ﬁnish, the ribs should be between 60 to 100% of the base wall and a draft angle per wall of 1° should be used. For structural foam applications where good surface ﬁnishing is not required,a rib wall equal to the base wall thickness and a draft angle of 1.0°per wall should be, used.

3.2.1

Rib Strength Analysis Method

0˚ 45' T= 0.4 W

Ribs are reinforcements used to improve the strength and rigidity of injection molded thermoplastic products. Ribs have been the key to the successful replacement of metals by engineering thermoplastics. Properly designed and located ribs not only increase the load-carrying ability of thermoplastic structural sections, but also lower manufacturing costs, increase cycle times, eliminate the use of heavy wall thickness that cause sink marks, reduce heat spots in the mold, lower warpage, and improve the dimensional control of the molded products. Rib design can be troublesome, especially when the product designer has to rely on guess work to determine the strength, geometry, size, and spacing of the ribs. Three basic rib designs have been developed for the plastics industry based on the type of thermoplastic resins and product application requirements. The ﬁrst rib design provides a good surface appearance, the second rib design is for structural applications that do not require good ﬁnish and are made of reinforced resins, the third rib design is for structural applications with poor surface ﬁnishing and are made of reinforced or foam resins. Figure 3-14 shows these rib designs. The rib strength analysis method requires the aid of computerized graphs to calculate the stress and the deﬂection of complex structures and predict the behavior of a ribbed thermoplastic structure. Four full size computerized graphs for calculating stress and deﬂection of a symmetrical tapered (1°) rib, where the rib wall thickness is either 60 or 100% of the base structure uniform wall, can be found in Figures 3-15, 3-16, 3-17, and 3-18.

W

Surface appearance rib design 1˚ T= 0.6 W

H

0.03 R.

W

Structural reinforced resins rib design 1˚

Each graph is a composite of several curves and is expressed as a dimensionless ratio to permit calculations in either the Standard International or English units. The graph curves are comparisons between the cross section moment of inertia of a hypothetical ﬂat rectangular plate and the ribbed wall structure.

T= W

0.03 R.

The wall thickness of a rectangular plate would be based on the calculations an engineer might make in substituting thermoplastic for other materials in a structure that must withstand a speciﬁed load. When the width of a rectangular cross section wall is analyzed, it is divided into smaller equal sections and the moment of inertia for a single section is calculated and compared with that of its ribbed equivalent structure. The sum of the small section moments of inertia is equal to that of the original rectangular cross section.

H

0.03 R.

H

W

Structural reinforced or foam resins rib design Figure 3-14 Three rib design geometries

216

3 Structural Designs for Thermoplastics R.

1˚ H

W

WS 0.6W B

B

10

9 0.62 8

1.87 5 2.5 4

5.0 6.25 7.5 10 12.5 15 20 25 37.5 50 75 150

3

2

Computer programmed curves for calculating stress of an injection molded plastic ribbed structure, with a tapered rib, lower wall thickness equal to 60% of the base structure wall thickness

1

BEQ

3.75

W

Wall thickness for stress

1.25

6

W

WS

Base equivalent width

1.0

7

0 0

1

2

3

4

5

6

7

8

9

10

H Height of rib W

Figure 3-15 Rib (60%) stress analysis graph

R.

1˚ H

W

WD 0.6W B

B

10

0.62 9

1.0 1.25

8

5

4

Computer programmed curves for calculating deflection of an injection molded plastic ribbed structure, with a tapered rib, lower wall thickness equal to 60% of the base structure wall thickness Figure 3-16 Rib (60%) deflection analysis graph

W

WD

3

150

2

1

0 0

1

2

3

4

5

6

H Height of rib W

7

8

9

10

Base equivalent width

3.75 5.0 6.25 7.5 10 12.5 15 20 25 37.5 50 75

6

W

2.5

BEQ

Wall thickness for deflection

1.87 7

217

3.2 Structural Rib Design R.

1˚ H

W

WS W B

B

10 1.0

9

2.5

5

3.75 5.0 6.25 7.5 10 12.5 15 20 25 37.5 50 75

3

2

1

W

6

BEQ

1.87

4

W

WS

Wall thickness for stress

7

Base equivalent width

1.25

8

Computer programmed curves for calculating stress of an injection molded plastic ribbed structure, with a tapered rib, lower wall thickness equal to 100% of the base structure wall thickness

150 0 0

1

2

3

4

5

6

7

8

9

10

Figure 3-17 Rib (100%) stress analysis graph

H Height of rib W R.

1˚ H

W

WD W

B

B

10 1.25 9 1.87 2.5

8

10 12.5 15

5

20 25

4

37.5 50 75

W

WD

3

Base equivalent width

6

W

5.0 6.25 7.5

BEQ

Wall thickness for deflection

3.75 7

150 2

1

0 0

1

2

3

4

5

6

H Height of rib W

7

8

9

10

Computer programmed curves for calculating deflection of an injection molded plastic ribbed structure, with a tapered rib, lower wall thickness equal to 100% of the base structure wall thickness Figure 3-18 Rib (100%) deflection analysis graph

218

3 Structural Designs for Thermoplastics The equations for the ribbed structure and the rectangular cross sections are developed to compare the equivalent terms used in the equations. t = T − 2 H × tan(θ°)

Draft angle ( ˚) t

A (Area) = B × W + H

R.

T

H (T + t ) 2

W

 3 B W 2 + 3 H t (H + 2 W ) + H (T − t )(H + 3 W )  y (Centroid) = H + W −   6A  

B

I (Moment of Inertia) =

1 [4 B W 3 + H 3 (3 t + T )] − A (H − y )2 12

WD = Wall thickness for deﬂection, WS = Wall thickness for stress W D or W S

I (Moment of Inertia) =

B

Z (Section Modulus) =

B WD3 (or B WS3 ) 12

B WD2 (or B WS2 ) 6

To deﬁne one of the smaller sections of the whole structure, the term Base Equivalent Width (BEQ) is used and is equivalent to the base width divided by the number of ribs. BEQ =

Total Base Width B = Number of Ribs N

Using these equations, the thickness ratios for the types of ribbed structures used in designs of thermoplastic products are determined. These equations are programmed to develop the family of curves shown in Figures 3-15 to 3-18. The results were plotted for a symmetrical tapered (1°) rib thickness equal to either 60 or 100% of the structure base wall thickness. The curves in the computerized graphs are given in terms of wall thickness for deﬂection (WD / W) or wall thickness for stress (WS / W). The abscissa is expressed in terms of the rib height (H / W). It is important to use the height of the rib rather than the total height of the structure [(H + W) / W]. Examples 3-1, 3-2, and 3-3 illustrate step by step how to use the computerized graphs to simplify the ribbed structure calculations for deﬂection and stress. Example 3-1 An injection molded nylon 6/6 part with 33% ﬁber glass reinforcement is being considered as a replacement for an aluminum plate. The replacement must have a stiffness at least equivalent to the 0.156 in thick aluminum plate. One side can be ribbed, but in order to ﬁt the existing housing, the overall thickness must not exceed 1.00 in. Aluminum has a modulus of elasticity of EA = 10,300,000 psi, while the modulus of elasticity for 33%

219

3.2 Structural Rib Design

ﬁber glass reinforced nylon 6/6 (50% relative humidity at 73 °F) has a value of EN = 900,000 psi. We selected a typical base wall thickness for nylon 6/6; W = 0.10 in. To achieve equal or greater stiffness, the deﬂection under a given load for the reinforced nylon 6/6 part must not exceed the deﬂection of the aluminum part under the same loading conditions. The deﬂection formula is proportional to 1 / (E · I), where E is the modulus of elasticity of the material and I is the cross section moment of inertia. I is proportional to the third power of the rib wall thickness (WD3). The equivalent wall thickness for deﬂection of a plain, ﬂat reinforced nylon 6/6 part can be determined by the following calculation: 1 1 = 3 E A × WA E N × WN3 1/ 3

 E × WA3  WN =  A   EN 

1/ 3

 10,300,000 × 0.1563  =  900,000  

= 0.351 in

The reinforced nylon 6/6 part equivalent wall thickness for deﬂection is WD = 0.351 in, but because this heavy wall thickness is not economical or practical for injection molding, we select a structural base wall thickness of W = 0.10 in with a base ribbed wall thickness of T = 0.06 in (60% W). Because the maximum structure height allowable is 1.00 in, the maximum rib height is H = 1.00 – 0.10 = 0.90 in. In order to use the rib (60% W) deﬂection analysis graph correctly, the height (H), the equivalent base width (BEQ), and the wall thickness for deﬂection (WD) values must be divided by the ribbed base structure wall thickness (W). 8.40 Alumin um pla

WD 0.351 = = 3.51 W 0.100

te

14.00

H 0.90 = = 9.0 W 0.10 Then, from Figure 3-16 we obtain:

0.156

1.00

BEQ / W = 28 or BEQ = 28 · 0.10 = 2.8 8.40

The number of ribs required for the 8.40 in height is 8.40 / 2.8 = 3 ribs. For the 14 in width it is 14 / 2.8 = 5 ribs. Figure 3-19 shows the original aluminum plate and the replacement ribbed structure made of reinforced nylon 6/6. Because the ribs intersect, the reinforced nylon 6/6 structure will be more rigid than the aluminum plate, thus providing an additional design safety factor.

Nylon 6

/6, 33%

reinforc ed

14.00

0.10

Figure 3-19 Aluminum plate and the equivalent nylon 6/6 design

220

3 Structural Designs for Thermoplastics

P (load) = 70.00 psi

0.15 inch 10.0 inches

4.0 inches

Example 3-2 A copper cantilever beam plate, ﬁxed at one end and subjected to a uniformly distributed load of 70 psi, is shown in Figure 3-20. Calculate the equivalent ribbed cross section in acetal homopolymer, its deﬂection, and stress. The modulus of elasticity for copper is: EC = 15,600,000 psi

Figure 3-20 Cantilever copper beam

The modulus of elasticity for acetal homopolymer is: EA = 410,000 psi Designing for the equivalent deﬂection, the equation is reduced to: 1 1 = 3 EC × WC E A × WA3 1/ 3

 E × WC3  WA =  C   EA 

or

1/ 3

 15,600,000 × 0.153  =  410,000  

= 0.50 in

A wall thickness of 0.50 in is not practical for thermoplastic structures, because of processing difﬁculties; therefore, an acetal homopolymer ribbed section is recommended. Modifying the known data in terms applicable for the computerized graphs and assuming a typical base structural wall thickness for injection molding of W = 0.12 in, calculate for a cantilever beam with nine equally spaced ribs. WA 0.50 = = 4.16 W 0.12 BEQ =

4.00 BEQ 0.44 B = = 3.70 = = 0.44 or W 0.12 N 9.00

From Figure 3-16, we obtain: H / W = 5.50, or H = 5.50 · 0.12 = 0.66 in. Applying the H / W = 5.50 and the BEQ / W = 3.70 to Figure 3-15, we obtain: WS / W = 2.75, or WS = 2.75 · 0.12 = 0.33 in. Calculate the moment of inertia and section modulus for the ribbed structure. I =

B × WA3 4.00 × 0.503 = = 0.041 in 4 12 12

Z =

B × WS2 4.00 × 0.332 = = 0.0726 in 3 6 6

221

3.2 Structural Rib Design

Maximum deﬂection of the cantilever beam at the free end: δMax. =

P × L3 70 × 10.03 = = 0.52 in 8 × E × I 8 × 410,000 × 0.041

Maximum stress of the cantilever beam at the ﬁxed end: σ Max. =

P×L 70 × 10 = = 4820.93 psi 2 × Z 2 × 0.0726

Because the acetal homopolymer has a tensile strength value of 10,000 psi, a design safety factor of 2.0 is obtained for this example.

Example 3-3 Calculate the deﬂection and stress for the structure shown in Figure 3-21. The ends are simply supported, with a uniformly distributed load and it is made of 30% ﬁber glass reinforced polyester (PET). Substitute the known data: BEQ =

B 2.50 = = 0.625 N 4.00

H = 0.75 − 0.12 = 0.63

BEQ 0.625 = = 5.20 W 0.12 H 0.63 = = 5.25 W 0.12

From Figures 3-17 and 3-18, the following results are found: WD /W = 4.30 or WD = 4.30 · 0.12 = 0.516 in WS / W = 2.70 or WS = 2.70 · 0.12 = 0.324 in

P (load) = 150.0 psi

Determine the moment of inertia and section modulus: I =

B × WD3 2.50 × 0.5163 = = 0.0286 in 4 12 12

Z =

B × WS2 2.50 × 0.3242 = = 0.0437 in 3 6 6

20.00 inch

2.50

Maximum deﬂection in the middle of the beam: δMax. =

0.12

5 × P × L3 5 × 150 × 20.03 = = 0.42 in 384 × E × I 384 × 1,300,000 × 0.0286

Because 30% ﬁber glass reinforced polyester (PET) has an approximate tensile strength value of 17,000 psi, a design safety factor of 2.0 is obtained in this example.

0.12 0.75

1˚

Figure 3-21 Beam, uniformly distributed load, ends simply supported

222

3 Structural Designs for Thermoplastics

3.3 Internal sharp corner

Poor design Pullout force Pullout force

Failure

Operational problem

r.

R.

Internal Sharp Corners and Notches

Internal sharp corners and notches are the leading cause of failure in injection molded thermoplastic parts. They are caused by the abrupt rise in stress concentrations in the internal sharp corners and are a function of the product design geometry, mold design, and the construction quality used in the mold. Figure 3-22 illustrates the failure problems caused by internal sharp corners and provides design recommendations to eliminate the internal sharp corner problems. All materials are notch-sensitive when subjected to high stresses at the notch and internal sharp corners result in crack failure. To avoid this problem, it is necessary to calculate the stress concentration factors for all internal corners and to check that all internal sharp corners are within safe stress limits. Figure 3-23 shows how to calculate the stress concentration factor of a cantilever beam and the ratio between the internal radius and the part wall thickness. With this graph, the internal radius dimension can be determined and should be approx. 50% of the part wall thickness. The internal radius should have a good surface ﬁnish, providing a streamlined path for the melt ﬂow, improving the impact strength, and resulting in easier ejection of the parts. The radii also give added life to the mold by reducing cavitation in the cavity’s metal surface. The minimum recommended internal radius is 0.031 in.

Good design Stress concentration factor

3.0

Figure 3-22 Internal sharp corner

Cantilever beam internal radius stress concentration factor graph

2.5

2.0

Cantilever beam P = Applied load R = Internal radius T = Wall thickness

P

Recommended

R. 1.5

0.2

0.5

0.6

0

0.4

1.0

Figure 3-23 Stress concentration factor vs. radius/wall thickness (Courtesy: Du Pont)

0.8

1.0

1.2

1.4

R.

T

Internal radius/part wall thickness (R/T)

3.4

Injection Molded Thermoplastic Bosses

Bosses are thermoplastic cylinders attached to a side wall or end corners. Special self-tapping screws are used to mount other components. The boss’s outside and hole’s inside diameters are based on size, depth and type of screws, pullout torque requirements, resin modulus of elasticity, creep, and boss weld line strength. Figure 3-24 shows operational problems when the wall of the boss is either too thick or too thin. The recommended design takes into account the melt ﬂow path around the correct boss wall and compensates for the low weld line strength with a rib. The use of large and long screws is not recommended; it is best to split the load using two or more smaller bosses to eliminate the heavy wall problems. Figure 3-25 shows the difﬁculties caused by large screws and provides design recommendations. When the boss is attached directly to a back wall or end corner molding, dimensional and breakage problems occur. Figures 3-26 and 3-27 show these operational problems, the recommended designs also shown eliminate these difﬁculties.

223

3.4 Injection Molded Thermoplastic Bosses

Boss wall too heavy Boss wall too thick

Poor design Poor design

Voids Hole size smaller

Warpage

Sink mark

Operational problems

Voids

Sink marks

Operational problems

Boss wall too thin

Poor design Good design split the load

Weld line failure

Figure 3-25 Thick boss walls cause problems that can be solved by using two smaller screws

Operational problems Correct boss wall thickness Rib Heavy cross sectional area Melt flow entrance

Voids

Rib

Sink marks

Neck down

Weld line

Poor design

Good design Figure 3-24 Boss wall thickness problems and recommended design

Good design

Operational problems

Figure 3-26 Molded bosses connected to a back wall

Heavy cross section area

Rib

Rib

Voids

Sink marks

Poor design

Molding problems

Figure 3-27 Molded bosses connected to an end corner

Good design

Good design

224

3 Structural Designs for Thermoplastics

3.5 Angled pins

Injection Molded Thermoplastic Threads

Internal and external threads can be manually or automatically molded into thermoplastic products; the speed cycle depends on the complexity of the mold. External threads are automatically molded by using the parting line of a two-plate mold or the parting line formed by angled pin slides as shown in Figure 3-28. Internal threads are molded by an automatic thread unscrewing device; the core moves forward for injection and retracts unscrewing the core from the molded product for ejection. Figure 3-29 illustrates the functional steps of this device.

Slides

Figure 3-30 illustrates a manual method used for threaded core inserts to encapsulate the threads where both items are ejected from the mold. Inserts are manually unscrewed for further use in the molding process.

Figure 3-28 External thread slides

3.6

Figure 3-29 Unscrew system

Manual core set-up

Gate

Collapsible Core for Molding Internal Threads

The collapsible core is a major breakthrough for injection molding thermoplastic products requiring internal threads, undercuts, and protrusions. This device incorporates only three moving parts, utilizing conventional mold movements. The collapsible core makes it possible to mold products that were previously considered impossible to mold. Products with internal protrusions, interrupted threads, and undercuts can be injection molded for high or low volume production applications. For conventional internal threaded products, the automatic operation of the collapsible core could reduce the injection molding cycle up to 30%. Root diameter of threads

Mold closed Core & part ejected Ejector pins

Cavity insert

Cavity

Molded cap Shut-off

Core insert

Stripper insert Collapsible core

Mold open

Collapsing segments

Figure 3-30 Manual process for molding internal threads

Positive collapse sleeve

Cooling hole

Ejector plate assembly

Figure 3-31 Collapsible core process for molding internal threads (Courtesy: DME)

Collapsible core open mold closed

Collapsible core closed mold open

225

3.7 Preferred Standard Thread Forms for Thermoplastics Figure 3-31 illustrates the collapsible core’s major advantages for injection molding thermoplastic quality end caps requiring internal threads without difﬁculty. The collapsible core makes it possible to design a simple mold that is easier to run in problematic product applications.

3.7

Preferred Standard Thread Forms for Thermoplastics

Several types of standard thread forms have been developed for common metal screw applications. The preferred types of thread forms for thermoplastic applications are those that have the largest internal radius with the load bearing face nearly perpendicular and with the highest depth thread. In the following, the standard thread forms are presented starting from the best to the worst for thermoplastic applications. The thread form selection will make a large difference in the performance of the thermoplastic part. The use of internal tapered (pipe) threads for any type of thermoplastic applications should be avoided. Buttress Threads The buttress threads are unique types of threads that are not symmetrical around a cross section center line; they are the preferred type of threads for thermoplastic applications. Buttress threads have advantages in load bearing applications where the load is in one direction only. Because the load bearing face is nearly perpendicular to the axis of the screw, the loads are transferred almost entirely along the axis rather than in the radial direction. The buttress threads have a standard root radius ranging from 0.035 to 0.070 × pitch. The largest possible root radius is suggested for thermoplastic products. Figure 3-32 shows the buttress thread details and applications. P (Pitch = 2 x screw dia./15) f (Flat = 0.125 x P) d (Depth = 0.75 x P)

45˚

r (Root radius from 0.035 to 0.070 x P)

Thread start (0.03 inch)

Thread end (0.03 inch)

Thread start (0.03 inch)

Figure 3-32 Buttress thread details and applications

226

3 Structural Designs for Thermoplastics (Depth of V-thread = 1.136 x P) s (Root & crest truncation = 0.26 x P) 47.5˚

r

d

s

H

British Association Standard Thread This thread form is recommended by the British Standards Institute for all screws smaller than ¼ in in diameter. This thread form, details are shown in Figure 3-33, is well suited for thermoplastic applications because of its thread face angle of 47.5° and its large root radius of 0.180 × pitch.

P r (Radius = 0.180 x P) (Depth of thread = 0.60 x P)

Figure 3-33 British Association standard thread details

(Depth = 0.640 x P)

The Whitworth thread form is an excellent thread design for thermoplastic applications, although this thread is being phased out in the industry. This type of thread, shown in Figure 3-34, has a generous radius of 0.137 × pitch at the root of each thread, which reduces the stress concentration effect.

American National or Unified Threads P

The uniﬁed thread shown in Figure 3-35 is the most common thread form used in designs. For a thermoplastic to metal thread joint, only coarse threads should be used to prevent cross threading and damage to the thermoplastic thread.

d 55˚

r (Radius = 0.137 x P)

Figure 3-34 Whitworth thread details

(Flat = 0.125 x P) P

f

Whitworth Threads

ACME Threads The ACME thread is used for power transmitting applications. This thread form, as shown in Figure 3-36, causes high stress concentration factors in the ﬂat root sharp corners, the hoop stress developed will concentrate at the weakest point and cause failure either immediately or over a period of time. This type of thread is not recommended for use for any type of thermoplastic applications.

d 60˚

Square Threads (Depth = 0.613 x P)

Figure 3-35 Unified thread details

(Depth = 0.5 x P) P

29˚ d

V-Threads The V-thread, as shown in Figure 3-38, is the worst screw thread form for thermoplastic applications. The sharp notch at the root of the thread causes extreme stress concentration factors (3.0 to 5.0), producing catastrophic product failure.

t (Width of thread = 0.5 x P)

Figure 3-36 ACME thread details

P

0.5 P

Square threads, as shown in Figure 3-37, are used for power transmission applications, such as jack power screws, because of their efﬁciency in transmitting power. They are not recommended for thermoplastic applications due to the stress concentrations induced into the base threads, causing the threads to shear off.

0.5 P

Figure 3-37 Square thread details

227

3.8 Injection Molded Products with Undercuts

3.7.1

Thermoplastic Threads Creep Effects

When threading assemblies between metal to thermoplastic, it is preferable to have the metal part external to the thermoplastic part. The male thread should be on the thermoplastic part. Careful considerations should be made in a metal/plastic assembly, because of the large difference in the coefﬁcients of linear thermal expansions of these materials. Thermal stresses created by this difference will result in creep of the thermoplastic part over a long time, if the assembly is subject to temperature ﬂuctuations or if the end use temperature is elevated. If the plastic part must be external to the metal, a metal backup sleeve may be needed.

3.8

Injection Molded Products with Undercuts

Injection molded thermoplastic products may have undercuts for functional reasons or for decorative effects. Undercuts in an injection molded product application will increase tooling costs and lead to longer injection molding cycles. When undercuts in a chamfered molded product are less than 5% of the diameter, it may be possible to strip or eject the product out of the mold cavity. However, if this solution is contemplated, the mold must be designed to operate so that ejection takes place only when the thin walled molded product is free to expand or compress. In such cases, it may be necessary to provide the mold with an ejection ring or plate, rather than ejector pins. The only suitable geometry for internal undercuts is circular. Other shapes, such as rectangular, can not be stripped because of the high stress concentrations. Several other mold undercut release techniques are available for thermoplastic products with undercuts greater than 5% of the diameter to avoid injection molding problems. Following are design suggestions for stripping undercuts: • The undercut thermoplastic product must be free to stretch or compress. • The undercut should be rounded and chamfered to permit easy slippage of the thermoplastic product over the core of the mold and to minimize stress concentration during the ejection stripping action. • Adequate mold ejection contact area should be provided to prevent penetrating or collapsing of thin walled thermoplastic product during the stripping action. • Some permanent deformation may occur when the undercut is stripped. The deformation depends on the thermoplastic product, mold design, and injection molding process variables. • Acetal Homopolymer Resins. It is possible to strip the thermoplastic products from the cavities, see Figures 3-39 and 3-40, if the undercuts are less than 5% of the diameter and are chamfered. Only circular geometries are suitable for undercuts. Other shapes, such as rectangles, have high stress concentrations in the corners which prevent successful stripping. Other methods should be used to obtain a satisfactory thermoplastic product for undercuts greater than 5%.

P

d 60˚

(Depth = 0.866 x P)

Figure 3-38 V-thread details

228

3 Structural Designs for Thermoplastics

Allowable internal undercuts (A - B) x 100 B

B

= % u ndercut

B A

A

Figure 3-39 Internal undercuts for injection molded products (Courtesy: Du Pont)

Allowable external undercuts

C B

(A - B) x 100

= % undercut

C

A Mold closed

C B A

Figure 3-40 External undercuts for injection molded products (Courtesy: Du Pont)

Core pin closed

Undercut Mold open Plastic part

Core pins mating face

Core pin open

Figure 3-41 Core pins internal undercuts

Core

Cavity insert

Plastic part

Ejector wedge

Ejected molded part

Figure 3-42 Ejector wedge internal undercut

• Unreinforced Nylon 6/6 Resins. Injection molded nylon 6/6 products having an undercut between 6 and 10% usually can be stripped from the mold. To calculate the allowable undercut see Figures 3-39 and 3-40. The allowable undercut will vary with thickness and diameter. The undercut should be chamfered to ease the removal from the mold and to prevent over-stressing of the molded nylon 6/6 product. • Reinforced Nylon 6/6 Resins. For glass reinforced nylon 6/6 resins, a collapsible core or split cavity undercut is recommended to minimize high stress conditions during the stripping action, see Figures 3-39 and 3-40. The undercut should be rounded and limited to 1%, if stripping from a 100 °F mold, or 2% from a 200 °F mold temperature. Internal Undercuts Using Core Pins and Ejector Wedge Internal undercuts can be injection molded by using two separate core pins or an ejector wedge that srips the thermoplastic product from the mold, as shown in Figures 3-41 and 3-42. The mold must be designed to permit the necessary deﬂection of the part as it is stripped from the undercut. This is a very practical method, but ﬂash must be controlled where the two core pins or the ejector wedge meet.

229

3.8 Injection Molded Products with Undercuts Internal Undercuts Using an Offset Ejector Pin

Mold closed Cavity insert

Figure 3-43 shows another method for molding internal undercuts using an offset ejector pin side movement, produced by the ejector plate assembly and the adjoining wall of the cavity. Offset ejector pins are used for internal side wall undercuts, but have limitations on the depth and size of the internal undercuts. UniLifter® Undercut Ejection System

Mold lower half cavity insert

This undercut releasing system uses standard components for simpliﬁed mold design and construction. Designs using angles from 5° to 10° will typically yield the best results. Angles up to 15° are permissible by using lifter guides in the bottom of the support plate, or whenever less than half of the core blades are bearing in the core insert. It is recommended that guided ejection be used in all designs.

Offset ejector pin Ejector plate assembly

Figure 3-44 illustrates the mold design and operation of the undercut releasing system.

Mold closed

Molded part

Offset ejector pin side movement

Ejection

Figure 3-43 Internal undercut, offset ejector pin side movement

Mold open

Figure 3-44 UniLifter ejector bar undercut (Courtesy: DME)

230

3 Structural Designs for Thermoplastics

Angled pin Heel block

Slide

Stop block Wear plate Slide retainer

Figure 3-45 External side undercuts using an angled pin slide

External Side Undercuts Using an Angled Pin Slide In this mold design, as shown in Figure 3-45, the side cavity block is attached to a carriage mounted in guides on the moving half of the mold. A suitable extension is machined in the ﬁxed mold half, adjacent to the cavity, to accommodate the side cavity assembly when the mold is closed. This mold design offers a positive lock applied by the locking heel to the side cavity assembly when the mold is closed. External Side Undercuts with Slides Actuated Hydraulically Figure 3-46 shows that both slides are actuated hydraulically and this system is not dependent on the opening movement of the mold. The slides in this mold design can be operated automatically by the operating control panel of the machine. This mold design does not rely on the hydraulic locking force alone to keep the slides closed during the melt injection in the mold cavities. It is better to use a locking plate assembly for large side projected area slides to avoid the use of heavy duty larger diameter hydraulic cylinders. However, if the total projected area of the undercut is relatively small, this system should be a good design option, the mold design and construction is considerably simpliﬁed.

Locking plate

Mounting plate Slides

Wear plate Hydraulic actuator

Slides

Core

Ram

Figure 3-46 External side undercuts, hydraulic actuation of slides

231

3.8 Injection Molded Products with Undercuts External Radial Undercuts with Angled Pin Slides In this system, two hardened steel angle pins mounted in the ﬁxed mold plate, control the slide cams. The slides are on the moving mold plate. Figure 3-47 shows a typical mold used to produce bobbins (spools); the top illustration shows the slides in the closed position. As the mold opens, the angled pins force the external cavity slide cams to move outwards, sliding on the mold plate as shown in the middle illustration. Once the angled pins move further out from the slide cams (mold fully open), the slide cam’s movement ceases immediately. The moving half of the mold causes the ejector sleeve system to operate and the molded bobbin is ejected from the mold automatically.

Mold closed, locking plate on both slides

Intermediate position of slides

Mold open, sleeve ejects the bobbin

Figure 3-47 External undercut with angled pin slides actuation

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3 Structural Designs for Thermoplastics

3.9

Injection Molded Integral Life Hinges

The injection molding of integral hinge offers a unique concept that has been successfully applied for the production of a variety of containers, cases, and similar products. Figure 3-48 illustrates the concept of producing complete integral hinge enclosures in one single operation, capitalizing on the functional advantage and an economical incentive in the manufacturing of the products.

180 ˚ Integral hinge, open position

Multi-part design is the main reason for implementing this technique in articles containing one or more integral hinges. However, this concept is considered difﬁcult because of the complex interaction of the many factors inﬂuencing the functional properties, and in particular the appearance, of the integral hinge enclosure. This section attempts to give an insight into the possibilities for, and the limitations of using injection molded thermoplastic integral hinge products. Considering the numerous possibilities in the product design and types of applications, it is obvious that the overall design aspects can only be discussed in general terms. The integral hinge is sometimes recognized as a separate device.

180 ˚ Integral hinge, closed position

90˚ Integral hinges, open positions

Figure 3-48 Injection molded integral life hinge applications

One aspect that should be considered is that the hinge acts as a second gate for the part of the box behind the hinge. The thin section formed by the hinge web (thickness between 0.001 and 0.015 in) causes a considerable pressure loss during the injection ﬁlling process of the hinge and lid. Rectangular or square boxes provided with straight sidewalls are prone to distortions with inward bending of the sidewalls of both box and lid. Consequently, the lid of the box, which is almost always the part behind the hinge, is very sensitive to the occurrence of mold shrinkage defects. It is therefore advisable to select a wall thickness for the lid that is smaller than for the box. The extent of the reduction in wall thickness in the hinge limits the possibility of ﬁlling the lid cavity. For the same reason, incorporation of ribs, bosses, etc. in the lid should be avoided. Even with correct gating, sink marks occur on the lid’s surface opposite the ribs and bosses. It is difﬁcult and sometimes impossible to prevent these sink marks. The best remedy is to hide this defect is to use a surface texture. The following factors affect the quality of molded integral life hinges: • Geometry and size of the injection molded life hinge • Product wall thickness and melt ﬂow distribution • Presence of melt ﬂow dividers and/or ribs • Locking mechanism used in the product • Surface ﬁnish required for the product • Life hinge design, quantity, location, and length of the hinge • Life hinge mold design, number of cavities, size and location of the gates, and mold cooling system • Processing conditions, type, conditions, and size of the injection molding machine used • Type of thermoplastic resin used

233

3.9 Injection Molded Integral Life Hinges

3.9.1

Injection Molded Integral Life Hinge Design

For injection molded integral life hinge applications requiring a long ﬂex life, it is essential that in designing the hinge section, optimum orientation of the polymer melt is achieved. The melt viscosity of the polymer used and the processing conditions are important parameters that inﬂuence the overall design of the life hinge and mold. Consequently, these factors set corresponding limits to the web thickness and land length of the hinge. As thickness does not inﬂuence the ﬂex life of a correctly designed hinge, this dimension is mainly chosen to suit the required stiffness and the angle through which the life hinge enclosure must operate. The hinge should be thin enough to allow easy bending through the angle of normal use, but thick enough to retain a sufﬁciently stiff connection between the box and the lid enclosure. For containers and cases that normally operate at zero load, the hinge web thickness should range from 0.010 to 0.015 in. For light duty applications, a hinge web thickness of 0.015 to 0.020 in can be used. For long hinges, the thickness should be kept within narrow tolerances in order to prevent adverse effects of nonuniform melt ﬂow through the hinge section. In general, it can be said that for 0.010 to 0.015 in thick hinges, the tolerances should be kept at ±0.0008 in. With a continuous hinge, the memory effect often results in a generally unwanted self-opening tendency of the enclosure. This tendency increases with thicker hinges and greater length. For some applications, this self-opening tendency can be an advantage, provided that it does not interfere with the container’s closing. An example where this characteristic is an advantage is the lid springs that open by a gentle pressure on the front of the enclosure. The land length should be properly dimensioned. Too short a land length may lead to insufﬁcient back pressure generated during the injection mold ﬁlling. This in turn may result in nonuniform ﬂow of the melt through the hinge. Weld lines may occur in the hinge section that will lead to premature failure of the hinge. On the other hand, too long a land length will cause a high pressure drop with the risk of shrinkage defects in the lid or even short shots. The land length of the hinge should be at least three times its thickness. If the ratio is less than 3 : 1, weld lines are likely to occur in the hinge, ﬂexing will create considerable stress, and the ﬂex life of the hinge will be reduced considerably. As a general guide, a land length of 0.030 to 0.060 in should be provided; however, this can be increased if greater angular movement is needed. Preferably, the inner shoulders of the hinge should be slightly recessed in order to ensure optimum ﬁt of the box and lid at the closing line. Because of necking-in of the polymer, hinge thickness and land length are slightly changed when the hinge is ﬂexed for the ﬁrst time. Stress concentration initiated by internal sharp corners or notches near the hinge area should be avoided. Therefore, all corners of the hinge area should be well rounded, with the hinge surfaces polished and smooth. This aspect should be carefully considered when, for instance, the hinge surfaces are provided with a leather grain or other types of surface textures. A minimum shoulder radius in the hinge of 0.015 in is recommended. Figure 3-49 shows a faulty hinge design often encountered in practice. In addition to the presence of sharp corners, this design shows the disadvantage that the hinge will most probably not bend along the desired line and/or bending will occur

0.01 to 0.015 inch

Box

Lid

Figure 3-49 Incorrect life hinge design

234

3 Structural Designs for Thermoplastics Lid

Only suitable for non critical applications 0.01 to 0.015 inch

0.01 to 0.015 Box

Lid

Figure 3-50 Incorrect life hinge design

0.03 R.

0.010 to 0.015 R. 0.03 to 0.06 Box

0.010 0.015 0.010

0.015 R.

Figure 3-52 Good angled polypropylene life hinge R. 0.03 R. Box

Recess for closing 0.010 0.015

0.02

0.01 0.063

0.03 R.

Box

Lid

1˚ Draft angle per wall (Typ.)

Figure 3-51 Good standard polypropylene life hinge

Lid 0.04

0.04

along different lines. This will result in a twisted hinge, poor closing of the box, and premature failure of the hinge. The design in Figure 3-50 is better, but is not recommended for long hinges and enclosures requiring relatively wide hinges. This type of hinge is only suitable for short hinges in noncritical applications.

0.08

Figure 3-53 Medical tablet dispenser life hinge design

Figures 3-51 and 3-52 are typical illustrations of polypropylene life hinge designs recommended for boxes and cases. Closing direction

H

0.010 0.015 0.020 to 0.040

L > 2H Lid Box

Figure 3-54 Video cassette case life hinge design

Box

Instead of using one long hinge it is sometimes better to use an interrupted hinge, reducing the self-opening tendency. It should be kept in mind, however, that a noncontinuous hinge gives rise to the occurrence of weld lines in the lid. Where extra resistance to torque or tearing is required, the thickness of the hinge web may be increased to 0.020 in. All transitions in the hinge web thickness should be adequately radiused (minimum 0.015 in). Figure 3-54 shows in principle the hinge design that is being successfully used for storage cases of video cassettes. The advantage of this design is that the gap between box and lid is masked by the hinge itself.

Closing direction 0.03 R. 0.06

A characteristic inherent in the use of polypropylene integral hinges is an imperfect ﬁt of the box rear wall and lid. It is difﬁcult to achieve an exact ﬁt, therefore the rear wall of the closed box shows a slight gap between the box and lid beside the hinge. This defect can be minimized and often eliminated by providing a slight recess (approximate 0.010 in) in the rear wall of the box part where the hinge section is located as shown in Figure 3-53.

0.047 0.010 Lid 0.015 Total flex angle 180˚

Figure 3-55 Medicine box life hinge design

The life hinge construction illustrated in Figure 3-55 is a small box for medicine tablets. All sidewalls, including the rear wall of the box, are convex to prevent distortions. Because the hinge should be straight, the convexity of the rear wall affects the overall hinge design. The hinge will extend outside the rear wall of the box to a greater extent than with a straight rear wall.

235

3.9 Injection Molded Integral Life Hinges

3.9.2

Mold Design Considerations for Hinges

Lid

The most important factor in mold design is the gating system. The selection of the number, size, and location of the gates is governed by technical, aesthetic, and economic considerations. However, the following conditions should be fulﬁlled: • The gate or gates should be located on the largest section of the enclosure, i.e., on the normally the box. • The major cavity should be completely ﬁlled before the melt ﬂow reaches the restricted hinge area.

Box Figure 3-56 Film gate in box front wall

• Polymer ﬂow in the hinge section should be perpendicular to the hinge line and should be as even as possible. This also applies when a multiple gate system is used. • With multiple gate systems, gates should not be located on either side close to the hinge, in order to prevent the formation of weld lines in the hinge. The ideal method of gating for shallow boxes is to use either a ﬁlm gate or two individual edge gates located in the front wall of the box as shown in Figures 3-56 and 3-57. Especially with the ﬁlm gate, mold ﬁlling is very uniform, resulting in optimum ﬂow of the melt through the hinge. In practice, however, neither of the systems is favored, because their location in the front wall will obstruct the construction of the locking mechanism of the box. In addition, with single cavity molds, both systems require the use of a hot runner or insulated runner mold, resulting in a relatively costly mold. The simplest and most suitable gating system applicable to both shallow and deep boxes is to locate the gate(s) at the bottom of the box. Figure 3-58 shows the gate close to the hinge area causing difﬁculty in ﬁlling the part; this gate location is not recommended. The disadvantage of this gate location is that the melt reaches the hinge section before the ﬁrst cavity is ﬁlled, restricting uniform melt ﬂow distribution and melt ﬂow orientation through the hinge. Because of the resultant higher pressure loss, a greater amount of shrinkage defects in the lid of the box can be expected. The choice between a single or a multi-gate system is mainly a matter of experience. In order to ensure that the ﬁrst cavity is completely ﬁlled ﬁrst, the gate for shallow boxes should be located as shown in Figure 3-59. With increasing box height, the gate can be moved toward the long axis of symmetry, which coincides with this line for deep boxes as shown in Figure 3-60. Figure 3-61 shows the correct location of the gates for a relatively shallow box. It can be seen that for this gating system the same principles apply as for the single gate system. When using multi-gate systems, attention should be given to balancing the melt ﬂow to minimize the formation of weld lines. Sometimes incorporation of thicker sections in the lid (e.g., ribs or bosses) is unavoidable. Especially with relatively large and high boxes, undesirable shrinkage defects in the appearance surface are likely to occur. As shown in Figure 3-62, it is possible in these cases to gate the article on both sides of the hinge. Attention must be given to the weld line formation and that ﬁlling of the ﬁrst part of the lid takes place from the cavity forming the box in order to ensure the quality of the hinge. This can be achieved by means of a multiple gate system in the box and a single gate in the lid.

Lid Weld Line

Box Figure 3-57 Two edge gates in box front wall

Box, integral hinge and lid Lid

Gate

Box

Gate Figure 3-58 Hinge, poor gate location

236

3 Structural Designs for Thermoplastics Gate locations for weld line strength

Deep box

Box Gate (A) 1 2

Gate

Shallow box

Lid

Gate Gate (A)

Lid

Lid

1 4 1 3

2 3

Figure 3-61 Hinge, good multi-gate system on box

2/3

1/3

Box Gate locations for weld line strength Box Lid

Gate Figure 3-59 Hinge, good gate location

Box ½

½ Gate

Figure 3-60 Hinge, good gate location

Gate (A) 1 2

Gate (A)

Gate (B)

1 4 1 2

2 3

Figure 3-62 Integral hinge good multi-gate system on box and lid

Short hinge (1)

Box Gate Weld line

Lid

Gate

Short hinge (2)

Figure 3-63 Two gates on box with two short life hinges

Gate

Gate

In those cases where it is desirable to reduce the self-opening of the lid arising from the memory effect of a continuous hinge section, interrupted small hinges can be used as shown in Figure 3-63. With respect to the location of the gates, the same rules are applicable here as discussed for continuous hinges. Because noncontinuous hinges give rise to the formation of one or more weld lines in the second cavity, special attention should be given to the correct balancing of the individual sections of the hinge and venting of the mold. Occasionally, opening and closing of the box is accomplished by means of a two-hinge construction as shown in Figure 3-64. In this case, the gate(s) should be located in the central section forming the rear wall of the box. Wall thickness of this section should be approximately 25% greater than the wall thickness for the box and lid.

3.9.3

Proper Gate Design for Life Hinges

In this section, some other aspects regarding the design of the mold, as well as the effect of processing conditions and polymer choice in the performance of hinged boxes are discussed.

Long hinge (1)

Box

The size of the box gates (A) should be equal and considerably greater than that used for partial ﬁlling of the lid (B). Moreover, the lid gate should be located at the maximum possible distance from the hinge. The mutual sizes of the gates should be established by trial and error and can be checked from the location of the resultant weld line in the lid and by producing short shots.

Small shallow boxes used for the packaging of medicine tablets, cosmetics, etc. are most economically produced in multi-cavity molds. In these cases, proper gate design is very important. Lid

Long hinge (2)

Figure 3-64 Two gates on base with two long life hinges

In practice, it can be seen that when using a two or four cavity mold layout, the gate (usually an edge gate) is located in the sidewall close to the hinge section as shown in Figure 3-65. This system is only suitable for boxes of optimum design and should be avoided for critical designs, i.e., boxes with thin straight side walls requiring a relatively long ﬂex life of the hinge. With this gate location, the resistance to ﬁlling the ﬁrst cavity may become greater than the resistance to

237

3.10 Conventional Types of Pin Hinges ﬂow through the hinge part located near the gate. This will obviously result in uneven ﬂow of the melt over the whole length of the hinge and create the risk of producing warped boxes, showing a very limited ﬂex life of the hinge due to occurrence of weld lines.

Weld line fail

Box

Lid

The best gating system with multi-cavity molds should be considered for each individual case. Sometimes it is possible to locate the gates in the center of the front wall of the boxes, but generally a system (as shown in Figure 3-66) gives the best compromise between all conﬂicting requirements. Edge gate

Because of their well known advantages, edge gates are preferred to other types of gates for multi-cavity mold layouts. Moreover, by using an edge gate, any shrinkage defects occurring in the lid or the box can be minimized. This type of gate enables a more ﬂexible control of the injection ﬁlling rate. The correct size of the edge gate should be dimensioned by having a gate land of 0.040 in and a gate depth of 50% of the box wall thickness; the gate width is established by trial and error, depending on the polymer melt viscosity. A good procedure is to start with a small gate; the size can then be enlarged as necessary.

Lid

Box

Lid

Runner

Edge gate

Lid

3.10

Box

Box

Sprue

A sprue gate is used in exceptional cases, e.g., for high and relatively narrow boxes where fast ﬁlling is required in order to minimize the pressure loss in the gated cavity (long melt ﬂow path) so that adequate pressure will be available to ﬁll the lid cavity.

Figure 3-65 Hinge enclosure incorrect edge gate location

Box

Conventional Types of Pin Hinges

Injection molded thermoplastic hinges can be inexpensive and practical. The main idea behind their design is to make the products as simple as possible and to avoid using metal hardware. There is a trade-off between the intricacy of a mold and the cost of assembly and ﬁnishing. For production volumes of more than 100,000 parts per year, it is deﬁnitely worthwhile to build an intricate mold that eliminates all the ﬁnishing processes. For small volume productions, trial runs and prototypes, the mold should be simple. The molded products can be drilled, heat staked, and hand ﬁtted together. Since there is sliding between mating parts, the hinge acts as a bearing. If the hinge is to be truly durable, at least one side of it should be made of nylon 6/6 or acetal homopolymer materials. The following methods of hinge formation do not require metal hardware in the hinge and are suitable for mass productions. Insert Molded Post Hinges A post made of nylon 6/6 can be insert-molded into a different material to form a strong and precise hinge. The nylon 6/6 post must be perfectly round and very smooth. The hinge will have enough friction to stay in any open position. With the proper selection of material, a hinge with an insert-molded post can last practically forever. Side Core Insert Mold Hinge Design If a small box has only two posts and socket hinges, the box can be molded with the posts interlocked in the sockets by using side cores in the mold as shown in Figure 3-67. Lid and box are molded in an interlock position. The thin steel sleeve from core separates the shaft on the lid from the bearing on the box.

Figure 3-66 Runner and mold layout, four hinge enclosures

Lid

238

3 Structural Designs for Thermoplastics

Lid Snap-in pins into side holes

Lid

Core

Core Core Box

Figure 3-68 typical snap-in pin hinge used in audio cassettes

Box

Lid and box are molded in an interlock position.The thin steel sleeve from core separates the shaft on the lid from the bearing on the box. Figure 3-67 Small box with two posts and lid with socket hinges

Snap-In Pin Hinge Design In many injection molded and blow molded cases, snap-in hinges work quite well. Figure 3-68 is a typical snap-in pin hinge used in audio cassette and compact disk cases made from polystyrene. These types of hinges can be molded so that they snap together during assembly. Sometimes, they consist of a pin that is supported at both ends molded into one part, and a “C”-shaped receptacle or semicircular opposing ﬂanges with interference so that the sides are deﬂected by the pin as it is inserted as shown in Figure 3-69.

Snap-in pins into holders

Standard Lug and Pin Hinge Design Figure 3-69 Enclosure case with rear snapin hinges

Top

Bottom

There are a variety of pin and hinge designs that can be injection molded into thermoplastic parts. Alternating ﬂanges with semicircular section geometries can be used in place of a tube so that slides and cams in the tool will not be needed. Pins are often molded with ﬂat sections, which are beyond the normal range of rotation so that the pin can be inserted from the side. It is also possible to mold alternating semicircular ﬂanges into both parts and then insert a separate metal or thermoplastic pin through the center. Pin and hinges can be designed so that the tube slides down over the pin and is locked in place when the hinge is closed but can be lifted off when the hinge is all the way open. This is often desired when cover removal is necessary for service of the internal components as shown in Figure 3-70. Co-extruded Thermoplastic Hinge Design

Pin

Engineering thermoplastic elastomers are process compatible with a number of other thermoplastic materials and adhere well to materials such as polyvinyl chloride. This fact suggests the possibility of using co-processing to arrive at designs that incorporate the properties of more than one thermoplastic. Coextrusion and insert molding are two obvious possibilities.

Figure 3-70 Standard lug and pin hinge

In some situations, the use of a small amount of engineering thermoplastic elastomer may provide the most cost effective solution to a design problem. An examples is the co-extruded hinge shown in Figure 3-71, which incorporates a ﬂexible ﬁlm as a hinge of copolyester thermoplastic elastomer with two rigid support sections of PVC.

Lugs and pin hinge cross section Top

Bottom

239

3.11 Metal Inserts for Thermoplastic Encapsulation

Copolyester hinge

Blind hole

Open hole

Blind hole counterbored

Blind hole protruding

PVC

Figure 3-71 Thermoplastic co-extruded hinge designs (Courtesy: Du Pont)

The high tear resistance, outstanding ﬂex life, and excellent low temperature properties of copolyester thermoplastic elastomers performing as thermoplastic hinges creates opportunities that were previously beyond the capability of thermoplastics. Also, this efﬁcient hinge design combines the different properties of each material, enabling different kinds of functions to be included in a hinge made in a single operation.

3.11

Metal Inserts for Thermoplastic Encapsulation

Protruding shaft

Protruding screw

Figure 3-72 Metal threaded and shaft inserts for plastic encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

The use of inserts in the injection molding process of thermoplastics presents some difﬁculties. Where inserts are required to add strength to the hold down screws, adding life to the screw thread, or when covers must be frequently removed for electrical conduction or other reasons, it must be realized that the encapsulation of inserts will slow down the molding cycle. It also requires a special type of injection molding machine, different mold and processing conditions, all adding to the manufacturing cost of the product. However, good design practices demonstrate that some types of inserts may be installed after molding. These metal inserts may be added by automatic means at a rate faster than is possible by incorporating such parts in the molding operation. For example, a tapped hole in the thermoplastic with a machine screw can be replaced with a self-tapping screw, with resultant savings in the cost of the injection molded products.

Drawn-pin

Drawn-shell

Several kinds of inserts are used in the injection molding process, including, metal inserts made by machining, by metal cold forging, by sheet metal stamping, and by sheet metal drawn processes. Several types of metal insert designs are shown in Figures 3-72 and 3-73. Maintaining a uniform accuracy in various dimensions of metal inserts has always been a problem for the plastics injection molding industry and for the metal insert manufacturers, mainly because of the lack of design information and standardization of metal insert dimensions.

Drawn-eyelet Figure 3-73 Metal drawn inserts for thermoplastic encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

240

3 Structural Designs for Thermoplastics Table 3-1 Machined Metal Screw Insert Tolerances

K E

45˚ B1

F C

H

B

A

G

D J

K E 45˚

Screw type

Description

Tolerance

A A1 A2

Tap size “American National” – “Class 2” Major diameter Minor diameter “regular” tolerance Minor diameter “precision” tolerance

(±) 0.002 in (±) 0.0005 in

B B1

Minor depth and major length diameter Number unusable threads top and bottom

C C1

Length “regular” tolerance Length “precision” tolerance Length of body male insert

(±) 0.010 in (±) 0.001 in (±) 0.010 in

D

Thread chamfer

45° · 0.003 in

E

Body chamfer

45° · 0.06 in

F

Coarse diamond knurl

G

Length of sealing diameter minimum

0.03 in

H H1

Length of usable thread 1.5 × diameter Length of usable thread

(H1 + B1 = B)

J

Sealing diameter tolerance

(±) 0.002 inch

K

Minimum bar stock diameter

F G

C1

B

H1

B2

D

A2 A1 J

Figure 3-74 Machined metal threaded insert tolerances (Courtesy: The Society of the Plastics Industry, Inc.)

3.11.1

Machined Metal Threaded Insert Tolerances

Dimensions and tolerances for typical metal female and male inserts are shown in Figure 3-74 and in Table 3-1. This data base used for machining as a single operation on an automatic metal screw inserts machine was compiled by the National Screw Machine Products Association. For dimensions A-2 (minor diameter) and C (overall length) the maximum “standard” tolerance should be speciﬁed whenever possible. However, for closer tolerances,“precision dimensions” can be speciﬁed when necessary. To maintain the “precision” tolerance, reaming and other additional operations will be necessary, at additional metal insert cost. If steel inserts are required and the recommended tolerances cannot be used in design without several modiﬁcations, the cost of this special steel insert will be higher than the common inserts made of brass, copper or aluminum. The minimum wall thickness of metal inserts depends on the accuracy of the insert mold shut-off dimensions. The stress produced by the thermoplastic mold shrinkage, combined with the injection molding pressure, may collapse the thin wall of the metal insert so that the mold shut-off diameter will be out of speciﬁed tolerances.

3.11.2

Thermoplastic Boss Wall Thickness for Metal Inserts

The thermoplastic boss wall thickness required around inserts depends on: • The proper metal insert design for thermoplastic encapsulation • Type of resin (thermoplastic or thermosetting)

241

3.11 Metal Inserts for Thermoplastic Encapsulation • The type of resin within each group • The mold shrinkage rate of the resin • The modulus of elasticity (rigidity) of the resin • The coefﬁcient of linear thermal expansion of the resin • The coefﬁcient of linear thermal expansion of the metal insert • End use temperature under which the product will have to function • Product dimensional changes caused by the post mold shrinkage and the moisture absorption characteristics of the resin • The resin creep characteristics (loss of properties) caused by aging of the product under load Very often, the thermoplastic injection molded product is designed ﬁrst and the important metal inserts then ﬁtted into the remaining space. If metal inserts are required for the application, they should be designed ﬁrst, and then the thermoplastic encapsulated product should be designed around the metal inserts. The shape, geometry, and requirements for the metal insert control the thermoplastic wall thickness to a high degree, especially when the metal inserts are of irregular shape (rectangular, square, hexagonal, star, or of any other shape having sharp corners). The most important properties of the resin are: the modulus of elasticity, the viscosity characteristics for ﬁlling and packing the mold cavity with an optimum injection pressure, a longer melt crystallization rate, good melt adhesion characteristics, high elongation rate, and low polymer melt/ mold processing temperatures. All these resin characteristics are very important for encapsulating metal inserts without cracking or moving the insert. No one property of the resin will solve the encapsulation molding problems. For instance, if using a resin having a low mold shrinkage rate of 0.002 in/in but with a low elongation rate and a high modulus of elasticity (rigid), the encapsulated molded product will crack. Resins with a higher mold shrinkage rate of 0.010 in/in but with a good elongation rate are capable of being stretched around the metal insert and will not crack despite a thin boss wall thickness. It is difﬁcult to set up thermoplastic boss standards (wall thickness around the insert and for the wall located underneath the insert) in relationship to the outside diameter of the metal inserts. Every encapsulated product presents different problems and must be engineered according to the design of the metal insert and the thermoplastic resin used. Figures 3-75, 3-76, and 3-77 show the metal insert orientation, the minimum clearance, and the problems caused by improper base wall thickness underneath the metal inserts. Parting line

Parting line

0.12 inch min.

Minimum distance between the inserts to avoid mold heat spots Figure 3-76 Inserts minimum clearance

Thin base wall

Blisters, poor surface finish

Thick base wall

Parting line

Voids

Good design (perpendicular)

Poor design (oblique)

Fair design (parallel)

Figure 3-75 Mold parting line orientation for metal inserts

Sink mark

Figure 3-77 Insert base wall thickness problems

242

3 Structural Designs for Thermoplastics Table 3-2 Plastic Boss Dimensions for Metal Insert Encapsulation B D

T

Metal insert Outside diameter “D” (in)

Unreinforced thermoplastic

Reinforced thermoplastic

Boss dia. “B” (in)

Thickness “T” (in)

Boss dia. “B” (in)

Thickness “T” (in)

0.156 0.187 0.218 0.250 0.281 0.343

0.312 0.344 0.406 0.453 0.484 0.578

0.062 0.062 0.078 0.078 0.093 0.093

0.312 0.360 0.422 0.500 0.531 0.609

0.062 0.062 0.078 0.078 0.093 0.109

Table 3-2 shows recommendations for round metal inserts (common insert sizes), the boss diameters, and the wall thickness for the section underneath the insert, required for the thermoplastic resins.

3.11.3

Press/Lock Slotted Metal Insert Installation After Molding

It is often more economical to assemble metal inserts into the product after its removal from the mold than encapsulating the metal inserts during the molding operation. Press/lock slotted metal inserts should be installed promptly after the molded product is removed from the mold (within 3 minutes, preferably immediately), in order to take advantage of the subsequent mold shrinkage of the thermoplastic, which will promote insert anchorage. Usually, the inserts are placed and pressed in with the aid of a suitably designed ﬁxture mounted in a small arbor press or punch press; allowing more than one insert to be be pressed in simultaneously. When an insert is to be pressed in after molding, the mold must be designed to provide a hole to receive it; the size of this hole must be designed on the basis of the size of the insert, the normal mold shrinkage of the thermoplastic, and any further allowance necessary to ensure anchorage. For small metal inserts, the diameter of the hole should be designed to be 0.001 in to 0.002 in smaller than the diameter of the insert, after completing the total shrinkage or when the product is at room temperature. For large metal inserts, and particularly when there is only a thin wall of thermoplastic surrounding the insert, an extra allowance for anchorage is not needed; on the contrary, the diameter of the hole may have to be designed on the basis of as little as 50% of the normal mold shrinkage factor in order to prevent cracking of the thermoplastic around the insert.

243

3.11 Metal Inserts for Thermoplastic Encapsulation Inserts with either coarse diamond or straight knurls can be pressed in satisfactorily. For some special applications, a coating of adhesive/sealer may be applied to the metal insert to promote anchorage and to ensure an airtight joint. For leakproof joint applications, the use of an “O” ring around the insert is recommended. A new type of insert (illustrated in Figure 3-78) has extended the practice of pressing in inserts. It can be pressed in with a single punch at the press, without the use of special ﬁxtures. It reduces rejection rate and increases the rate of production. This press/lock slotted metal insert consists of an internal threaded brass circular shield ring insert with four slots/knurled ears and a spreader screw. This brass insert can be fed by hand or by machine into the thermoplastic molded hole. The spreader screw is then threaded downwards, putting the screw pressure on the four slots/knurled ears’ surfaces. The slots expand the shield against the wall of the hole, without harming the internal threads. When the spreader screw has reached the end of its travel, the brass insert is locked into position by a shoulder on the inside of the insert shield.

Spreader screw

Slot ear insert Molded hole

The diameter of the hole to receive the press/lock insert should be 0.002 in greater than shield size. The bottom of the hole, upon which the insert rests, must be ﬂat. As with all inserts, maximum torque strength is achieved when the screw is long enough to utilize all of the threads of the insert and at the same time short enough so that the underside of its head comes down tight on the surface.

3.11.4

Cold Forged Metal Inserts for Encapsulation

Large volumes of cold forging inserts are needed to compete economically with other processes. Secondary operations are required for this process, such as turning, drilling, tapping, and others.

Plastic part

Pressed-in slot ear insert

Shield

The tolerances of cold forging inserts are the same as those for screw machined inserts, when machining, drilling, reaming, or tapping is involved. There are no speciﬁc formulas to control the diameters and widths of the collar’s dimensions with respect to the shank, or controlling the external geometry, such as ribs, shut-off shoulder, etc., which may be combined with other symmetrical or asymmetrical shapes in one piece. The solution for each problem should be reached through cooperation between the product designer, tool engineer, process engineer, manufacturing, resin supplier, and the insert manufacturer. Types of Materials Used for Cold Forged Inserts

Slot ears closed Spreader screw expands slot ears

Almost any metal can be cold forged, but the following materials are preferred in the order they are listed: • Aluminum and aluminum alloys • Brass • Copper and copper alloys • Carbon steels • Alloy steels • Stainless steel

Slot ears locked

Figure 3-78 Press/lock slot ears insert post molding installation

244

3 Structural Designs for Thermoplastics Tolerances without Finishing Operations Tolerances for any dimension such as length or diameter vary with the material and with the sizes and proportions of the insert, as they in turn determine the equipment or method of heading to be used. In general, the following tolerances can be considered for commercial inserts without needing ﬁnishing operations, although in some cases, special care must be exercised to meet the required speciﬁcations: Length shoulder Radii (R.) Sealing diameter Sealing diameter draft

(±) 0.010 in 0.03 in (±) 0.020 in (±) 1°

(maximum) (minimum) (minimum) (maximum)

Tolerances with Finishing Operations Figure 3-79 Commercial threaded female metal inserts for plastics

Whatever tolerance is needed can be met by adding ﬁnishing operations. For example, aircraft studs, bolts, and special components are commonly made in production today to tolerances as close as (±) 0.0005 in.

Sharp corners

3.11.5 Shrink stress

Poor design

Sharp corners

Shrink stress

Threaded Female Metal Inserts

Designing injection molded thermoplastic products requiring threaded female metal inserts for encapsulation requires a lot of planning and preliminary engineering work. The product design engineer, tool engineer, process engineer, manufacturing, resin supplier, and the insert manufacturer must cooperate to obtain simplicity of design, which will result in the production of satisfactory products and will promote economical production. Figure 3-79 shows four typical commercial threaded female metal inserts used with thermoplastic products. Typical applications for these inserts are electrical bobbins, purge solenoid magnetic spools, and ABS sensors; leads and insert connections are encapsulated inside of the plastic. Figure 3-80 shows three designs that could affect the performance of the thermoplastic product encapsulating one of these metal insert designs. Differential Shrinkage Effects on Encapsulation

Fair design

Figure 3-81 shows the after-molding effects on encapsulation caused by the high mold shrinkage of the thermoplastic resins and the small dimensional changes of the metal insert.

Metal insert

R.

R.

Plastic part

Differential shrinkage

Good design Figure 3-80 Threaded female metal insert design variations

Figure 3-81 Encapsulation effects caused by differential shrinkage

245

3.11 Metal Inserts for Thermoplastic Encapsulation Coefficient of Thermal Expansion Effects on Material Selection The correct selection of metals for inserts is essential, because of the differences in the coefﬁcient of linear thermal expansion between the various metals and thermoplastics. The values for the coefﬁcient of linear thermal expansion for ﬁber glass reinforced thermoplastic resins are controlled by the following parameters: • Fiber glass orientation during the injection molding process • Base polymer thermal properties • Additives compounded with the polymer • Process/equipment required to compound the polymer The coefﬁcient of linear thermal expansion for a polymer oriented parallel to the melt ﬂow direction (also known as MD) is generally the smaller value reported by the plastics suppliers. The ﬁber glass becomes chemically coupled and anchored with the polymer and provides reinforcement to the polymer, thus restricting changes of dimensions of the molded part. However, these beneﬁts are only applicable to sections of the molded part that have been aligned towards the MD direction. The coefﬁcient of linear thermal expansion increases with the ﬁber glass orientation angle changing from parallel or 0° (MD minimum) to perpendicular or 90° (TD maximum).

Table 3-3 Coefficient of Linear Thermal Expansion (0 °C to 200 °C) – (cm x cm/°C)

Materials

Coefﬁcient · 10–6

Engineering Thermoplastic Nylon 6/6, unreinforced Nylon 6/6, FR. unreinforced Nylon 6/6, reinforced Nylon 6/6, FR. reinforced Nylon 4/6, FR. reinforced HTN Nylon, reinforced Nylon 6/12, reinforced PBT, reinforced PET, reinforced PPS, reinforced LCP, reinforced

70 MD – NA TD 60–70 MD – NA TD 20 MD – 55 TD 20 MD – 55 TD 25 MD – 60 TD 15 MD – 45 TD 23 MD – 50 TD 25 MD – 60 TD 25 MD – 60 TD 22 MD – 55 TD 14 MD – 40 TD

Metals Aluminum 2S (99.2% Al) Brass (67% Cu, 33% Zn) Bronze (90% Cu, 10% Zn) Phosphor bronze Phosphor bronze 30 (95.5% Cu, 4% Sn, 1% Zn) Copper (99.9+%) Steel (99% Fe, 1% C) Cold roll steel Stainless steel (90.2% Fe, 8% Cr, .4% Mn, .12% C) Monel (60% Ni, 12% Fe, 11% Cr, 2% Mn) Nickel Silver (92.5% Ag, 7.5% Cu) Solder, half and half Zinc (95% Zn, 5% Al)

23.94 18.5 18.8 16.8 18.90 17.71 12 14 11 14 12.9 18 24 28

246

3 Structural Designs for Thermoplastics The coefﬁcient of linear thermal expansion for the transverse direction of melt ﬂow (perpendicular to the orientation of the ﬁber glass, also known as TD), has a much higher value than the coefﬁcient for the MD direction. It is very close to the coefﬁcient of linear thermal expansion value reported for the unreinforced polymers. The ﬁber glass oriented in the direction perpendicular to the melt ﬂow provides very little reinforcement for the polymer (a slight improvement provided by the ﬁber glass TD cross section areas). Plastics suppliers do not publish the coefﬁcient of linear thermal expansion for the transverse direction. These properties are available from the suppliers’ technical service departments. The coefﬁcient of linear thermal expansion for common thermoplastic resins and metals are shown in Table 3-3.

3.11.6

Metal Inserts Anchorage for Thermoplastic Encapsulation

Firm and permanent anchorage of metal screw inserts is essential and because there is no chemical or natural adherence between thermoplastics and metal inserts, anchorage must be obtained by mechanical means. The slight anchorage that is obtained by the mold shrinkage of thermoplastic around the metal screw insert is never sufﬁcient. Metal inserts must be anchored sufﬁciently to prevent turning when torque is applied and to prevent pulling out of the thermoplastic when subjected to tension. Internal stresses in the encapsulated thermoplastic must be kept to a minimum.

Flashing fill threads

Poor torque resistance

Poor design

Insert shoulder seals-off flashing

Coarse knurl anchorage

Good design Figure 3-82 Metal screw inserts anchorage for encapsulation

In the early days of plastics, it was customary to use hexagon metal stock for inserts. Other than in some special applications, this is mechanically incorrect. Hexagonal stock provides torsional anchorage only. Grooves must be machined to obtain sufﬁcient anchorage in tension. Combinations of sharp corners and grooves on hexagonal stock set up certain internal stresses in the thermoplastic, which often result in cracking. In almost all applications, round metal stock is recommended, so that a coarse diamond knurling can be obtained. Coarse diamond knurling provides the most satisfactory anchorage from the standpoint of torque and tension and minimizes possible cracking around the insert. Knurling of metal inserts is best accomplished in screw machines with end knurling tools. Grooves can be used in conjunction with a coarse diamond knurl. Sharp corners must be avoided when machining the grooves. When using grooves, one wide groove in the center of the insert is preferred over two grooves. The center groove allows the thermoplastic material to shrink or creep toward the center and minimizes the formation of strain in the encapsulated molded product. Figures 3-82, 3-83, 3-84, 3-85, 3-86, 3-87, and 3-88 show poor and good anchorage designs, selection of metal insert designs required for the thermoplastic encapsulation and types of anchorage for round bars and ﬂat sheet metal inserts.

247

3.11 Metal Inserts for Thermoplastic Encapsulation Horizontal seal only No seal or shoulder

Poor design

0.125 min.

Improved design

Limited vertical & horizontal seals 0.062 min.

Horizontal & vertical seals

Protruding metal screw inserts should have a shoulder to seal out any thermoplastic ﬂashing around the insert threads during molding.

Good design

Best design

Figure 3-83 Selection of metal screw inserts for encapsulation

Dual protruding screw inserts should have a shoulder extend above the top and below the bottom of the molded part to prevent ﬂashing on the threads. Mold misalignment may cause damage to the mold when closing. Figure 3-84 Dual protruding metal screw insert for encapsulation

Metal insert in bosses should extend to within one material thickness of the opposite wall. Ribs can be added for additional strength to the boss weld line. Poor desig n

Good design

Figure 3-85 Threaded metal insert boss design for encapsulation

248

3 Structural Designs for Thermoplastics No seal or shoulder

Horizontal seal only

Poor design

Improved design

Limited vertical & horizontal seals

Horizontal & vertical seals

Good design

Best design

Threaded metal inserts should have a shoulder to seal out any thermoplastic ﬂashing that might be forced inside the insert threads during molding encapsulation. Figure 3-86 Selection of threaded metal inserts for encapsulation

Typical anchoring designs used to encapsulate the sheet metal stamping inserts with thermoplastic. Figure 3-87 Types of sheet metal stamping inserts for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

Typical anchoring designs used to encapsulate the round cold metal forged insert bars with thermoplastic. Figure 3-88 Types of cold metal forged inserts for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

3.11 Metal Inserts for Thermoplastic Encapsulation

3.11.7

Metal Insert Encapsulating Process Problems

3.11.7.1

Metal Insert Floating (Movement)

The ﬂoating of metal inserts can be controlled or prevented by applying the following procedures: • The retaining pins that hold the insert in position may be tapered slightly, starting the taper at the ﬁllet and carrying it up to 1/3 of the length of the pin. If too much taper is allowed, making the insert too tight on the retaining pin, the insert may pull out of the encapsulated thermoplastic material. • The surface holding the retaining pin should be machined using a coarse straight knurl to improve the holding strength between the retaining pin and the insert holding area. • Split retaining pins are practical for holding blind hole inserts. • An extended circular shoulder should be provided on the insert to obtain a good sealed-off area; this shoulder bore in the ﬁxed half of the mold allows the proper positioning of the insert. This technique is ideal for preventing the insert from ﬂoating, although it is not permissible when inserts must be ﬂush with the mold parting line surface of the thermoplastic material. • For male screw metal inserts, a tapered hole can be provided for a drive ﬁt, if close accuracy of the inserts is maintained. In this case, 0.0005 to 0.001 in for the depth of the hole is sufﬁcient. • When precise location of the metal insert is essential, removable threaded pins are provided in the mold. Inserts are screwed into these pins. Subsequent removal of the ﬂash from the thread is in most cases avoided. However, this procedure increases the cost of the mold, causing a longer molding cycle and higher production costs. 3.11.7.2

Thermoplastic Melt Flashing Inside the Insert Blind Hole

Thermoplastic melt ﬂashing inside the insert blind hole is not as frequent as with the open hole type of inserts. Flashing is caused by loose retaining pins which allow the insert to move with the melt ﬂow, poor mold insert surface, or knurling on the entire outside diameter of the insert, leaving extended burrs on the face, which do not permit the insert to rest ﬂat on the surface of the mold or the surface of the retaining pin. It is good practice to provide a slight recess in the mold, accommodating the outside diameter of the insert. When the shoulder diameter of the retaining pin is the same as that of the insert and sharp corners can be retained in the hole, an 0.003/0.005 in increase in depth is sufﬁcient to prevent ﬂashing. It is desirable for the insert to protrude above the thermoplastic product surface, especially when electrical contacts are made. 3.11.7.3

Protruding Metal Insert Encapsulation Process

Protruding threaded metal inserts are frequently required and are encapsulated in thermoplastics for speciﬁc purposes. In most cases, the protruding section is used for assembly or for bearing points where a mechanical action is required. In special cases, especially with large metal inserts where the product encapsulated in thermoplastic is subjected to considerable torque in order to obtain a tight connection, the protruding section is necessary.

249

250

3 Structural Designs for Thermoplastics Hex metal insert

Smooth hex metal insert

Round metal insert

Radius all hex corners

Sharp corners

Circular shoulder Cracks

Bead Holes

Shrinkage stresses cause molding problems

Coating

Poor design

Holes

Good design

Figure 3-89 Protruding metal inserts encapsulation process

Figure 3-89 shows the process of encapsulating a protruding circular shoulderthreaded metal insert in thermoplastic part. Using a hexagonal metal insert can cause high stress concentration factors and product failure or cracks in the thermoplastic around the hex insert’s sharp corners.

Bead

Coating

Fair design

Coating

Figure 3-90 Thin tubular metal insert anchorage for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

3.11.8

Special Metal Inserts Anchorage for Encapsulation

For thermoplastic encapsulation applications where the product wall thickness is limited, the metal insert anchorage design is provided by a coarsely diamond knurled circular metal insert with a protruded shoulder without any sharp corners. A groove can be added to increase anchorage for tension. However, sharp corners must be avoided. When encapsulating a hexagonal metal insert with a protruded circular shoulder, all sharp corners must be eliminated and a circular groove be provided at the center of the external surface of the insert for tension anchorage. Thin Tubular Metal Insert Anchorage for Encapsulation

Countersunk holes

These inserts are extremely difﬁcult to anchor properly. When the tubular insert is located away from the molded part, it is possible to provide round holes or invert a bead from the tubular insert which will act as a satisfactory anchorage. The bead can be used outside or inside of the inserts as shown in Figure 3-90. When encapsulating the inside or outside of a tubular metal insert and a good joint is required, it is necessary to apply a thin coating to the metal insert surface (secondary operation) using an elastomeric polymer (liquid injection molding silicone, neoprene, vinyl, or a coupling agent) to improve the bonding strength between the thermoplastic and the thin tubular metal insert surface.

Angled lugs anchorage

Flat Plate Metal Insert Anchorage for Encapsulation

Figure 3-91 Flat plate metal insert anchorage for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

Flat plate metal inserts can be anchored by means of countersunk holes wherever it is permissible, or the section can be partially cut out and bent over to provide anchorage. All edges of the insert should be bevelled, except when certain sections of the insert are not required for the functioning of the product, as shown on Figure 3-91. When the ﬂat plate metal inserts must be thick, angled

251

3.11 Metal Inserts for Thermoplastic Encapsulation end lugs can be extruded and slightly ﬂared to provide satisfactory anchorage. Anchorage may also be obtained by spot welding lugs to the underside of the ﬂat plate metal insert. Drawn Shell Metal Insert Anchorage for Encapsulation When using a drawn shell metal insert slotted in the middle in a thermoplastic product with minimum wall thickness, extreme caution must be exercised to provide proper anchorage for encapsulation. Figure 3-92, top illustration, shows a poor anchorage design, because it allows insufﬁcient thermoplastic product wall thickness to avoid cracking. The middle illustration shows a fair anchorage design, because the encapsulated thermoplastic product has a chance to slide (ﬂashing) over the insert without providing good anchorage. The bottom illustration provides the best possible anchorage, the drawn shell metal insert slotted in the middle has an 80° shutoff parting line and the insert has been ﬂared slightly, providing good anchorage. With this design, the encapsulated thermoplastic product actually has a chance to anchor the insert and to creep while shrinking. Drawn Pin Metal Insert Anchorage for Encapsulation Figure 3-93, left illustration, shows a poor anchorage design; the drawn pin open end metal insert is encapsulated with thermoplastic during the molding operation and then the insert is countersunk after molding. A slight bead has been provided as an undercut for anchorage of the insert. This type of anchorage is basically useless for holding the insert properly.

Poor design

Fair design ˚

80

Good design Figure 3-92 Drawn shell metal insert anchorage for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

The middle illustration shows a good anchorage design for this type of insert; piercing pins should be provided in the mold, so that the insert can be pierced during the molding operation. The insert is ﬂared out to provide proper anchorage and the necessary thermoplastic countersink molded hole is formed. The right illustration shows the best anchorage design; the drawn pin open end metal insert is encapsulated with thermoplastic. Partial anchorage is obtained

Anchorage folding lugs

Poor design

Good design

Best design

Figure 3-93 Drawn pin metal insert anchorage for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

252

3 Structural Designs for Thermoplastics

Anchorage lugs

by shearing and folding two lug segments of the insert during the molding operation. To minimize thermoplastic melt ﬂashing into the insert, the ﬂoating pressure type of piercing pins are recommended as part of the mold. Large Drawn Shell Metal Insert Anchorage for Encapsulation

Tap threads after molding

Figure 3-94 shows a large drawn shell metal insert encapsulated with thermoplastic. In this application, the thread length or the insert space is limited. Because the drawn shell metal insert is thin, approx. 50% of the thread depth is used. The four ﬂared lugs of the insert provide satisfactory anchorage for encapsulation. However, it is impossible to provide sealing surfaces if a threaded insert is used, because the hot thermoplastic melt ﬂashes into the internal threads of the insert. Therefore, tapping threads inside the encapsulated insert after molding is recommended for satisfactory results. Large Surface Metal Insert Anchorage for Encapsulation

Anchorage lugs

Figure 3-94 Drawn shell metal insert anchorage for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

To encapsulate one or more large surface metal inserts on one side of a circular thermoplastic product, as shown in Figure 3-95, the circular inserts should have at least two grooves for anchorage. To retain the inserts in the mold, locating cavity grooves in one half of the mold are used; to avoid movement of the inserts during melt injection, ﬁxed retainer pins in the other half of the mold are used, which cause holes in the surface “B”. Special action retainer pins are used to plug the holes, the pins are moved forward to retain the inserts at mold closing, then the melt is injected in the cavity; just before the packing time starts, the retainer pins are retracted, allowing the melt to ﬁll the holes before mold ejection.

Metal inserts

Delay ejection retainer pins

Surface "A"

Surface "B"

Figure 3-95 Large surface metal inserts anchorage for encapsulation (Courtesy: The Society of the Plastics Industry, Inc.)

These inserts cause nonuniform mold shrinkage and warpage of the product. Surface “A” becomes convex while surface “B” becomes concave after the part is allowed to cool and age. If a ﬂat surface is required, the encapsulated product should be annealed after molding and, as the last step, the surface should be machined. Poor design

Delay ejection retainer pins

Good design Figure 3-96 Irregular shaped metal inserts anchorage to encapsulate

Irregular Shape Metal Insert Anchorage for Encapsulation These irregularly shaped metal inserts cause great difﬁculties in encapsulation. Figure 3-96, top, is a poor design, where a long “U” shaped insert is partially encapsulated, causing the one-piece insert to ﬂoat, molded-in stresses (differential shrinkage/thermal expansion), and cracking of the thermoplastic. The bottom illustration is a good design, where two separated inserts are supported by either ﬁxed or delayed ejection retainer pins. This design has several advantages, among them better encapsulated product dimensional control, product free of moldedin stresses, improved quality control, reduction of product defects, and better process efﬁciency in long molding production runs.

253

3.11 Metal Inserts for Thermoplastic Encapsulation

3.11.9

Electrical Lead Inserts for Encapsulation

If electrical contacts are required for the application, the insert can be made solid, ﬂat, thin, and narrow. The electrical lead insert should be curved (serpentine like, winding curves) to compensate (a slight give in the curved insert) for the thermoplastic melt shrinkage around the insert, as shown in Figure 3-97. Straight inserts do not provide a ﬁrm and permanent anchorage required for a good joint between the electrical lead and the encapsulated thermoplastic product.

Electrical leads with 90° directional bends reduce the encapsulation problems caused by the different coefficient of linear thermal expansions of metal and thermoplastic. Top view A

When a long electrical lead type insert is used, it is advisable to provide an anchorage in the middle of the insert by means of retainer pins with a coarse diamond knurl on the clamping surface. The center anchorage will allow the thermoplastic melt to creep along the surface of the insert while it is shrinking toward the center. Where dimensional accuracy is required, full allowance for the differential shrinkage of the materials should be made. Leakproof Encapsulation of Various Inserts Because of the difference in the behavior between thermoplastics and metals (e.g., difference in the coefﬁcient of linear thermal expansion) and the problems of providing proper insert anchorage, it is impossible for the insert to remain airtight within the thermoplastic, even under low injection pressures. When several inserts are used to encapsulate a leakproof product, special processing techniques are used to withstand high internal injection pressures, while retaining a leakproof joint. A leakproof encapsulated joint between two materials can be obtained by controlling the insert and the encapsulated product expansion and contraction. It is necessary that the inserts (before encapsulation) be precoated with a thin elastomeric adhesive and the walls around the insert be thin and uniform to compensate for the difference in the coefﬁcient of linear thermal expansions.

Electrical leads A Cross section view "A - A" Lead insertion mold guides

Electrical leads Scale = 2:1

Figure 3-97 Electrical leads multidirectional bends encapsulation (Courtesy: Delphi Automotive Systems)

A few successful techniques for leakproof joints of inserts encapsulated with thermoplastics are reviewed. The external surface of the insert should be modiﬁed for anchorage with two or more grooves, about 0.03 in wide and 0.02 in deep. The front or anchorage of the insert should be dipped in neoprene, polyvinyl chloride acetate, or other rubbery material, and then cured before using. This will supply sufﬁcient coating on the insert to give it the necessary cushioning action. It is also possible, especially on round inserts, to provide a large enough groove in the anchorage head of the insert so that a high temperature “O”-ring can be used. Under normal encapsulated molding conditions, the “O”-ring will produce satisfactory leakproof results. For some applications, such as the bobbin shown in Figure 3-98, a retaining groove for the “O”-ring is molded on the thermoplastic bobbin insert, sealing the metal can insert and providing a leakproof joint to stop the thermoplastic melt ﬂow inside the bobbin wound wires during the encapsulation. Encapsulation of Reinforced Metal Inserts It is necessary to encapsulate the metal inserts with thermoplastics to reinforce the product, providing greater strength, rigidity, safety and accuracy. Rather than molding a thermoplastic product with a thick wall, which can cause severe operational problems, using correctly designed inserts will not only produce greater rigidity with a minimum product wall thickness, but it will also assist in maintaining better dimensional accuracy.

"O" ring viton Thermoplastic encapsulation

Thermoplastic insert electrical bobbin Metal can Electrical insert lead

Figure 3-98 Leakproof thermoplastic bobbin insert encapsulation (Courtesy: Delphi Automotive Systems)

254

3 Structural Designs for Thermoplastics Figures 3-99 and 3-100 show encapsulated thermoplastic gears on hexagonal and square metal inserts. Figure 3-101 shows encapsulated thermoplastic pulley, using a ball bearing as an insert. These illustrations show typical poor product and insert designs, operational problems caused by insert sharp corners and thermoplastic mold shrinkage, and design recommendations to overcome the cracking problems. Hex metal insert

High stress concentration

Circular metal insert

Functional problems

Good design

Cracks

Poor design

Figure 3-99 Encapsulated thermoplastic gear on metal insert hub (Courtesy: Du Pont)

Square metal insert

High stress concentration

Reduced stress level

Cracks

Poor design

Functional problems

Good design

Figure 3-100 Encapsulated plastic gear on metal insert square hub (Courtesy: Du Pont)

Post molding shrinkage Double edge seal

Single edge seal

Cracks Encapsulated thermoplastic

Ball bearing

Open end for plastic expansion

Poor design

Functional problems

Figure 3-101 Encapsulated thermoplastic pulley on ball bearing (Courtesy: Du Pont)

Good design

3.11 Metal Inserts for Thermoplastic Encapsulation Nonmetallic Inserts for Encapsulation Inserts made of various materials are used successfully in encapsulation applications with thermoplastic polymers. Wooden or preformed inserts, for example, in bowling pins or golf balls use a core that is encapsulated with Surlyn resin to improve the surface ﬁnishing, hardness, scratch resistance, color, and toughness of the product, as shown in Ionomer Surlyn Applications, Chapter 1, page 17. The use of wooden inserts in applications, such as doorknobs or automobile gear shift knobs, saves considerable material and shortens the encapsulation molding cycle.

3.11.10 Inserts Preparation for Molding Encapsulation Cleaning the Inserts Proper cleaning or washing of inserts prior to molding encapsulation is essential, particularly when using metal screw machine inserts. If inserts are improperly washed, even though they appear clean, there may be loose metal chips hanging on to the threads or ﬁne metal dust in the knurls. The latter is often rolled onto the surface by the process of knurling and it is not easily washed off, but it will be loosened by the melt ﬂow of the thermoplastic. These metal chips may ﬂow up to the surface and impair the appearance of the molded product. The most serious difﬁculty, however, is in electrical applications, where a small particle or a slight amount of metal dust may cause a total electrical breakdown. Grease and oil also are detrimental to the appearance of molded products and should be thoroughly washed off. Cleaning processes are divided into three types: • Mechanical, including hand polishing, tumbling, sandblasting, or washing with solvent or alkali • Chemical, removal of iron rust and silver tarnish by an acid bath • Use of electrolytic cleaners

Oil and machining chips can best be removed by a well-stirred alkali bath followed by a rinse with hot water, except where the nature of the metal, such as aluminum, rules out the alkali in favor of degreasing with a solvent. In many cases, a reasonable amount of tarnish can do no harm, but where the function or the appearance of the piece demands chemically clean inserts, an acid dip is necessary. For brass and bronze, a mixture of nitric and sulfuric acids or nitric alone is commonly used. Silver tarnish can be removed with nitric acid or a diluted solution of one of the cyanides. Trisodium phosphate can efﬁciently remove iron rust. Preheating the Metal Inserts Large inserts should be preheated (above the mold cavity temperature if possible) prior to molding encapsulation. This will allow the maximum expansion and improve the melt ﬂow and delay the crystallization rate of the thermoplastic. With thermoplastic materials, preheating of inserts will reduce the possibility of weak weld lines, which often result in cracking of the thermoplastic after molding. Cleaning Flash from Metal Inserts Most of the difﬁculty with ﬂash can be avoided in the design of the thermoplastic product and the insert by providing good sealing areas so that the melt ﬂow of

255

256

3 Structural Designs for Thermoplastics thermoplastic is shut off. However, even with the best design there will be some material on the inserts, especially when the mold wears or close tolerance on inserts is not maintained. Several methods are recommended to minimize this ﬂashing problem, e.g., lubricating the insert prior to molding encapsulation with wax, soap, grease, or oil. Plating or hardening the inserts minimizes the tendency to produce ﬂash. To remove the ﬂash, cut the ﬂashing area, using special deﬂashing tools close to the molded product and peel it off. Burn the ﬂashing area using a small ﬂame from a low temperature portable torch, designed for trouble shooting the nozzle/melt freeze-off in an injection molding machine. For some polymers, a mild solution of caustic soda will loosen the ﬂashing area so that it can be easily removed. This method, however, requires extreme caution because extended contact or too strong a solution will harm the surface of the molded product and loosen the insert anchorage. Relieving Molded-In Stresses Considerable stresses are created in thermoplastic molded encapsulated products of irregular design, such as those having both thin and thick sections and especially those with metal inserts. The best method to relieve stresses is to allow the molded encapsulated product to cool slowly. The ideal situation would be to carry the molded encapsulated product on a conveyor through an oven that has various stages of temperatures, starting at 122 °F below the melting point of the polymer, then gradually decreasing until the molded encapsulated product is cooled to room temperature. The next best method requires two ovens, one at approximately 225 °F and the second at 150 °F. The molded encapsulated product remains in each oven successively until its temperature is reduced to oven temperature. The ﬁnal step is cooling to room temperature. In the case of thermoplastic materials, molding stresses are relieved by using an air oven, or submerging in a liquid solution at suitable recommended temperatures and procedures. Following are guidelines for the design of inserts for molding encapsulation: • Design the product without inserts if at all possible. • Do not encapsulate an insert unless the mold was designed for it. • Do not use an open hole or internal threaded insert if it can be avoided. • Do not leave sharp corners on inserts. Chamfer wherever possible. • Do not encapsulate inserts without proper anchorage. • Do not ﬂatten inserts if the insert is loose on the retaining pin. • Do not inject thermoplastic melt or ﬂashing into the insert to hold it on the retaining pin, this will damage the mold, necessitating repairs. • Do not encapsulate inserts unless they are clean. • Do not use the metal drawn type eyelet inserts unless necessary. • Do not design the product with a very thin wall for the base of the insert. This wall will corrugate and appear like a blister caused by a high heat spot. • Do not design the product with a very thick wall for the base of the insert; this can produce sink marks and internal voids. • Do not allow a very thin wall around the insert, the boss will crack.

257

4

Thermoplastic Gearing Design

Gears are used to transmit power and/or angular motion from one shaft to another. Examples of power gearing are the transmission and differential gears in automobiles and the gears in electric kitchen appliances and hand drills. Some gears, such as the gears between the hour and minute hands of a clock and the gears in some cameras used to focus the lenses from the range ﬁnder, are used primarily to transmit angular motion. Most gears used to transmit angular motion also transmit some power, such as in telescopes and radar drives. In this chapter, we will review the gear technology that is of importance when considering methods of gear manufacture. Gear manufacturing processes can be grouped into the following categories. • Molding Injection molding thermoplastic or powdered metal molding and casting: the completely formed gear is made in a mold having the shape of the ﬁnal gear, where the material is forced into the mold cavity in a melt or liquid form. • Metal Removal Such as milling, hobbing, and shaping, which cut the tooth spaces into prepared gear blanks leaving the desired teeth. • Finishing Processes Such as shaving, grinding, and lapping, which improve the accuracy and surface ﬁnish of previously prepared gear teeth. • Chipless Methods Cold rolling, used for small worms, some gears, and splines; hot rolling used for larger gears and other proprietary processes in which gears are formed by squeezing and/or rolling with forming gears; impact forging in which metal is forced into a gear shaped die by extreme impact. The selection of a speciﬁc manufacturing process to produce gears is inﬂuenced by the type of gear to be made. Because of the complex nature of gears, considering their accuracy requirements and the exacting demands in the shape of their teeth, the manufacturing processes are highly specialized. This limits the kinds and sizes of gears that can be produced on any one machine or with a single process. When evaluating processes that may be used to manufacture gears, it is desirable to understand the geometry of the gears to be made. Without exception, every manufacturing process imposes compromises on the production of perfect gears. The determination of the degree to which a gear can be compromised requires an understanding of its geometry so that the exact effects of the compromise can be evaluated. Many product designers and manufacturers feel that gears designed to very strict standards can be manufactured only with special gear molds. In fact, the involute gear tooth form is one of the most versatile geometries known to thermoplastic gearing engineering design. With a reasonable knowledge of involute geometry, the product designer can apply standard gear mold designs to an extremely wide variety of tooth forms, solving many special problems, none of which appear to be standard by ordinary engineering techniques.

258

4 Thermoplastic Gearing Design In this chapter, we will review the classiﬁcation and application of gears, the standard spur and helical gear technology developed for metal applications, the new gear tooth forms “Plastic Gear Technology” (PGT) developed for injection molded thermoplastic gears, molding and operational problems caused by gear design/gating, engineering analysis methods to demonstrate how to calculate these special tooth forms by working several examples to calculate the strength, size, and tolerances of gears and ﬁnally, several illustrations of gear design documentation required to design/construct the mold and for molding production quality control standards of the gear.

4.1

Classification of Gears

To understand gearing, it is desirable to classify the more important characteristics of the gear application, such as the relationship of the shaft axes on which the gears are mounted. Shafts may be parallel or nonparallel to the shaft axis. If non-parallel, they may be intersecting or nonintersecting.

4.1.1

Gears Parallel to the Shaft Axis

Spur Gears Injection molded thermoplastic spur gears are the most common type of gears; they have a cylindrical form and the involuted teeth are parallel to the axis.

Figure 4-1 Spur gears

The American National Standard (ANSI B6.1-1968) provides two involute spur gear tooth forms. These two forms are identical, except than one has a 20° and the other has a 25° pressure angle while both tooth forms have a minimum allowable tooth number of 18. A gear tooth standard is established by specifying the tooth proportions of the basic rack. In recent years, the established standard of 20° pressure angle has become the universal standard spur gear tooth form. Figure 4-1 shows a typical spur gear set. Helical Gears

Figure 4-2 Helical gears

Helical gears have a cylindrical shape with an envoluted tooth form. The gear teeth may lie in one or two rows around the structure. The helical gear tooth orientation in the pinion and gear must be of opposite hand. The tooth elements are helixes about the axis of the gear. Figure 4-2 shows a large helical gear set. The American National Standard (ANSI B6.7-1967) provides a 20° tooth form for helical gears of 20 diametrical pitch and ﬁner. The tooth proportions of ﬁne pitch helical gear teeth are based on the normal diametral pitch and are the same as for ﬁne pitch spur gears. Single Helical Gears

Figure 4-3 Single helical gears

If the teeth lie in a single row, the gear is called single helical. Injection molded thermoplastic single helical gear systems are used for low speed/torque applications, especially with the pinions encapsulated around metal shafts. Figure 4-3 shows a typical large single helical gear set. In recent years, this application has increased with the introduction of new thermoplastic materials developed for the gear market.

259

4.1 Classiﬁcation of Gears Double Helical Gears If the teeth lie in two rows, separated by a narrow toothless zone, the gears are called double helical gears. The helixes are of opposite hands. Because of the limited strength of the plastic materials, they are not recommended for these types of gears, because the double helical gears are used in large and high torque transmission applications as shown in Figure 4-4. Herringbone Gears Herringbone gears are cylinders with two rows of helical teeth that join in the middle, the gears have a herringbone appearance, thus their name.

Figure 4-4 Double helical gears

Because of the limited strength of the plastic materials, they are not recommended for these types of gears; most herringbone gears are used in large and high torque transmission applications. Herringbone gears are used when a smooth, continuous action is essential, as in high speed drives. These high speeds are encountered in marine transmission gears, particularly in connection with steam turbine and electric motor drives. Figure 4-5 shows a typical herringbone gear set. Internal Gears Internal gears have a cylindrical shape with envoluted teeth and have either spur or helical teeth. By deﬁnition, a left handed internal helical gear is one that meshes with a left hand mating pinion. An internal gear is like a standard spur gear turned “outside in”, as shown in Figure 4-6. To avoid interference in the tooth form, the gear’s internal diameter is increased and the mating pinion outside diameter is also made larger. Injection molded thermoplastic internal gears are commonly used in many applications, such as in the planetary gear system drives and others.

4.1.2

Figure 4-5 Herringbone gears

Bevel Gears, Nonparallel and Intersecting Shafts

Bevel Gears Conical in form, the proﬁles of the teeth are of specially generated forms, depending on the application, economics, and other factors. There is a variety of special tooth proﬁle forms identiﬁed as follows:

Figure 4-6 Internal gears

Straight Bevel Gears Straight bevel gears have straight tooth elements that, if extended, would pass through the point of intersection of their axes. Injection molded thermoplastic straight bevel pinion gears are commonly used in many low speed/torque transmission applications. These straight bevel teeth forms, as shown in Figure 4-7, are the most commonly used in gear reduction applications, but their tooth sides are tapered so that they would intersect the axis at a common point called the pitch cone apex if the bevel pinion is extended inward.

Figure 4-7 Straight bevel gears

260

4 Thermoplastic Gearing Design Spiral Bevel Gears Spiral bevel gears have curved oblique teeth on which contact begins gradually and continues smoothly from end to end. They mesh with a rolling contact similar to straight bevel gears. Spiral bevel gears run smoothly, quieter, and with reduced vibrations at high speeds. Injection molded thermoplastic spiral bevel pinion gears have limited use in low speed/torque transmission applications, because of the mold complexity. Figure 4-8 shows these types of gears. Figure 4-8 Spiral bevel gears

Zerol Bevel Gears The teeth of Zerol bevel gears are curved but lie in the same general direction as the teeth of straight bevel gears. The face cone elements of the Zerol bevel gears do not pass through the pitch cone apex but are approximately parallel to the root cone element of the mating gear to provide uniform tooth clearance. Injection molded thermoplastic Zerol bevel gears have a limited number of applications, because they are not commonly used. Figure 4-9 shows a typical Zerol bevel gear set. Crown Gears

Figure 4-9 Zerol bevel gears

Crown gears have teeth forms either straight or curved that lie in a plane pitch surface. Plastic materials are not used for this gear system. Skew Bevel Gears Skew bevel gears are similar to the crown gears, except that the tooth forms are straight and oblique, as shown in Figure 4-10. Thermoplastic materials are not used for this gear system. Miter Gears Miter gears are similar to bevel gears, having an equal number of teeth with axes at right angles. The specialized miter gear tooth forms are:

Figure 4-10 Skew bevel gears

• Coniﬂex, which has straight, crowned teeth. • Formate, in which the gear member of the pair has nongenerated teeth, usually with straight proﬁles and the pinion has generated teeth that are conjugate to the gear. • Revacycle, which has straight teeth generated by a special process with a special tooth form. Face Gears Face gears consist of a spur or helical pinion in combination with a conjugate gear of disk form; the axes are at right angles, either intersecting or nonintersecting. Injection molded thermoplastic face pinion gears have limited use in low speed/ torque transmission applications. Figure 4-11 shows a typical face gear set.

Figure 4-11 Face gears

261

4.1 Classiﬁcation of Gears

4.1.3

Hypoid Gears, Nonparallel and Nonintersecting Shafts

Hypoid Gears Hypoid gears are similar to bevel gears, but operate on nonintersecting axes. They have tooth forms that are curved and oblique. The tooth surfaces of both gear and pinion are cut or generated by the same or similar tools. Hypoid gears resemble spiral bevel gears, except that the axis of the pinion is offset relative to the gear axis. If there is sufﬁcient offset, the shafts may pass one another, permitting the use of a compact straddle mounting on the gear and pinion. The advantage of this design is that the pinion diameter is increased which makes it stronger than a spiral bevel pinion. Because of the limited strength of the plastic materials and complexity of the mold, they are not recommended for these types of gears. Hypoid gears, as shown in Figure 4-12, are used in high gear reduction and high torque transmission applications.

Figure 4-12 Hypoid gears

Worm Gears The worm gear system consists of a worm pinion and a worm gear with their axes at right angles to each other. Cylindrical Worm Gears

Figure 4-13 Cylindrical worm gear

A cylindrical worm gear has a tooth form like the threads of a screw. The tooth form commonly used is an involute helicoid. The cylindrical worm gears have one or more involuted helicoid threads like the screw threads on a cylinder, as shown in Figure 4-13. Hour Glass Worm Gears The hour glass worm gears, as shown in Figure 4-14, have one or more threads with their diameters increasing from the middle section towards both ends of the threads, conforming to the curvature of the mating worm pinion gear. The hour glass worm gears are also termed enveloping worm gears.

Figure 4-14 Hour glass worm gear

Single Enveloping Worm Gears A single enveloping worm gear is a mate to a worm. A worm gear that is completely conjugate to its worm has a single involuted helicoid contact surface; this type of gear interaction is termed as single enveloping. These types of worm gears are injection molded from thermoplastic polymers for many automotive and electrical gear reducer applications. The single enveloping worm gear reduction ratio between the speed of the worm gear and the speed of the worm wheel may range from 1.5 to 100. Worm gears having high ratios are not very efﬁcient as transmitters of power because of the effect of the low lead angle. The single enveloping worm gears are not used when the primary purpose is to transmit power efﬁciently. Figure 4-15 shows a typical single enveloping worm gear set. Double Enveloping Worm Gears A double enveloping worm gear system consists of an hour glass worm mated with a fully conjugated throated worm gear. The contact between the worm and the worm wheel is theoretically a line contact; however, due to the tooth

Figure 4-15 Single enveloping worm gears

262

4 Thermoplastic Gearing Design deﬂection under load, the contact line is increased to a narrow contact zone. The larger tooth bearing area and multiple tooth contact increases the load carrying capacity, using smaller sizes of worm gears. Because of the limited strength of the plastic materials and the complexity of the mold, they are not recommended for these types of gears. Double enveloping worm gears are used in high gear reduction ratios and high torque transmission applications. Figure 4-16 shows a typical double enveloping worm gear. Figure 4-16 Double enveloping worm gear

Crossed Axial Helical Gears The crossed axial helical gears operate on crossed axes and may have the teeth oriented in the same or opposite direction. Figure 4-17 shows a crossed axial helical gear set. Injection molded thermoplastic crossed axial helical gears have limited used in low speed/torque transmission applications only. Spiroid Gears Spiroid gears, as shown in Figure 4-18, have their axes at right angles to each other. The pinion member is conical in shape and the mating member is a facetype gear. They have tooth forms that are curved and oblique. Figure 4-17 Crossed axial helical gears

The Spiroid gears resemble the spiral bevel gears, except that the axis of the pinion is offset relative to the face-type of gear axis. The advantages of the Spiroid gears are the high gear reduction ratios, fast operating speeds, and high torque transmission performance. Because of the limited strength of plastics and the complexity of the mold, they are not recommended for use in Spiroid gear applications. Helicon and Planoid Gears These types of gears are members of the Spiroid family, a Helicon pinion gear is a Spiroid pinion gear without a taper tooth form.

Figure 4-18 Spiroid gears

4.1.4

Gears for Straight Linear Motion

Spur Rack Gear A spur rack gear is a straight bar having a rectangular cross section area with a spur tooth form, spaced along a straight line and perpendicular to the top surface of the bar. Spur Pinion Gear

Figure 4-19 Spur rack and pinion gears

The spur pinion gear has a cylindrical form with involuted straight teeth at right angles to the direction of motion. These types of gears, as shown in Figure 4-19, have a metal insert centrally located for mounting a shaft. The insert has been encapsulated with a thermoplastic material using an injection molding process. Helical Rack Gear A helical rack gear is a straight bar having a rectangular cross section area with a helical tooth form that is oblique to the direction of motion, spaced along a straight line, and perpendicular to the top surface of the bar.

4.2 Standard Injection Molded Thermoplastic Gears

4.2

Standard Injection Molded Thermoplastic Gears

Thermoplastics are very good materials for low speed, light load gears for reasons of economy, performance, production efﬁciency, and lower weight. Engineering thermoplastic polymers are well suited for gears because of their combination of thermal, mechanical, electrical, environmental resistance, and ﬂame retardant properties. Gear loads are transmitted during the gear motion by mating teeth which mesh with each other. This action tends to bend the teeth and subject them to a certain sliding motion. The wear characteristics and lubricity of the gear material are very important. All gear types (spur, worm, bevel, helical, annular internal, and external) experience some combination of these rolling and sliding forces. In general, wear resistance is critical for the tooth form and the dimensional stability is critical for overall gear performance. The limits of performance for the injection molded thermoplastic gears are determined by the following parameters: • Environmental temperature generated on the teeth by friction • Bending stresses on the tooth root • Fatigue or impact resistance • Wear resistance and coefﬁcient of friction • Chemical resistance It is only a question of operating conditions, whether an engineering thermoplastic material can be considered a good candidate for a gear. If it appears that an engineering thermoplastic material will meet the speciﬁed requirements, a product designer must realize that the use of an engineering thermoplastic material also provides a number of advantages and design ﬂexibilities that cannot be achieved when using metal materials, such as: • Post machining operations or removing of burrs are not required • Possible combination of gears with other elements, such as springs, bearings, ratchets, cams, and other gears • Corrosion resistance • Electrical insulation • Better dry running behavior than metal materials • Lower noise levels Thermoplastic gears can be molded by using special thermoplastic compounds developed for gear applications. In spite of the introduction of many new materials, the majority of applications still require gears made from acetal homopolymer and nylon 6/6. With the recent introduction of these new materials, such as Kevlar® ﬁber reinforced acetal homopolymer, nylon 6/6, PET and co-polyester elastomers, high performance gears are produced with these resins. In special circumstances, polycarbonate, polysulfone, polyurethane, and thermoplastic elastomer resins have been used for injection molded gears. Compounded thermoplastic materials have been modiﬁed to improve their mechanical properties. The reinforcements used in thermoplastic materials

263

264

4 Thermoplastic Gearing Design for gears are ﬁber glass and Kevlar® ﬁbers. The additives compounded with the plastic matrix to reduce the coefﬁcient of friction and to improve the wear resistance properties are zinc or aluminum stearate, polytetraﬂuoroethylene (TFE) Teﬂon® powders and ﬁbers, silicones, graphite, and molybdenum disulﬁde. Fiber reinforcements increase the tensile strength, the modulus of elasticity; they also improve the dimensional stability of the gears and the end use temperature of the molded gears. Powders do not improve the mechanical strength properties; they reduce the mold shrinkage properties improving the dimensional control of the molded gears. Fiber glass and Kevlar® reinforcement with coupling agents increases the level of mechanical properties, reduces the moisture absorption characteristics, and also reduces the coefﬁcient of linear thermal expansion to one third of the original matrix polymer value. Engineering thermoplastics containing both ﬁber reinforcement and internal low coefﬁcient of friction additives are very important for gear applications. The following specialty compounded resins were developed for gear applications: High viscosity acetal homopolymer, Kevlar® reinforced acetal, chemically lubricated acetal, acetal with TFE Teﬂon® ﬁbers, unﬁlled nylon 6/6 internally lubricated, unreinforced nylon 6/6 with a molybdenum disulﬁde, Kevlar® reinforced nylon 6/6, Kevlar® reinforced PET, unreinforced copolyester elastomer, Kevlar® reinforced copolyester elastomer, unreinforced lubricated polycarbonate, glass reinforced polycarbonate, lubricated polysulfone and polyurethane.

4.2.1

Selection of Thermoplastic Resins for Gears

As with any other material, the selection of injection molding thermoplastic resins is governed by the size and nature of the load to be transmitted, the speed, the life required, the environment in which the gear will operate, the type of lubrication, and the degree of precision necessary. The horsepower Eq. 4-1, 4-2 and 4-3 used for gear analysis in conjunction with Tables 4-1 to 4-3 will help to determine, which of the thermoplastic resins are good gear materials in terms of the mechanical strength required. These equations do not consider the end use temperature, the impact strength, coefﬁcient of friction to reduce heat dissipation, the coefﬁcient of linear thermal expansion, wear resistance, dimensional stability, or injection molding conditions. The equations are simple variations of the Lewis equations and assume the use of standard tooth forms used in metal gears. The results are conservative, but are of sufﬁcient accuracy to help in making a decision of whether a further in-depth study is warranted. The design analysis will require machining prototype gears, testing the gears using the same speciﬁcations and requirements (load, velocity, endurance, temperature, environmental conditions) of the gears designed for the speciﬁc application under review. The safety stress is calculated by dividing the tensile stress (Table 4-2) by a safety factor (e.g., 1.50), assuming initial lubrication. The values given for tensile strength of ﬁber glass reinforced plastics should be used with discretion. The values for ﬁber glass reinforced materials are superior to those for unreinforced thermoplastics, which are excellent for certain gear applications. For other applications, requiring greater strength, the ﬁber glass decreases the wear resistance, dimensional control (differential mold shrinkage); the teeth are rigid, they do not ﬂex the way we expect from a thermoplastic gear.

265

4.2 Standard Injection Molded Thermoplastic Gears Table 4-1 Gear Service Factors (SF)

Type of load

Occasional 1/2 hour/day

Intermittent 3 hours/day

8 to 10 hours/day

24 hours/day

Heavy shock Medium shock Light shock Steady

1.25 1.00 0.80 0.50

1.50 1.25 1.00 0.080

1.75 1.50 1.25 1.00

2.00 1.75 1.50 1.25

Table 4-2 Physical Properties of Thermoplastics

UL max. temperature °F

Coef. of linear thermal expansion, 10–5 in./lin. °F

2.3

200

6.8

M94 0.20 R120

–

Kevlar® reinforced acetal

14.3

6.5

5.0

200

6.8

M94 0.62 R120

3.9

Chem. lubricated acetal

9.5

4.1

1.4

200

6.8

M94 0.10 R120

–

Acetal with TFE ﬁber

6.9

4.2

0.7

200

6.8

M94 0.08 R120

–

Nylon 6/6 lubricated

10.2

1.8

1.2

250

5.0

M59 0.44 R108

22.9

Nylon 6/6 with MOS2

10.5

4.08

0.9

250

5.0

M57 0.18 R108

–

Kevlar® reinforced nylon

17.3

8.02

1.2

–

–

–

0.39

6.00

Glass reinforced nylon 6/6

18

9.0

2.5

310

1.3

M101

–

–

Kevlar® reinforced PET

14.8

8.33

1.6

356

–

–

0.26

0.50

Copolyester elastomer

4.7

0.01

–

275

–

–

0.77

420

Kevlar® copolyester TPE

4.7

0.05

–

275

–

–

0.68

1.1

Polycarbonate lubricated

8.7

3.15

15.0

250

3.7

M74 R120

–

–

Glass reinforced PC

19.0

11.0

2.0

–

1.2

M92 R120

–

–

Polysulfone

10.2

3.9

1.3

298

3.0

M69 R120

–

–

Polyurethane

3.9

0.01

–

240

–

–

1.05

380

Wear rate, 10–5 in./hr.

Notched Izod impact (ft-lb/in)

3.8

Static coef. of friction

Flexural modulus (105 psi)

10

High viscosity acetal

Rockwell hardness

Tensile strength (103 psi)

Thermoplastic resins

266

4 Thermoplastic Gearing Design Table 4-3 Lewis Tooth Form Factor (Y)

Number of teeth

14.5° involute or cycloidal

20° full depth involute

20° stub tooth involute

20° internal full depth Pinion

Gear

12 13 14 15 16 17 18 19 20 21 22 24 26 28 30 34 38 43 50 60 75 100 150 300 Rack

0.210 0.220 0.226 0.236 0.242 0.251 0.261 0.273 0.283 0.289 0.292 0.298 0.307 0.314 0.320 0.327 0.336 0.346 0.352 0.358 0.364 0.371 0.377 0.383 0.390

0.245 0.261 0.276 0.289 0.295 0.302 0.308 0.314 0.320 0.327 0.330 0.336 0.346 0.352 0.358 0.371 0.383 0.396 0.408 0.421 0.434 0.446 0.459 0.471 0.484

0.311 0.324 0.339 0.348 0.361 0.367 0.377 0.386 0.393 0.399 0.405 0.415 0.424 0.430 0.437 0.446 0.456 0.462 0.474 0.484 0.496 0.506 0.518 0.534 0.550

0.327 0.327 0.330 0.330 0.333 0.342 0.349 0.358 0.364 0.371 0.374 0.383 0.393 0.399 0.405 0.415 0.424 0.430 0.437 0.446 0.452 0.462 0.468 0.478 –

– – – – – – – – – – – – – 0.691 0.679 0.660 0.644 0.628 0.613 0.597 0.581 0.565 0.550 0.534 –

4.2.2

Horsepower Equations for Gears

Spur gear HP (external and internal) HP =

σS × F × Y × V 55 (600 + V ) P × SF

(4-1)

Helical gear HP (external and internal) HP =

σS × F × Y × V 423 (78 + V 0.5 ) PN × SF

(4-2)

Straight bevel gear HP HP =

(σ S × F × Y × V ) (C − F ) 55 (600 + V ) P × C × SF

Where: σS = Gear safety stress. σS = σ / Safety factor (psi) σ = Material tensile stress (psi) F = Tooth face width (in) Y = Lewis tooth form factor (Table 4-3) V = Velocity at pitch circle diameter (fpm) P = Diametral pitch PN = Normal diametral pitch SF = Service factor C = Close mesh center distance between gears (in)

(4-3)

4.2 Standard Injection Molded Thermoplastic Gears

Example 4-1 Determine the thermoplastic material for a spur gear that transmits 0.125 horse power (HP) at 320 revolutions per minute (RPM). The gear will run under a steady load for eight hours per day. Basic information for a standard spur gear: Number of teeth (N) Diametral pitch (P) Pressure angle (φ) Pitch diameter (DP) Face width of teeth (F) Transmission load (HP) Gear rotation (RPM) Tooth velocity (V) Service factor (SF) Lewis tooth form factor (Y) Tensile stress (σ) Safety working stress (σS)

70 32 20° 2.1875 0.375 in 0.125 HP 320 RPM fpm 1.0 (Table 4-1) 0.430 (Table 4-3) psi psi

Solution Calculate the spur gear tooth velocity (V): V =

(RPM) × π × DP 320 × 3.1416 × 2.1875 = = 183.26 fpm 12 12

Spur gear horsepower (Equation 4-1): HP =

σ × F ×Y ×V 55 (600 + V ) P × SF

Spur gear tooth tensile stress: 55 (600 + V ) P × SF × (HP) 55 (600 + 183.26) 32 × 1.0 × 0.125 = F ×Y ×V 0.375 × 0.430 × 183.26 = 5,831.37 psi

σ =

The safety working stress: σ S = 5,831.37 × 2 = 11,662.75 psi There are several ﬁber reinforced thermoplastic materials that could be used for molding the gear based only on Table 4-2. However, other physical and chemical properties of these thermoplastics must be studied depending to the environment in which the spur gear is to operate: e.g., whether these thermoplastic materials need to resist lubricants (oil, grease, etc.), UV and pollution exposure from the weather, they need to be dimensionally stable at high velocities, have excellently wear resistance characteristics and a low coefﬁcient of friction in relation to the material of construction used for the mating gear.

267

268

4 Thermoplastic Gearing Design

4.2.3

Spur Gear Terminology and Definitions Outside dia. (DO)

Pitch dia. (Dp)

Working depth (hK)

Base dia. (Db)

Clearance (c)

Driv

e r ge ar

Pitch point

Pressure angle ( )

Line of action

p To

lan

d

e Fa c nk Fl a

Addendum (a) Dedendum (b)

h idt ew Fa c (F)

Whole depth (ht) Circular pitch (p)

Ro

Circular tooth thickness (t)

ot dia . (D

Fillet radius (rf)

)

R

Dri

ve n gear

Figure 4-20 Standard spur gear terminology

Addendum (a): The height of a gear tooth outside the pitch circle; the radial distance from the pitch circle to the outside diameter of the addendum circle, numerically equal to the reciprocal of the diametral pitch. Addendum circle: The circle at the top of the gear tooth. AGMA number: A number indicating the relative quality of a gear as speciﬁed by the American Gear Manufacturers Association, a higher number indicating higher quality. Angle of action: The angle through which one tooth travels from the time it ﬁrst reaches its mating tooth on the line of action until contact ceases. This is divided into the angle of approach and the angle of recess. Backlash: The amount by which a tooth space exceeds the thickness of the meshing tooth, provided to compensate for thermal expansion; the difference between tooth thickness and tooth space as measured on the pitch circle. Base circle diameter (Db): The diameter of the base circle. Base pitch: The normal pitch of an involute gear; the distance between two successive parallel involutes which form the proﬁles of two adjacent teeth, equal to the circumference of the base circle divided by the number of teeth on the gear. Center distance (C): The distance between the centers of the mating gears. Circular pitch (p): The length of an arc of the pitch circle that corresponds to one tooth interval, equal to the circumference of the pitch circle divided by the number of teeth on the gear. Circular tooth thickness (t): The thickness of a single tooth measured along the pitch circle; for an unmodiﬁed tooth, equal to one-half the circular pitch.

269

4.2 Standard Injection Molded Thermoplastic Gears Clearance (c): A small space provided so that the top of a meshing tooth will not touch the bottom land of the other gear as it passes the line of centers. Dedendum (b): The depth of a tooth space below the pitch circle; the radial distance from the pitch circle to the root circle, equal to the addendum plus the tooth clearance. Diametral pitch (P): The ratio of the number of teeth to the pitch diameter of a gear, representative of the number of teeth per inch of pitch diameter (DP). P=

NG + NP π×C

(4-4)

Face width (F): The surface of a gear tooth lying between the pitch circle and the addendum circle. Fillet radius (rf ): The radius of curvature of the corner where a tooth joins the root circle. This corner is radiused with extra material to reduce the build-up of stress concentrations. Flank: The surface of a gear tooth between the root diameter (DR) and the pitch diameter (DP). Gear ratio: The ratio of the number of teeth in the gear to the number of teeth in the pinion. Interference: A term relating to conditions that permit contact between mating teeth away from the line of action to interfere with the transmission of uniform motion. Involute: A system of gearing; the principal proﬁle of a gear tooth; a curve generated on a circle, the normal of which are all tangent to that circle. Line of action: The line along which correct contact between mating teeth is made, resulting in transmission of uniform motion from one gear to another. Number of teeth in gear (NG): The total number of teeth of the gear. Number of teeth in pinion (NP): The total number of teeth of the pinion. Outside diameter (DO): The external diameter of the gear. Pitch circle: A circle that represents a smooth disc that would transmit the desired relative motion by friction. Pitch diameter (DP): The diameter of the pitch circles of mating gears. Pressure angle (φ): The angle between the line of action and a line perpendicular to the common center line of two mating gears or the angle cutting the tooth face at the pitch point and the tooth face itself. The most common pressure angles are 14.5°, 20°, and 25°, of which 20° is the most commonly used. Root circle: The circle at the bottom of the tooth spaces in a gear. Root diameter (DR): The diameter of the root circle. Whole depth of tooth (ht): The total depth of the tooth space in a gear measured radially between the addendum circle and the root circle. Working depth of tooth (hK): The depth that the teeth of one gear extend into the spaces of its mating gear, equal to the sum of the addenda of mating gears; also equal to the whole depth minus the clearance.

270

4 Thermoplastic Gearing Design Spur Gear Tooth Stress For standard injection molded thermoplastic spur gear designs, as with any other material, it is advisable to calculate the increase in gear size at the highest end use temperature and provide sufﬁcient backlash to prevent binding. Since the mechanical properties of the engineering thermoplastic resins are temperaturedependent, the load-bearing capacity of the injection molded thermoplastic spur gear decreases with increasing temperature, depending on the thermal characteristics of the material selected. Spur gears are also very susceptible to stress build-ups at the roots of the teeth caused by shock loading. Careful consideration has to be given to the notch impact sensitivity of the engineering thermoplastic material selected for the application. It is best to extend a full radius at the tooth root of the gear. Backlash can effectively allow for dimensional changes in the gears. Radii at the tooth root can improve the shock load compatibility. One of the advantages with the teeth of an injection molded thermoplastics spur gear is that the teeth bend without causing ﬂexural stress or fracture; the endurance limit of the material should be used when designing to allow for cyclic reduction of mechanical stress. The spur gear tooth form, however, does not conform to the idealized conﬁguration of the three point bending tests used to determine the ﬂexural strength by ASTM D790. Therefore, it is often preferable to use the tensile strength of the thermoplastic material and to provide a greater margin of safety. Gear Safety Stress (σ S ) =

Maximum Tensile Strength of the Plastic Safety Factor

Spur Gear Tooth Tangential Force The spur gear tooth load limits or the design requirements can be calculated using the Lewis equation: W =

W

σS × F × t 2 6 × ht

σS =

6 × W × ht F × t2

where W is the tangential force at the end of tooth, F is the uniform face width of tooth, ht is the whole depth of the tooth decreasing towards the loaded end. The outline of the tooth form is equivalent to a parabola with vertex at the loaded end.

F

Spur Gear Tooth Transferred Torque Torque is transferred by a gear through the tangential forces acting on the gear teeth. If only one tooth of each gear in contact at or near the pitch point is carrying the entire load, a simple method of calculating the tangential force on a gear tooth is given by:

t

W =

σS × F × Y P

(4-5)

W =

2×T DP

(4-6)

ht

Figure 4-21 Tooth tangential force

271

4.2 Standard Injection Molded Thermoplastic Gears Rearranging: T =

σ S × DP × F × Y 2×P

(4-7)

Where: T σS t ht F W Dp F Y P

= Torque of gears (in-lb) = Safety working stress (psi) = Circular tooth thickness = Tooth whole depth = Tooth face width = Tangential force on tooth (lb) = Pitch diameter (in) = Face width of tooth (in) = Lewis tooth form factor = Diametral pitch

Many gears are subjected to a “stall torque” that is signiﬁcantly higher than the actual operating torque. It is advisable to use the yield stress of the material at the operating temperature, calculating the bending stress of a cantilever tooth. Factors Affecting Spur Gear Tooth Loading To calculate the gear tooth load accurately, several factors must be known: • The spur gear reduction ratio • The center distance limits • The spur gear or pinion operating speed • The spur gear or pinion horsepower or torque • The type and duration (time) of loading • The Maximum Operating Temperature • The type of operating service requirement • Type of lubrication The spur gear reduction ratio is determined by the desired rate of motion between the driving (or input) spur gear and the driven (or output) spur gear. The center distance is set by the space allocated between the spur gear and the pinion. The speed of either the driving spur gear or the driven spur gear must be known, as well as either the horsepower or the torque to be transmitted. The expected spur gear service life should be stated along with the operating conditions (type of loading, duration, temperature, environment, and lubrication). In addition, the spur gear pressure angle and the diametral pitch are critical factors, because they determine the size and form of the spur gear teeth and, consequently, the load sharing ability and the strength of the teeth. The most common spur gear pressure angles are 14.5°, 20°, and 25°, with the 20° pressure angle being used most often. The 20° pressure angle has a higher load bearing ability than the 14.5° angle and it allows fewer pinion teeth to be used, avoiding the undercutting of the gear teeth before it becomes a manufacturing concern.

272

4 Thermoplastic Gearing Design

4.3

Properties Required for Injection Molded Thermoplastic Gears

Injection molded thermoplastic spur gears depend on similar properties as those used in the principles of design established for standard metal gears. However, the mechanical, thermal, and creep characteristics of the thermoplastic polymers make it essential to adhere to these properties more rigidly than would necessarily be the case in designing metal spur gears that will be machined. The coefﬁcients of linear thermal expansion of reinforced thermoplastics are approximated in 1/3 of those of unreinforced resins. The expansion of metal gears with an increase in temperature is not a signiﬁcant design concern to take into consideration. However, for injection molded thermoplastic spur gears, it is necessary to calculate the amount by which thermoplastic gears will expand at the highest temperature to which they will be subjected and to provide sufﬁcient backlash between the teeth to prevent binding. For some hygroscopic thermoplastic polymers, such as nylons, the size of the gear increases slightly when absorbing moisture. This dimensional change in nylon gears is rarely a problem, but additional backlash should be allowed between the teeth if the gear mechanism may be unused for long periods in damp ambient operating conditions. The teeth of heavily loaded metal spur gears in critical drives are given a degree of tip relief and have full ﬁllet root radii to reduce fatigue stresses. These modiﬁcations should also be speciﬁed for the teeth of all thermoplastic spur gears. In designing a pair of metal spur gears, variations in both gear teeth addendum lengths are frequently used. If the pinion has a small number of teeth, these teeth may be undercut. Undercutting weakens the strength of the gear teeth, causes wear, and reduces the life expectancy of the gear. The undercutting process can be eliminated by increasing the addendum of the pinion teeth and decreasing addendum of the gear teeth. The elimination of undercutting is beneﬁcial and it is recommended in the design of injection molded thermoplastic spur gears, because it provides better tooth strength and gear design ﬂexibility. Injection molded thermoplastic spur gears offer several beneﬁts. The designer is freed from many of the limitations imposed by using machine tools to make the blanks, standard lobs and cutters, in shapers to form the teeth of metal gears. The expensive machining and assembly operations are eliminated by designing the gear as an integral part of the mechanism, allowing desirable modiﬁcations to the standard tooth to be speciﬁed without increasing the price of the gear with little or no additional cost for the molding tool. Today, there are instrument mechanism gears of such complexity being molded that their fabrication in metal would not be economically feasible. Consulting with the tool, process, and manufacturing engineers, the resin supplier’s technical assistants, and the mold maker before the ﬁnal gear designs are approved may well result in considerable cost savings in the overall project by combining several features in a single multifunction injection molded gear design.

273

4.4 Thermoplastic Spur Gear Design Requirements

4.4

Thermoplastic Spur Gear Design Requirements

Single cavity mold Web = H x 1.1

When designing thermoplastic gears, it must be remembered that the gears are not only supposed to fulﬁl the expected mechanical function, but that they should also be dimensioned in a way to facilitate correct and efﬁcient injection molding operations. The simpler the geometric shape, the easier it will be to ﬁll the cavity properly and, if required, to achieve tight tolerances.

Whole depth (H)

Hub = H x 1.2

Two ideal spur gears and gate designs are shown in Figures 4-22 and 4-23. To provide mechanical strength, the rim wall section supporting the spur gear teeth should be at least the same tooth whole depth dimension (H). The vertical cross sections below the rim and hub wall thickness depend on the functional requirements of the gear, the type of gate, and the location of the gate. Figure 4-22 shows a diaphragm gate, centrally located at the spur gear hub; a wall thickness 20% larger than the rim and 10% larger for the vertical web section is preferred. Figure 4-23 shows a three or four pin point gate (requires a threeplate mold) located in the spur gear web; the web wall thickness should be 20% larger than the rim and the hub wall thickness should be 10% larger than the rim. Proportionally distributed wall thicknesses of the gear and the correct type of gate are recommended to obtain excellent dimensional control, including close TIR tolerances, without warpage.

Sprue Sprue puller

Rim (H)

Figure 4-22 Spur gear with sprue diaphragm gate

Multi-cavities, 3 plate mold Web = H x 1.2

Thermoplastic Spur Gear Classic Designs Whole depth (H)

Figures 4-24, 4–25, and 4–26 show three spur gear case designs illustrating a poor spur gear design, molding problems caused by an inadequate gear design, and design recommendations to overcome these types of molding problems.

Hub = H x 1.1

Spur Gear Case Design “A” Figure 4-24, left illustration, shows a spur gear with a thin wall thickness web that does not allow the rim to pack. The middle illustration shows internal voids around the rim, caused when the thin web froze-off and the rim could not be ﬁlled with more thermoplastic melt. The internal voids cause higher shrinkage than in the rest of the gear, producing poor dimensional control (roundness) and the middle of the tooth face width becomes concave. The right illustration shows proportional wall thickness as recommended for the gate type in Figures 4-22 and 4-23.

Thin web

Sucker pin

Pin point gate Rim (H)

Figure 4-23 Spur gear with 3 or 4 pin point gates

Sink teeth (concaved)

Gate

Voids

Poor design

Operational problems

Figure 4-24 Spur gear case design “A”

Recommended design

274

4 Thermoplastic Gearing Design

Web warpage

Web offset Voids Hub, heavy wall section Hub sink surface

Poor design

Operational problems

Recommended design

Figure 4-25 Spur gear case design “B”

Spur Gear Case Design “B” Figure 4-25, left illustration, shows a gear with a heavy hub wall thickness, with the web located at the rim’s edge. The middle illustration shows operational problems, with internal voids around the hub, because the heavy hub wall cannot be packed. The voids high shrinkage causes the hub’s inside diameters to sink. The offset web causes warping of the gear (uneven mold heat distribution), poor dimensional control (roundness), and bending of the tooth’s face width. The right illustration shows the recommended gear design, a uniform and balanced wall thickness, a distributed web that controls both mold halves’ cooling temperature in the cavities, keeping mold shrinkage well under control. Spur Gear Case Design “C” Figure 4-26 left illustration, shows a heavy web wall with some material removed by the four holes. The middle illustration shows operational problems, voids inside the web, because the heavy wall of the web cannot be packed. The web and hub voids cause high mold shrinkage, warping both sides of the teeth, poor dimensional control (roundness), sink marks on the hub’s outside and inside diameter. The right illustration shows the recommended design, with the web wall and cross sections uniformly/equally spaced, so that they function like four proportional webs/ribs to support the rim and the spur gear teeth.

Voids

Voids

Hub sink Web sink Teeth sink

Poor design Figure 4-26 Spur gear case design “C”

Operational problems

Recommended design

275

4.4 Thermoplastic Spur Gear Design Requirements

4.4.1

Gating Effects on Thermoplastic Gear Roundness Dimensions

High shear rates produce two effects that signiﬁcantly affect the spur gear performance. The plastic molecules become aligned as a result of the high shear rates so that the melt in the walls is highly oriented in the melt ﬂow direction. This effect is undesirable, because the strength in the direction perpendicular to the ﬂow direction is reduced and the gear has a tendency to crack along the weld line. The other effect is the dimensional change caused by the orientation of the melt ﬂow. Figure 4-27 shows the gear elongation and compression problems caused by an edge gate located on the spur gear tooth. When the melt ﬂow is split by the core of the gear hub, both melt fronts, which are the coldest part of the melt, are reunited at the back side of the core causing a weld line area that is the weakest section of the gear. The ﬂow direction of the melt has the higher injection pressure, causing less mold shrinkage (larger dimensions). The pressure perpendicular to the ﬂow direction is lower, causing higher mold shrinkage (smaller dimensions). The mold shrinkage becomes more pronounced with ﬁber reinforced thermoplastics; the ﬁber oriented in the ﬂow direction does not compress and has lower shrinkage, while the ﬂow in the perpendicular direction has higher mold shrinkage. Figure 4-28 shows a single pin point gate located in the web wall of the spur gear. This type of gate requires a three-plate mold to break the gate and to mold automatically. A hot drop gate from hot runnerless molds can also be used for molding gears. This type of gating causes the same molding problems as the edge gate, but provides better results, because the gate is more centrally located. The weld line is improved and the differential mold shrinkage is reduced. The molded gears have better TIR (roundness) than the edge gate shown in Figure 4-27.

Weld line Transverse flow compression

Melt flow behavior

Edge gate

Flow direction elongation

Figure 4-27 Single edge gate on the spur gear tooth width

Weld line Web Reduced compression

Gate

Melt flow behavior

Figure 4-28 Single pin point gate on the spur gear web wall

Reduce elongation

276

4 Thermoplastic Gearing Design Hub

Gate

Weld line

Figure 4-29 shows a pin point gate located in the hub wall of the spur gear. This type of gate requires a three-plate mold to break the gate and to mold automatically. Some hot drop gates from hot runnerless mold systems can also be used for molding gears. This type of gating causes the same molding problems as the single pin point gate in the web, but provides better results, because the gate is more centrally located. It has a stronger weld line, better warpage control, and reduces differential shrinkages caused by the melt ﬂow or ﬁber orientation. Molded gears have better TIR (roundness) than the two previous cases (Figures 4-27 and 4-28). Figure 4-30 shows a spur gear with three pin point gates equally spaced around a circle in the web wall of the gear. To avoid molded-in stresses and premature freeze-off of the gates, a diffuser (circular indentation) is provided behind each gate on the web’s back side. This gating requires a three-plate mold to run a multi-cavity tool automatically. Hot drop gates from some hot runnerless molds are also used for injection molded thermoplastic spur gears. This type of gating is very common in molding close precision tolerances for high velocity gear applications. This system provides strong weld lines, better warpage control, and minimum differential shrinkages caused by the melt ﬂow or ﬁber orientation.

Melt flow behavior

Less compression

Figure 4-31 shows a sprue diaphragm gate being used for molding a single cavity spur gear on a two-plate mold. It may also be used for molding multi-cavity gears using some hot runnerless molds. This type of gate allows uniform melt ﬁlling, equal pressure, shrinkage, and melt ﬂow distribution inside the cavity. The sprue diaphragm gate allows the best precision dimensional control (roundness) without weld lines. The diaphragm gate disk connected to the sprue requires a post molding operation to remove the diaphragm gate disk and sprue from the inside diameter of the hub. For the multi-cavity hot runnerless molds, it requires a post molding operation to remove only the diaphragm gate disk.

Less elongation

Pin point gate

Figure 4-29 Single pin point gate on the spur gear hub wall

Weld line (strong)

Diffuser

Sucker pin Web wall

Melt flow behavior

Figure 4-30 Three pin point or runnerless gates on gear web wall

Diaphragm gate disk Sprue puller Sprue or hot mold runnerless drop Melt flow behavior

Figure 4-31 Sprue diaphragm gate in the middle of spur gear hub

Hub

277

4.4 Thermoplastic Spur Gear Design Requirements

4.4.2

Multifunction Designs with Thermoplastic Gears

The most important advantage of injection molded thermoplastic gears is the possibility of multifunctional designs, reducing the number of components for a given device. Figures 4-32, 4-33, 4-34, and 4-35 show multifunctional design applications. • Figure 4-32 shows a multifunctional gear system; a spur gear made of acetal homopolymer is provided with two molded-in springs, acting on a ratchet wheel that is combined with the second spur gear made of unreinforced nylon 6/6. These types of ratchets function without any problems, as long as the springs in the ratchets are intermittently working and releasing the load in a free stage position, avoiding creep problems of the spring arms. • Figure 4-33 shows a multifunctional gear design made of acetal homopolymer; the gear teeth are protected against impact loads, by connecting the hub and the rim by properly dimensioned ﬂexible elements. This principle is used on printing wheels to obtain consistent printing results in spite of injection molding dimensional discrepancies.

Nylon 6/6 spur gear

Spring arm

Ratchet

Ratchet

Acetal spur gear & spring arm

Figure 4-32 Two spur gears combined with two spring ratchets

Flexible elements

Figure 4-33 Self-centring and flexible spur gear, printing wheel

278

4 Thermoplastic Gearing Design • Figure 4-34 shows a gear design system for a backlash-free motion transmission between two gears. The main gear is equipped with four molded-in springs ﬁtting into corresponding slot pockets on the second gear. When assembled with the pinion, the two tooth crowns are slightly offset, causing the springs to be loaded and thus suppress any backlash. The stress relaxation caused by the creep effects on the spring arms decreases the spring arms’ resistant stresses. This backlash-free gear principle is adequate for only small torques, such as in instrument dials or clock adjusting mechanisms. • Torque limiting devices are often very useful for thermoplastic gear applications to prevent tooth damage when overloading occurs (for instance on high torque transmissions, such as meat grinders, can openers, and hand tool drills). Figure 4-35 shows one solution of many other possible designs. In this torque limiting device, it is essential that the pivoting springs do not remain accidentally in the loaded position. In this design, the torque limiting is achieved by three pivoting springs and radial guiding pockets.

Pinion gear Slot pocket Main gear

Spring arm Second gear

Figure 4-34 Backlash-free thermoplastic gear system (Courtesy: Du Pont)

Pivoting spring

Pocket

Figure 4-35 Torque limiting thermoplastic gear system (Courtesy: Du Pont)

279

4.4 Thermoplastic Spur Gear Design Requirements A product designer must be aware that it is difﬁcult to combine both design conditions (precision and multifunction) in a thermoplastic gear. Either the application requires an accurate gear, using the simple symmetrical shape or it combines as many functions as possible in one part, resulting in a more complicated shape with less accuracy.

4.4.3

Mounting Thermoplastic Gears on Metal Shafts

There are several methods for mounting thermoplastic gears on metal shafts. Some of these methods work well for metal gears, but cause severe problems with thermoplastic gears. Figure 4-36 shows the hubs of a plastic gear drilled and tapped for a metal set screw; the plastic threads can be stripped by the pull out force of the set screw, when the set screw is torqued against the metal shaft. Figure 4-37 shows a metal-threaded insert ultrasonically welded in the gear hub for a metal set screw; the metal insert is stripped from the plastic hub by the pullout force of the set screw when the set screw is torqued against the metal shaft. Recommended methods for mounting plastic gears on metal shafts include coarse knurling the metal shaft and encapsulating the plastic gear around the shaft as shown in Figure 4-38. Figure 4-39 shows a method of drilling a hole through the plastic hub and metal shaft, and then a dowel pin is inserted through the assembly. Figure 4-40 shows a method of drilling and pressing a dowel pin on the shaft, then the two snap locks molded in the plastic gear hub are snapped and locked together for mounting the plastic gear to the dowel pin and metal shaft.

Plastic gear

Set screw

Metal shaft

Hub

Figure 4-36 Set screw threaded on gear hub

Metal threaded insert ultrasonic welded

Hub

Set screw Plastic gear

Metal shaft

Figure 4-37 Ultrasonic welded metal insert and set screw on the gear hub

Plastic gear

4.4.4

Standard Spur Gears, Equations, and Calculations

Knurled shaft

Example 4-2 A pair of gears having a ratio of 3 is to be used at a center distance (C) of 8.48 in. If one gear (NG) has 60 teeth and the other (NP) has 20 teeth, determine the diametral pitch (P).

Metal shaft

Solution Figure 4-38 Encapsulated knurled metal shaft with plastic gear

To calculate the diametral pitch, select Equation 4-7. Diametral pitch (P) = t =

π 2P

Dowel pin Hub

Plastic gear

Dowel pin

Metal shaft

Plastic gear

Snap lock

Metal shaft

Figure 4-40 Dowel pin pressed metal shaft and snap lock hu

Hub

Figure 4-39 Dowel pin drilled through plastic hub and shaft

280

4 Thermoplastic Gearing Design Table 4-4 Standard Spur Gear Equations Knowing the Diametral Pitch (P)

Coarse – Pitch lower than 20

Fine – Pitch higher than 20

Addendum

a=

1.000 P

a=

1.000 P

Dedendum

b=

2.000 P

b=

1.20 + 0.002 P

Working depth

hK =

2.000 P

hK =

2.000 P

Whole depth

ht =

2.250 P

ht =

2.20 + 0.002 P

Circular tooth thickness

t =

π 2P

t =

π 2P

Clearance

c=

0.250 P

c=

0.200 + 0.002 P

Pitch diameter

DP =

N P

DP =

N P

Outside diameter

DO =

N +2 P

DO =

N +2 P

Root diameter

DR =

N − 2.5 P

DR =

N − 2.2 + 0.004 P

Fillet radius

rf =

0.30 P

rf =

0.30 P

Example 4-3 Find the circular tooth thickness on the pitch circle of a 20°, full depth involute tooth having a diametral pitch (P) of 12. The circular tooth thickness is given by the equation in Table 4-4: t =

π 3.1416 = = 0.1309 in 2×P 2 × 12

Example 4-4 Find the outside diameter (DO) on the tooth circle of a 20°, full depth involute tooth having a diametral pitch (P) of 12 and number of teeth (N) of 40. The outside diameter is given by the equation in Table 4-4: (DO ) =

N + 2 40 + 2 = = 3.50 in P 12

281

4.4 Thermoplastic Spur Gear Design Requirements

4.4.5

Spur Gear Pitch Backlash

Figure 4-41 shows a pitch backlash or the tangential clearance between two meshing teeth ﬂanks. The backlash is known as “circumferential play” or “tangential clearance”; backlash is generally deﬁned as the distance by which tooth space exceeds tooth thickness, as measured on the pitch circle. The purpose of backlash is to prevent gears from binding or seizing-up by making simultaneous contact on both sides of meshing teeth. The pitch backlash prevents the mating teeth from binding and it is inﬂuenced by the following factors:

Pitch circles Pitch backlash or tangential clearance

Figure 4-41 Pitch backlash between mating spur gears

• Operating temperatures • Coefﬁcient of linear thermal expansion • Mounting tolerances • Type of gear loading • Center distance tolerances • Tooth shape, size, complexity and accuracy (roundness) • Post molding dimensional changes • Type of lubrication • Run-out tolerances of shaft bearings • Speed and running conditions Gears operating under moderate loads and at moderate speeds at room temperature will be less affected by small variations in backlash. At high load, high speed or high temperatures, gears should have both greater tooth accuracy and additional backlash to compensate for the coefﬁcient of linear thermal expansion. For thermoplastic gears at room temperature, the following backlash values are suggested: This is only a starting point, since the operating temperature is dependent on other factors, such as gear load, running speed, and tooth size. Gears operating at room temperature at high permissible loads and speeds can experience a rise in tooth temperature of 100 °F. Any heat management method at the tooth surfaces will decrease binding caused by insufﬁcient backlash. One method is through continual lubrication. Another is a thermoplastic gear with a metal mate having a higher heat dissipation factor. Backlash may be increased by extending the center distance of mating gears. This provides an advantage of more clearance between the outside diameter of one gear and the base diameter of the other, allowing for expansion at high temperatures without radial interference. However, increasing center distance causes the teeth to mesh outside the pitch circle, which can result in greater wear. Insufﬁcient pitch backlash may go as far as to cause seizing and rapid spur gear failure during operation. In addition, the speciﬁed center distance can be affected by other factors. If, for example, other components of the gear assembly are injection molded of thermoplastics, their dimensions may change over time due to the coefﬁcient of linear thermal expansion or post mold shrinkage from

Table 4-5 Backlash Values for Thermoplastic Gears

Diametral pitch (P)

Pitch backlash (in)

16 20 32

0.004–0.006 0.003–0.005 0.002–0.004

282

4 Thermoplastic Gearing Design high temperatures. The type of gears can also inﬂuence the center distance, as with bearing run-out caused by the axial thrust of helical gears. Therefore, providing pitch backlash by means of increased center distance is a method that should be approached with the utmost caution. 3P

4P

5P

The pitch backlash and center distance measurements and adjustments should be made under actual operating conditions over time, using injection molded prototypes to accurately predict the performance of the thermoplastic gears. It is essential to measure and adjust the correct pitch backlash at operating temperature and under real working conditions. Many gears, even though correctly designed and molded, fail as a result of incorrect backlash at operating conditions. When the gear housing is also injection molded of thermoplastics, the same considerations apply to this product. The values can be different, because the gear and housing have different geometries and the housing may be molded in another resin. Consequently, the center distances may also vary and inﬂuence the pitch backlash. It is often easier to determine the center distance after having produced and measured the gears. It is important to note that this procedure may produce more wear as the gears will no longer mesh exactly on the theoretical pitch circle.

6P

Measuring, testing, and adjusting the pitch backlash using the same end use conditions is essential for a successful thermoplastic gear application.

4.4.6

7P

Standard Spur Gear Tooth Size Selection

The standard spur gear tooth size or the diametral pitch (P) is determined as a function of allowable bending stress and general operating conditions. 8P

From a strictly functional and technical point of view, there is no reason to choose a bigger spur gear tooth size than required. However, the selection of thermoplastic gears is more conservative than for metal gears.

9P

For a given pitch diameter, a smaller spur gear tooth size is often chosen for the following reasons: • Additional load distribution safety

10 P

• Less critical injection molding tolerances • Less sensitivity to thermal variations, post mold shrinkage, and dimensional stability

12 P

Figure 4-42 shows the standard spur or helical gear tooth sizes of the most common diametral pitches (P). These spur or helical gear teeth are drawn using a full size scale; they can be used to quickly scale the overall size and tooth form details of standard spur or helical gears.

14 P

16 P

It can safely be said that the smallest standard spur gear tooth proﬁle that fulﬁls the strength requirements is usually the best solution.

18 P

20 P

Figure 4-42 Standard spur gear tooth size per diametral pitch (P) (full size)

When standard spur gear testing is used for the tooth size selection, accelerated tests at speeds higher than required of a given application are of no value. Increased temperature may cause rapid failure, while under normal working conditions the part may perform well.

4.4 Thermoplastic Spur Gear Design Requirements Test conditions should always be chosen to simulate the real running conditions. The following examples further explain the need for meaningful end use testing. • Appliance thermoplastic gears under a high load that operate only intermittently: the gears should not be tested continuously; use cycles that allow the whole gear assembly device to cool down to room temperature between each test cycle. • Window blinds thermoplastic gears operate infrequently at reduced speeds: these gears can be tested at the same speed continuously; when increasing the tooth surface temperature, the test results are of no value. • Automotive windshield wiper thermoplastic gears reach their maximum working temperature quickly, at which they operate most of their service life. These thermoplastic gears should be tested continuously. Injection molded thermoplastic gears operate very closely to the endurance limit of the resin and the breaking torque test should not be considered as valid in all cases. If the breaking torque proves to be 8 to 10 times the operating load, it can usually be taken as an indication that it will provide a long service life in use. If two mating spur gears have standard circular tooth thicknesses and are brought into close mesh, the distance between their centers will be half the sum of their standard circular pitch. However, two spur gears having standard circular tooth thicknesses can operate at the standard center distance only if both spur gears were perfect. Any errors in the spur gears would cause the tooth face surface to bind at some point in their rotation. The pitch backlash must be checked before all tests. Once a thermoplastic gear has failed during the test, it is almost impossible to determine whether an incorrect pitch backlash was partially or entirely responsible for the failure. The successful development of an injection molded thermoplastic gear requires sound experience, careful product design, detailed study in selecting the best injection molding thermoplastic resin, gear tooth form selection, gear assembly analysis (tolerances), proper gear data (documentation), excellent mold design and construction, correct injection molding process, and meaningful tests.

4.4.7

Standard Gear Total Composite Tolerances

The total composite tolerances of a gear are the summation of the following dimensional variables: • Pitch tolerance • Proﬁle tolerance • Run-out tolerance (roundness) • Lateral run-out tolerance (wobble) The sum of the pitch tolerance and the proﬁle tolerance of a gear is known as the tooth-to-tooth composite tolerance. When the run-out tolerance is added to the tooth-to-tooth composite tolerance, this value is called the total composite tolerance of a gear. When an injection molded thermoplastic gear is loaded in the center, the center distance measuring instrument must be in close mesh with the master spur gear.

283

284

4 Thermoplastic Gearing Design

Tooth to tooth composite Total composite tolerance tolerance

Run-out tolerance

One revolution of test gear

Figure 4-43 Gear total composite tolerance graph test results

By rotating the test gear in close mesh with a master gear of known accuracy in a variable center distance ﬁxture, the run-out tolerance, the tooth-to-tooth composite tolerance, and the total composite tolerance can be determined by measuring and plotting the radial displacements as shown in the general test graph results in Figure 4-43. There are various types of center distance measuring instruments available. The simpler models are equipped with a dial indicator and require the operator to measure and plot the radial displacements as the gear is rotated manually through 360° in close mesh with the master gear. Figure 4-44 shows a manual center distance measuring instrument. The more sophisticated models trace the radial displacements through an electronic device on a moving chart. The American Gear Manufacturers Association has developed a system to classify the dimensional accuracy of standard gear tooth forms by a number in accordance with their maximum tooth-to-tooth and the total composite tolerances allowed for the standard gears. This number is known as the AGMA Quality Number speciﬁed for the standard gear. The AGMA Quality Numbers, the tooth-to-tooth composite tolerances, the run-out tolerances, and the total composite tolerances, corresponding to the diametral pitch, the number of teeth, and the pitch diameter of a standard gear are listed in the American Gear Manufacturers Association Handbook, the Machinery’s Handbook, and other reference manuals. If a standard spur gear is assigned an AGMA Quality Number of Q7, with a diametrical pitch of 20, number of teeth 18, and a pitch diameter is 0.90, to meet these requirements, the standard spur gear maximum run-out tolerance should be 0.0008 in, the maximum tooth-to-tooth composite tolerance 0.0019 in, and the maximum total composite tolerance 0.0027 in. Figure 4-45 shows the measuring instrument test results of an injection molded thermoplastic standard spur gear in close mesh with the master spur gear. The errors in the test gear are at the maximums allowed by these tolerances. Testing gear

Master gear

0.0019 inch

Cent Dista er nce

Figure 4-44 Gear center distance manual measuring instrument

0.0027 inch

Figure 4-45 Measuring instrument spur gear tolerances test results

0.0008 inch

285

4.4 Thermoplastic Spur Gear Design Requirements To allow for the tolerances in two mating gears, either the operating center distance must be made greater than the calculated close mesh center distance by an amount equal to the sum of half the total composite tolerances, or the circular tooth thicknesses must be thinned by an equivalent amount. AGMA Quality Numbers must be chosen for a pair of mating gears at an early stage in the design procedure and the ﬁnished gears must be inspected by being run in close mesh with a master gear, to ensure that the maximum tolerances allowed are not exceeded. Together with the gear drawing, gear data documentation should be a part of the design, where the “gear testing radius” is speciﬁed. The center distance between the gears and the test radius is calculated in Example 4-5. This gear testing radius has maximum and minimum values corresponding to the maximum and minimum values of the calculated circular tooth thickness and the maximum total composite tolerance. Example 4-5 A standard spur gear has 72 teeth, a diametral pitch of 32, and a pressure angle of 20°. The gear is required to have an accuracy corresponding to AGMA Quality Number Q7. The standard pitch diameter is 2.250 in. Calculate the testing radius. Assume that the gear will be inspected by being run in close mesh with a master gear having 64 teeth, a pitch diameter of 2.00 in, and a circular tooth thickness of (1.5708 / 32) = 0.0491 in. From the AGMA it is found that the tooth-to-tooth composite tolerance is 0.0014 in and the total composite tolerance is 0.0032 in. The standard spur gear has a circular tooth thickness of (1.5708 / 32 – 0.0032) = 0.0459 in maximum and (0.0459 – 0.0014) = 0.0445 in minimum. First we need to calculate the center distance between the gears, when the standard spur gear circular tooth thickness is at 0.0459 in and 0.0445 in. C=

(N G + N M ) cos φ 2 × P × cos φ1

(4-8)

Where: C = Close mesh center distance (in) NG = Number of teeth in gear NM = Number of teeth in master gear P = Diametral pitch φ = Pressure angle (degrees) P (t G + t M ) − π φ1 = Angle whose involute is = + inv φ NG + NM tG = Circular tooth thickness of gear (in) tM = Circular tooth thickness of master gear (in) NG = 72, NM = 64, φ = 20°, P = 32, tG = 0.0459 / 0.0445, tM = 0.0491 inv φ1 =

32 (0.0459 + 0.0491) − 3.1416 + 0.014904 = 0.0141564 72 + 64

φ1 = 19.66° , cos φ1 = 0.9417 , C = inv φ1 =

(72 + 64) × 0.939692 = 2.1204 2 × 32 × 0.9417

32 (0.0445 + 0.0491) − 3.1416 + 0.014904 = 0.013827 72 + 64

286

4 Thermoplastic Gearing Design

φ1 = 19.51 , cos φ1 = 0.94258 , C = Close mesh center distance (C) =

(72 + 64) × 0.939692 = 2.1184 2 × 32 × 0.94258

2.1204 in 2.1184

To complete the calculation, half the total composite tolerance is added to the maximum close mesh center distance, half is subtracted from the minimum, and from both results, half the pitch diameter of the master gear is subtracted. 2.1204 + (0.0032/2) − (2.0000/2) = 1.122 2.1184 − (0.0032/2) − (2.0000/2) = 1.116 Testing Radius =

1.122 in 1.116

Figure 4-46 shows the circular tooth thickness and testing radius tolerances between the gears, the measurement results were superimposed for comparison. Figure 4-46 shows the test radius tolerance as the difference between the maximum test radius of a standard spur gear and the minimum test radius of the master gear. The circular tooth thickness of a standard spur gear differs from the master gear by an amount equivalent to the radial tolerance displacement caused by the tooth thickness difference found in both gears. If the total composite tolerance in each gear is less than what is allowed by the tolerance, then the radial tolerance displacement could be greater than the difference between the maximum and minimum calculated circular tooth thicknesses as speciﬁed on the drawing. Therefore, the standard spur gear would be acceptable if the testing radius checked within the maximum and minimum speciﬁed values. This is the reason why the circular tooth thickness of a gear appears on the drawing as a “basic speciﬁcation” rather than being included in the “manufacturing and inspection” data and why it is referred to as the “calculated” circular tooth thickness. It is also the reason why the measurement over pins has a proviso to the effect that this measurement is to be used “for set-ups only”. After choosing the number of teeth for a standard spur gear, the tooth form, the helix angle, and the AGMA Quality Number and after having determined the tooth thickness, the remaining gear data is obtained by a mathematical computation of the gear assembly system, including the various tolerances required for the design.

Maximum testing radius Maximum center distance Radial tolerance caused by tooth thickness difference Minimum center distance Minimum testing radius

Figure 4-46 Measuring test tolerances comparison between gears

Testing radius tolerance

4.5 Tolerances and Mold Shrinkage of Thermoplastic Gears Specifying the circular tooth thickness and accuracy of a thermoplastic gear with a testing radius precludes any possibility of misinterpretations and makes inspection a simple and quick operation. If the gear checks within the maximum and minimum values speciﬁed for the testing radius and if it satisﬁes the total composite and tooth-to-tooth maximum tolerances, the gear design is correct. Because bevel gears and worm gears are commonly manufactured as mating pairs, designing and specifying injection molded thermoplastic gears present problems not encountered with spur and helical gears. The gear project requires close working communication between the gear mold maker and the manufacturer of the metal pinion or worm before completing the design.

4.5

Tolerances and Mold Shrinkage of Thermoplastic Gears

Molded thermoplastic gears are as accurate as metal machined gears. Whether a molded thermoplastic gear or metal machined gear is characterized by AGMA Quality Number Q8, both gears meet the same tolerances. However, thermoplastic gears have not yet been molded to the highest precision obtainable by machining; gears requiring such precision represent a very small percentage of the gear markets. Fine-pitch instrument gears are molded to tolerances that would have been considered impossible to achieve in the past. For example, a four-cavity mold for a ﬁne-pitch spur gear made of acetal homopolymer produced gears that were within the tolerances of AGMA Quality Number Q12. The tooth-to-tooth composite tolerance was 0.0003 in, with a total composite tolerance of 0.0005 in or less. All materials shrink when they are transformed from a solid cold state to a hot transitional liquid melt and back again to a solid state, as the material cools off to room temperature. As a consequence, all mold cavities must be made larger than the molded gears to compensate for the mold shrinkage effects caused by the thermoplastic polymers. For example, if a molded gear is to have an outside diameter of 1.500 in and the thermoplastic resin has a mold shrinkage of 0.020 in/in, the outside diameter of the cavity must be 1.530 in. In the design and fabrication of an injection mold for thermoplastic resins, the gear cavity for the mold should be constructed using different fabricating procedures than those used to cut the teeth in the gear if it were to be machined. It is necessary to cut an oversized mold gear cavity that, in turn, would be used to transform the melt into the correct size gear. If the gear cavity is not oversized, it will result in a smaller gear with serious dimensional problems. Figure 4-47, left illustration, shows an enlargement of a standard spur gear tooth form (dotted line) having a diametral pitch of 32, a pressure angle of 20°, and the proﬁle of a standard oversize spur gear tooth form. The right illustration shows the proﬁle of an injection molded thermoplastic spur gear tooth form obtained from an oversize gear cavity and a standard spur gear tooth form (dotted line). The injection molded thermoplastic spur gear tooth form is considerably different from the standard spur gears. It is thicker at the root and thinner at the tip; it has a larger pressure angle of 25°, as shown in Figure 4-47, right illustration.

287

288

4 Thermoplastic Gearing Design Standard spur gear tooth form

Standard oversize spur gear tooth form Oversize mold cavity thermoplastic gear

Figure 4-47 Different spur gear tooth forms comparison

It is obvious that the mold shrinkage of the teeth in the tool cavity must be carefully compensated for so that, when the molded thermoplastic gear solidiﬁes and becomes stable, the gear will have the correct tooth form. The mold design of a helical gear is more complicated; the axial (cross ﬂow direction) mold shrinkage is quite different from the radial (melt ﬂow direction) mold shrinkage. Compensating correctly for the mold shrinkage in a gear requires the mold designer to thoroughly understand gear geometry, gating, mold venting, cooling, processing, and to have considerable experience with the mold shrinkage behavior of thermoplastics. The importance of correctly compensating for the mold shrinkage of the thermoplastic resins cannot be over-emphasized. For example, the gear center distance measuring instrument tolerance test results shown in Figure 4-45 show that this gear can have a total composite tolerance of 0.0027 in and a tooth to tooth composite tolerance of 0.0019 in. If the tooth-to-tooth composite tolerance is at the maximum, the run-out tolerance must be held to 0.0008 in. But if the tooth-to-tooth composite tolerance is reduced to 0.0005 in, an amount easily achieved if the mold gear cavity is correctly designed and accurately constructed, the run-out tolerance can go as high as 0.0022 in. Because it is more difﬁcult to control run-out tolerance than tooth-to-tooth composite tolerance, it is important that the thermoplastic molded gear tooth form be as accurate as possible (several degrees more accurate than a comparable machined metal gear). Mold shrinkage does not need to affect accuracy other than for the very minor variations in mold shrinkage that occur during the production molding run. It is not an uncommon practice for product designers to specify close tolerances for the outside diameters of thermoplastic gears and leave everything else wide open. This is probably done in the mistaken belief that the outside diameter of a molded gear is a measure of overall accuracy and therefore this is the easiest dimension to measure. In fact, this close outside diameter tolerance ensures that all the tolerances present in a lot of molded gears will have the same tolerance magnitude from one gear to the next gear. Except in rare cases, the outside diameter of a gear is, within limits, a matter of no consequence. If it is speciﬁed that the circular tooth thickness of a molded gear is to be held to +0.000/–0.001 in, then the outside diameter must be allowed to vary within a tolerance band of at least 0.0027 inch in the case of 20° pressure angle gears for AGMA Quality Numbers Q7. To specify close tolerances for the outside diameter of a molded gear, except where the outside diameter is functional as in pump gears, these close tolerances can make for unnecessarily high tooling costs and increase the molding cost of the gear, without obtaining the performance required for the gears. It cannot be emphasized strongly enough that the accuracy of thermoplastic gears should be speciﬁed for AGMA Quality Numbers and should be inspected by the center distance measuring method.

289

4.6 Standard Helical Gears

4.6

Standard Helical Gears

A helical gear is essentially a spur gear with teeth slanting across its face in a cylindrical spiral, or helix, to the axis. Helical gears are preferred over spur gears in many applications because of their smoother, quieter operation with fewer tendencies to squeak. However, they require not only perfect tooth proﬁles but also exactly matching helix angles. These requirements can be difﬁcult to fulﬁl, particularly when mating gears are made of dissimilar materials. Helical gears generate axial thrust that may create problems. It is advisable to use helix angles not greater than 15°. Compared to a spur gear having the same tooth size, a helical gear provides slightly improved tooth strength. As small helix angles are most commonly used, this fact can be neglected when determining the diametral pitch and it should only be considered as an additional safety factor. Helical gears have teeth that are formed on a spiral that winds around the axis of the shaft running through the gear (as opposed to spur gear teeth that are formed parallel to this axis). They are high efﬁciency gears (98–99%) and are typically used when high speeds and high horsepower are involved. The gear helix may be either left or right handed and can have various helix angles. These angles cause the pair of gears to exert end thrust on the bearings that carry the shafts on which the gears are mounted. Provisions should be made to compensate for thrust in the bearings. The design equations for helical gears are similar to those for spur gears, with a modiﬁcation to account for the helix angle. This overlap is what gives helical gears their excellent smooth and quiet operation. The involute tooth proﬁle is normally used for helical gears as small variances in center distance will not affect tooth action. Helical gears require perfect tooth proﬁles and exactly equal helix angles for proper performance. Table 4-6 Basic Helical Gear Equations for Known Diametral Pitch (P)

Find

Rules

Equations

Addendum

Divide 1 by the pitch diameter

a=

Lead to tooth helix

Multiply the pitch diameter by 3.1416 by cot of tooth helix angle

L = π × DP × cot ψ

Center distance

Add the pitch diameters of both gears and divide by 2

C=

DP1 + DP2 2

Whole depth of tooth

Divide 2.157 by the diametral pitch

ht =

2.157 P

Tooth thickness at pitch line

Divide 1.571 by the diametral pitch

t =

Pitch diameter

Divide the number of teeth by pitch dia. × cos of tooth helix angle

DP =

Outside diameter

Add twice the addendum to the pitch diameter

DO = DP + 2 × a

1.000 P

1.571 P N P × cos ψ

290

4 Thermoplastic Gearing Design

4.7

Standard Straight Bevel Gears

Standard straight bevel gears are widely used in applications involving power transmission, for right angle drives, and for providing high efficiency in operation. They may also be used to transmit power between shafts at any angle. Bevel gears can be made with either straight or spiral teeth, which taper in both thickness and height to almost zero at the axis of the gear. Figure 4-48 shows a standard straight bevel gear; the teeth are formed with all elements on planes that intersect at the axis of the gear. Bevel gear teeth are tapered in both thickness and height, with the outer portion (the heel) longer than the inner part (or toe). Table 4-7 is used for calculating the common variables required for the pinion and gear. Bevel gears exert both axial thrust and radial loads on the shaft-support bearings. If bevel gears have too few teeth, undercutting can be a problem. Bevel gear geometry can lead to molding and/or functional problems. On heavily loaded large bevel gears, thrust load on the tooth crown may become considerably stressed and the use of ribs is not recommended.

Pitch apex to back Pitch apex to crown

Crown to back

Crown Dedendum angle Pitch apex

ce

tan

is ed

Face angle

Shaft angle

n Co

PINION

Roo ang t le Pitch angle

Face width

GEAR Uniform clearance

Pitch diameter

Back cone

Ba ck co ne dis tan ce

Outside diameter

Figure 4-48 Standard system straight bevel gear nomenclature

291

4.7 Standard Straight Bevel Gears Table 4-7 20° Straight Bevel Gear for 90° Shaft Angle Equations

Item

Pinion

Gear

Working depth

hK =

2.000 P

hK =

2.000 P

Whole depth

ht =

2.188 + 0.002 P

ht =

2.188 + 0.002 P

Pitch diameter

DP =

Pitch angle

β = tan −1

Cone distance

CD =

Circular pitch

p=

Addendum

aP = hK − aG

Dedendum

bP =

Clearance

c = ht − hK

Dedendum angle

δP = tan −1

Face angle

γ F = γ + δG

ϕF = ϕ − δP

Root angle

γ R = γ − δP

ϕR = ϕ − δG

Outside diameter

DOP = DP + 2 aP cos γ

DOP = DG + 2 aP cos γ

Pitch apex to crown

XP =

Circular thickness

tP = p − tG

Chordal addendum

aCP = aP +

Tooth angle

3438  t P   + bP tan φ  CD  2

3438  t P   + bP tan φ  CD  2

Limit piont width

CD − F (t G − 2 bP tan φ) CD

CD − F (t G − 2 bP tan φ) CD

NP P

NG P

DG = NP NG

DP 2 sin ϕ

π P

ϕ = 90° − β

CD =

p=

2.188 − aP P

DP 2 sin ϕ

π P

bG =

2.188 − aG P

c = ht − hK

bP CD

DG − aP sin γ 2

δG = tan −1

XG =

tG = t P2 cos γ 4 DP

bG CD

DP − aG sin γ 2 p − (aP − aG ) tan φ 2

aCG = aG +

t G2 cos ϕ 4 DG

292

4 Thermoplastic Gearing Design

4.8

Unreinforced nylon 6/6 worm pinion

Standard Worm Gears

Most machined worm gears are made with a throated shape that provides a contact line of a certain length on the worm. Because this system cannot easily be applied on injection molded thermoplastic gears, a simple helical gear is normally used. Consequently, the load is transmitted on very small contact points that could lead to excessive pressure, surface temperature, and wear. Therefore, a helical gear meshing with a worm has limited possibilities and should be investigated carefully. Various attempts have been made, aimed at improving wear and increasing power transmission by changing the contact points to contact lines. Figure 4-49 shows a practical application demonstrating the use of an unreinforced nylon 6/6 for the worm pinion and acetal homopolymer for the worm gear.

Acetal homopolymer one piece worm gear

Figure 4-49 One-piece thermoplastic worm gearing system

Nine radial slide cores

Figure 4-49 shows a one-piece injection molded worm gear made of high viscosity acetal homopolymer, meshing with a worm pinion made of unreinforced nylon 6/6 for a hand operated device. The undercut resulting from the throated shape amounts to about 4% and can therefore be ejected from the mold without problems. It is noteworthy that this particular worm with seven leads cannot be molded in a two plate mold with the parting line in the center. Because the lead angle of 31° being greater than the typical pressure angle of 20°, this results in an undercut along the parting line. Therefore, the worm must be unscrewed from the mold. This principle of injection molding and ejecting a one-piece worm gear from the mold is used in several applications, even though it requires good processing experience and skill to design and to construct the proper tool. Figure 4-50 shows an automotive windshield wiper gear made of high viscosity acetal homopolymer, injection molded in a different way. Because of the undercut of about 7% and the rigid structure, ejecting the worm gear from the mold becomes impossible. To compensate for this deﬁciency, the mold is made with nine radial slide cores, each of which covers six teeth. This injection molding procedure produces an excellent worm gear, but the worm gear mold only has a single cavity, causing rather high tooling and manufacturing costs. Figure 4-51 shows another automotive windshield wiper with a half throated worm gear that is based on an intermediate solution. It is composed of a half throated and a helical gear portion. The tooth contact takes place on the curved section, while the helical part merely improves the tooth strength and the stall torque. Though not ideal, this solution nevertheless offers a signiﬁcant advantage over a simple standard helical gear.

Figure 4-50 Worm gear mold with nine radial slide cores

Figure 4-52 shows a split snap-ﬁtted worm gear made of acetal homopolymer. The two worm gear halves are designed in such a way that each component is injection molded using the same mold cavity. The two worm gear halves are post-mold assembled by rotating 180° until facing each other and pressing both components until they are snap-ﬁtted together. The snap ﬁt mechanism provides the centering action with the teeth perfectly aligned by means of lugs ﬁtting into corresponding holes. A single cavity provides a complete gear assembly that is held together by means of snap ﬁts, ultrasonic welding, or rivets. The worm gears can be made as wide as necessary, limited only by proper meshing. Split snapﬁtted worm gears are especially recommended for larger worm gear diameters that require high performance in the end use application.

Figure 4-51 Half throated worm gear

The following limitations should be kept in mind when designing split snapﬁtted worm gears compared to standard helical gears:

293

4.8 Standard Worm Gears • Higher tooling cost • Requirement of perfect centering of the worm pinion and the gear. Even small displacements cause the load to be carried by only a portion of the tooth width, resulting in increased wear or rapid failure. • The worm gear drive is more sensitive to discrepancies of the lead angles, which must match perfectly. • The worm pinion and the gear must be assembled in a certain way. If, for instance, the worm pinion is mounted ﬁrst into the housing, the split snapﬁtted worm gear or the half throated worm gear can only be added in a radial direction, while a standard helical gear can be mounted from the side.

4.8.1

Radial snap hole Radial snap arm

Standard Worm Gear Analysis

Except for very slow running and hand operated worm gear drives, the power transmission is essentially limited by the heat build-up on the tooth surfaces. Temperature rise is not only a result of speed and load, but is also inﬂuenced by other factors, such as overall design, heat dissipation through the housing, additional cooling by external fan blowers, or heat being carried into the gear from the electric motor. Another important factor is the efﬁciency of the initial lubrication, which very often determines the total service life. For the type of worm gears commonly used with thermoplastic materials (a single threaded worm and a helical worm gear), point contact occurs between the meshing teeth and limits the ability of these gear systems to carry high loads. This point contact can be increased by the various worm gear designs discussed previously. Since the teeth on the helical worm gear are weaker than the threads of the worm gears, the maximum output torque of the worm will be limited by the torque capacity of the helical worm gear. The same common equations presented for helical gears can be applied for worm gears. A liberal safety factor (2–3) should be applied to take into account the stress concentration caused by theoretical point contact. An additional problem in worm gearing is the high sliding velocity that increases tooth temperatures and results in a lower yield strength and higher tooth wear rate. Therefore, it is suggested that the use of acetal homopolymer in the worm pinions or worm gears be restricted to low loading and reduced rubbing velocities of less than 250 feet per minute for these applications. For moderate operating conditions, a worm gear made of acetal homopolymer may be used successfully with a metal worm pinion. As a guide, it is suggested that the maximum rubbing velocity of the worm gear be limited to 250 feet per minute for continuous operation and initial lubrication. The equation used to determine the rubbing velocity is: VR =

Left half worm gear

DP × n 12 × cos α

Where: VR = Rubbing velocity (fpm) DP = Worm pitch diameter (in) n = Worm speed (rpm) α = Lead angle (degrees)

(4-9)

Figure 4-52 Split snap-fitted acetal homopolymer worm gear

Right half worm gear

294

4 Thermoplastic Gearing Design A rubbing velocity of about 500 feet per minute may be used for worm gears made of acetal homopolymer, when continuously lubricated worm gears or intermittent operations have been provided for the gearing application.

4.10

Plastic Gearing Technology Designs

Designing injection molded thermoplastic gears and establishing the manufacturing and quality control requirements for the production gears involves spending an extensive amount of engineering time working on complex gearing calculations. The designer will need several gear handbooks and standards, thermoplastic material speciﬁcations and molding characteristics, a set of mathematical tables, a computer or a calculator, and countless hours to spend trying to arrive at a satisfactory gear design. The design of injection molded thermoplastic gears requires consideration of the following important factors: • The power to be transmitted • The gear ratio required • The selection of the type of gear and the tooth design geometry • The type of thermoplastic resin for injection molding • The tolerances required for the gear • The center distance dimension tolerances and alignment for the housing to be manufactured, the bearing run-out tolerances • The end use temperature and the type of lubrication The ABA Tool & Die Company, Inc. and its afﬁliate Plastic Gearing Technology, Inc. from Manchester, Connecticut have developed four tooth form design systems known as PGT-1, PGT-2, PGT-3, and PGT-4 in addition to the standard gear tooth forms for thermoplastic gears. Plastic Gearing Technology Inc. has also developed a Plastics Gearing Design Manual and a PGT Computer Program. The company also offers thermoplastic gear and mold design services, precision mold making, and custom molding services. The PGT gear tooth form design systems are used only for spur and helical gears. The required equations, several gear design example calculations, and the gear design drawing speciﬁcations are presented for the PGT systems. The injection molding process provides gear design freedom. The product designers are free to introduce any gear tooth design geometries and choose the correct diametral pitch for the application. These departures from the standard gear design do not increase the manufacturing costs of the injection molded gears. When a gear tooth deﬂects under load, the trailing tooth is out of position for truly smooth engagement with the oncoming tooth of the mating gear. All thermoplastic gear tooth designs should be given a degree of tip relief. The teeth are gradually thinned from halfway up the addendum to the tip, providing a relieved effect. The gear teeth are also given a full ﬁllet radius between two teeth at the root. This full ﬁllet root radius can increase the fatigue strength and service performance of the thermoplastic gear using the PGT systems.

4.10 Plastic Gearing Technology Designs

4.10.1

Spur and Helical Gears PGT-1 Tooth Design

The standard ﬁne-pitch metal gears provide a relatively greater amount of clearance between the tooth tips and the roots of mating standard metal gears, than the standard coarse pitch gears. Greater clearance is required for the standard ﬁne pitch gears to improve the gear wear resistance of the tips after prolonged use. PGT-1 tooth design for spur and helical thermoplastic gears functions perfectly with a standard mating gear tooth form having either a ﬁne pitch or a coarse pitch. The ﬁne pitch gears include all gears of 20 diametral pitch (P) and higher, the coarse pitch gears include all gears with diametral pitches (P) of less than 20. Figure 4-53 shows the spur and helical PGT-1 tooth form design for 20° pressure angle, together with the PGT-1 tooth form equations that were developed by Plastic Gearing Technology Inc. for calculating the various gear dimensions. PGT-1 tooth form ﬁne-pitch gears used for instrument gear applications should have teeth that are longer than the standard gear tooth form. The relatively large coefﬁcients of linear thermal expansion of injection molded thermoplastic gears make the use of longer PGT-1 tooth forms mandatory for most thermoplastic gears employed in instrument movements. A pair of PGT-1 tooth form ﬁnepitch mating gears must be designed so they will not bind at the highest end use temperatures. Injection molded thermoplastic PGT-1 tooth form ﬁne-pitch gears are used in electric clocks, control mechanisms, meters, cameras, and similar applications. Tables 4-8, 4-9, 4-10, and 4-11 show the precalculated values for the minimum circular tooth thickness (t), outside diameter (DO), and root diameters (DR) per the most common number of teeth (N) that were developed using a 1.0 diametral pitch (P). For other gears of different diametral pitches, divide the selected value by the diametral pitch speciﬁed.

295

296

4 Thermoplastic Gearing Design 1

DP

a=

= 20°

2P

90°

1 P

1.0469 h= P

ht =

2.33 P

R= 4 p rf=

0.43 P

b=

1.33 P

p=2xt 2.3329 - (0.0426 x N) P 1 DO = [(N - 2.3158) + 2.7475 x t x P], D P = D O - 2 a, P 1 DR = [(N - 6.9758) + 2.7475 x t x P] P (Up to 18 teeth) t =

P = Diametral pitch, p = Circular pitch, 20° = Pressure angle (φ), a = Addendum, b = Dedendum, ht = Whole depth, rf = Fillet radius, h = Depth straight tangent with fillet radius, t = Circular tooth thickness, N = Teeth number, DR = Root diameter, DO = Outside diameter, DP = Pitch diameter, R = Tangent relief tooth tip radius. Figure 4-53 Thermoplastic gear PGT-1 tooth design (Courtesy: Plastic Gearing Technology Inc.)

Table 4-8 PGT-1 Thermoplastic Gear Tooth Design Data (Courtesy: Plastic Gearing Technology Inc.)

Number of teeth (N)

Min. circular tooth thickness (t × P) in

Outside diameter (DO × P) in

Root diameter (DR × P) in

6 7 8 9 10 11 12 13 14 15 16 17 18

2.0773 2.0347 1.9921 1.9495 1.9069 1.8643 1.8217 1.7791 1.7365 1.6939 1.6513 1.6087 1.5708

8.9254 9.9477 10.9578 11.9577 12.9234 13.8064 14.6893 15.5723 16.4553 17.3382 18.2212 19.1041 20.0000

4.7316 5.6145 6.4975 7.3805 8.2634 9.1464 10.0293 11.9123 11.7952 12.6782 13.5611 14.4441 15.3400

297

4.10 Plastic Gearing Technology Designs

4.10.2

Spur and Helical Gears PGT-2 Tooth Design

1

a=

= 20°

2P

DP

90° h=

1.15 P

1.248 P

ht =

2.63 P

R= 4 p rf=

0.352 P

b=

1.48 P

p=2xt t=

2.4793 - (0.0426 x N) (Up to 22 teeth) P

DO =

1 [(N - 2.0158) + 2.7475 x t x P], P

DR =

1 [(N - 7.2758) + 2.7475 x t x P] P

D P = D O - 2 a,

P = Diametral pitch, p = Circular pitch, 20° = Pressure angle (φ), a = Addendum, b = Dedendum, ht = Whole depth, rf = Fillet radius, h = Depth straight tangent with fillet radius, t = Circular tooth thickness, N = Teeth number, DR = Root diameter, DO = Outside diameter, DP = Pitch diameter, R = Tangent relief tooth tip radius Figure 4-54 Thermoplastic gear PGT-2 tooth design (Courtesy: Plastic Gearing Technology Inc.)

Table 4-9 Thermoplastic Gear PGT-2 Tooth Design (Courtesy: Plastic Gearing Technology Inc.)

Number of teeth (N)

Min. circular tooth thickness (t × P) in

Outside diameter (DO × P) in

Root diameter (DR × P) in

6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

2.2237 2.1811 2.1385 2.0959 2.0533 2.0107 1.9681 1.9255 1.8829 1.8403 1.7977 1.7551 1.7125 1.6699 1.6273 1.5847 1.5708

9.0970 10.1247 11.1398 12.1446 13.1405 14.1287 15.3915 16.0852 17.0549 18.0196 18.9234 19.8064 20.6893 21.5723 22.4552 23.3382 24.3000

4.8338 5.7168 6.5997 7.4827 8.3656 9.2486 10.1316 11.0145 11.8975 12.7804 13.6634 14.5463 15.4293 16.3123 17.1952 18.0782 19.0400

298

4 Thermoplastic Gearing Design

4.10.3

Spur and Helical Gears PGT-3 Tooth Design

1 2P

DP

= 20°

a=

90°

R=

4

1.25 P

1.380 h= P

ht =

2.83 P

p rf=

.30 P

b=

1.58 P

p=2xt t=

DO =

2.5768 - (0.0426 x N) (Up to 24 teeth) P

1 [(N - 1.8158) + 2.7475 x t x P], P

D P = D O - 2 a,

1 [(N - 7.4758) + 2.7475 x t x P] P P = Diametral pitch, p = Circular pitch, 20° = Pressure angle (φ), a = Addendum, b = Dedendum, ht = Whole depth, rf = Fillet radius, h = Depth straight tangent with fillet radius, t = Circular tooth thickness, N = Teeth number, DR = Root diameter, DO = Outside diameter, DP = Pitch diameter, R = Tangent relief tooth tip radius DR =

Figure 4-55 Thermoplastic gear PGT-3 tooth design (Courtesy: Plastic Gearing Technology Inc.)

Table 4-10 Thermoplastic Gear PGT-3 Tooth Design (Courtesy: Plastic Gearing Technology Inc.)

Number of teeth (N)

Min. circular tooth thickness (t × P) in

Outside diameter (DO × P) in

Root diameter (DR × P) in

6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24

2.3212 2.2786 2.2360 2.1934 2.1508 2.1082 2.0656 2.0230 1.9804 1.9378 1.8952 1.8526 1.8100 1.7674 1.7248 1.6822 1.6396 1.5970 1.5708

9.3103 10.2413 11.2597 12.2677 13.2666 14.2577 15.2419 16.2200 17.1924 18.1598 19.1226 20.0810 21.0355 21.9863 22.9231 23.8061 24.6890 25.5720 26.5000

4.9017 6.1447 6.6676 7.5466 8.4335 9.3165 10.1994 11.0824 11.9653 12.8483 13.7313 14.6142 15.4792 16.3801 17.2631 18.1460 19.0290 19.9120 20.8400

299

4.10 Plastic Gearing Technology Designs

4.10.4

Spur and Helical Gears PGT-4 Tooth Design

1 2P

= 20°

DP

a=

1.35 P

90°

R= 4 p

rf=

0.248 P

h=

1.520 P

b=

ht =

3.03 P

1.68 P

p=2xt t= DO =

2.6751 - (0.0426 x N) (Up to 26 teeth) P

1 [(N - 1.6158) + 2.7475 x t x P], P

D P = D O - 2 a,

1 [(N - 7.6758) + 2.7475 x t x P] P P = Diametral pitch, p = Circular pitch, 20° = Pressure angle (φ), a = Addendum, b = Dedendum, ht = Whole depth, rf = Fillet radius, h = Depth straight tangent with fillet radius, t = Circular tooth thickness, N = Teeth number, DR = Root diameter, DO = Outside diameter, DP = Pitch diameter, R = Tangent relief tooth tip radius DR =

Figure 4-56 Thermoplastic gear PGT-4 tooth design (Courtesy: Plastic Gearing Technology Inc.) Table 4-11 Thermoplastic Gear PGT-4 Tooth Design (Courtesy: Plastic Gearing Technology Inc.)

Number of teeth (N)

Min. circular tooth thickness (t × P) in

Outside diameter (DO × P) in

Root diameter (DR × P) in

6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26

2.4195 2.3769 2.3343 2.2917 2.2491 2.2065 2.1639 2.1213 2.0787 2.0361 1.9935 1.9509 1.9083 1.8657 1.8231 1.7805 1.7379 1.6953 1.6527 1.6101 1.5708

9.3236 10.3580 11.3795 12.3907 13.3926 14.3867 15.3737 16.3545 17.3296 18.2996 19.2650 20.2260 21.1830 22.1363 23.0861 24.0326 24.9760 25.9164 26.8541 27.7891 28.7000

4.9718 5.8547 6.6377 7.5936 8.5036 9.3826 10.2695 11.1525 12.0354 12.9184 13.8013 14.6843 15.5673 16.4502 17.3332 18.2161 19.0991 19.9820 20.8650 21.7479 22.6400

300

4 Thermoplastic Gearing Design PGT-4

PGT-3

PGT-2 PGT-1 DO STANDARD

DP t

rf DR (Scale = 25.00

1.00)

Gear of 20 (P) diametral pitch, 13 (N) teeth and 20° ( ) pressure angle

Figure 4-57 Standard vs. PGT system spur gear tooth comparison

Table 4-12 Standard and PGT Gear Tooth Form Comparison (Scale = 25.00 : 1.00) Gear of 20 (P) diametral pitch, 13 (N) teeth, and 20° (φ) pressure angle

Tooth form

Outside Circular tooth dia. (DO) thickness (t)

Pitch dia. (DP)

Root dia. (DR)

Fillet radius (rf )

Standard

0.07854

0.7500

0.6500

0.5250

0.0150

PGT-1 PGT-2 PGT-3 PGT-4

0.08890 0.09630 0.10115 0.10606

0.7786 0.8042 0.8110 0.8177

0.6786 0.6892 0.6860 0.6827

0.5456 0.5507 0.5541 0.5576

0.0215 0.0176 0.0150 0.0124

4.10.5

Plastic Gearing Technology Tooth Form Design Variables

Circular Tooth Thickness (t) The tooth thickness of a standard spur gear is always speciﬁed as being the circular tooth thickness on the standard pitch circle. But, for the PGT tooth form system, the circular tooth thickness will depend on the number of teeth times a constant divided by the diametral pitch. The PGT tooth form systems increase in thickness based on the incremental increase of the constant; therefore, the circular tooth thickness for PGT-4 is wider than PGT-3, PGT-2, or PGT-1, as shown in Figure 4-57. The addendum of each PGT tooth form is lengthened by a constant divided by the diametral pitch. The PGT-4 has a longer addendum, while PGT-1 has a shorter addendum; however, longer than the standard gear. In other words, the longer the addendum, the thicker the circular tooth thickness.

4.10 Plastic Gearing Technology Designs There is a direct relationship between the circular tooth thickness speciﬁed for a spur gear and the outside and root diameters. This relationship can best be explained by working through a calculation example.

Example 4-6 A spur gear has 16 teeth, a diametral pitch of 24 and the PGT-1 tooth design. The PGT-1 circular tooth thickness is 1.6513 ÷ 24 = 0.0688 in (using Table 4-8). Calculate the outside and root diameters. 1) Determine the standard pitch circle diameter (see Table 4-4). N = 16, P = 24, Standard pitch circle diameter (DP) = N/P = 16 / 24 = 0.6666 2) Determine the standard addendum (see Table 4-4). Standard addendum (a) = 1/P = 1 / 24 = 0.04166 in 3) Determine the standard circular tooth thickness (see Table 4-4). Standard circular tooth thickness (t) = π / (2 × P) t = 3.1415926 / (2 × 24) = 0.06545 in 4) The PGT-1 circular tooth thickness speciﬁed is 0.0688 in. The increase over the standard is 0.0688–0.06545 = 0.00335 in. To achieve this increase, the line of action is held back by 0.00335 ÷ 2 tan φ, where φ is the pressure angle. The pressure angle of the PGT-1 tooth design is 20°, therefore the line of action is held back by the following amount: 0.00335 / (2 × 0.36397023) = 0.0046 in 5) The standard addendum is 0.04166 in. The addendum corresponding to the speciﬁed PGT-1 circular tooth thickness of 0.0688 in is the standard addendum plus the amount of the line of action that has been held back. Addendum (a) = 0.04166 + 0.0046 = 0.04626 in 6) Determine the outside diameter (DO). The outside diameter of a gear is the standard pitch diameter (DP) plus two addendums (a). Outside diameter (DO) = 0.666 + (2 × 0.04626) = 0.7585 in 7) Determine the root diameter (DR). The root diameter (DR) is the outside diameter (DO) minus two whole depths (ht). Find the PGT-1 whole depth (ht) by using the equation in Figure 4-53 Whole depth (ht) = 2.33 / P = 2.33 / 24 = 0.09708 in Root diameter (DR) = 0.7585 – (2 × 0.09708) = 0.5643 in Generating the teeth by a line of action is only one of a number of methods employed to form involute teeth, but the relationship between the circular tooth thickness, the outside and root diameters applies no matter how the teeth are formed.

301

302

4 Thermoplastic Gearing Design Given the circular tooth thickness of a gear having one of the four PGT tooth form designs, the outside and root diameters are readily obtained by using the appropriate equations given in Figures 4-53, 4-54, 4-55, and 4-56). Example 4-7 Determine the outside and root diameters of the gear in the previous Example 4-6 by using PGT-1 equations. This gear has (N) 16 teeth, a diametral pitch (P) of 24, and a circular tooth thickness (t) of 0.0688 in. P = 24, N = 16, t = 0.0688 in Outside diameter (Figure 4-53) DO = (1/P) × [(N – 2.3158) + 2.7475 × t × P] DO = (1/24) × [(16–2.3158) + 2.7475 × 0.0688 × 24] = 0.7592 in Root diameter (Figure 4-53) DR = (1/P) × [(N – 6.9758) + 2.7475 × t × P] DR = (1/24) × [(16–6.9758) + 2.7475 × 0.0688 × 24] = 0.565 in

4.10.6

Maximum Allowable Outside Diameter DO (Max.)

If the two involute curves forming the proﬁle of a tooth are continued out, they will eventually cross and the tooth will be pointed. If a gear has a small number of teeth and an enlarged tooth thickness, the teeth may become pointed at a diameter less than the outside diameter given by the DO equation. To avoid specifying an outside diameter impossible to attain, Equation 4-10 is used as a check. The smaller value obtained by calculating DO and DO (Max.) is the outside diameter speciﬁed. Equation 4-10 provides the maximum outside diameter that will still provide the teeth with an adequate top land for all four PGT tooth form designs. DO (Max.) =

N × 0.93969262 P × 1.017 × cos φ1

(4-10)

Where: DO (Max.) = Maximum allowable outside diameter (in) N = Number of teeth P = Diametral pitch t = Circular tooth thickness (in) (t × P) φ1 = Angle whose involute is = N + 0.01490438 Example 4-8 A spur gear has 10 teeth (N), a diametral pitch of 44 (P), PGT-1 tooth design, and a circular tooth thickness (t) of 0.0433. Find the outside and the root diameter. Outside diameter (Figure 4-53) DO = (1/P) × [(N – 2.3158) + 2.7475 × t × P] DO = (1/44) × [(10–2.3158) + 2.7475 × 0.0433 × 44] = 0.2936 in

303

4.10 Plastic Gearing Technology Designs

Maximum outside diameter DO (Max.) =

N × 0.93969262 P × 1.017 × cos φ1

Inv φ1 = (t × P) / (N + 0.01490438) = (0.0433 × 44) / (10 + 0.01490438) = 0.1902364 φ1 = 43°32′, cos φ1 = 0.72497 DO (Max.) =

10 × 0.93969262 = 0.2896 in 44 × 1.017 × 0.72497

The outside diameter to be speciﬁed should be the smaller calculated dimension obtained by these two equations (0.2896 in). DR = (1/P) × [(N – 6.9758) + 2.7475 × t × P] (Figure 4-53) DR = (1/44) × [(10–6.9758) + 2.7475 × 0.0433 × 44] = 0.1878 in

4.10.7

Spur Gear Tooth Form Comparison

If a standard gear has a small number of teeth, the base circle will be greater than the root diameter. Because no tooth action can take place below the base circle, that section of the tooth inside the base circle is nonfunctional. To accommodate the tip of the mating tooth, the nonfunctional lower portion is undercut. Undercutting the teeth is not recommended, it weakens the teeth and causes premature wear of the gear. Figures 4-58 and 4-59 show two separate spur gear tooth forms. Both gears have the same number of teeth and the same diametral pitch. The gear tooth in Figure 4-58 has the standard circular tooth thickness (t = π / 2 P). For the gear tooth in Figure 4-59 (PGT-1), the circular tooth thickness has been increased above the standard. The PGT-1 tooth is a better design; it has a more functional proﬁle and the undercut is reduced signiﬁcantly. The equations to determine the minimum circular tooth thickness, shown in Figure 4-53, were developed to ensure adequate involute proﬁles and to avoid the objectionable undercutting of the teeth in the gears.

t Standard

Figure 4-58 Standard gear circular tooth thickness

Example 4-9 A spur gear has 12 teeth (N), a diametral pitch (P) of 32, and PGT-2 tooth design. Find the minimum circular tooth thickness (t), outside diameter (DO), and the root diameter (DR). 1) Minimum circular tooth thickness (t), see Figure 4-54 t =

2.4793 − (0.0426 × N ) 2.4793 − (0.0426 × 12) = = 0.0615 in P 32

2) Outside diameter (DO), see Figure 4-54 DO = (1/P) × [(N – 2.0158) + 2.7475 × t × P] DO = (1/32) × [(12–2.0158) + 2.7475 × 0.0615 × 32] = 0.4809 in Check for the maximum possible outside diameter DO (Max.), Equation 4-10

t Increased

Figure 4-59 PGT-1 design, circular tooth thickness increased

304

4 Thermoplastic Gearing Design

DO (Max.) = Inv φ1 =

N × 0.93969262 P × 1.017 × cos φ1

t×P 0.0615 × 32 = = 0.16379 N + 0.0149 12 + 0.0149

φ1 = 41°44′, cos φ1 = 0.74625 DO (Max.) =

12 × 0.93969262 = 0.4643 in 32 × 1.017 × 0.74625

The smaller outside diameter obtained by using both calculations should be the one speciﬁed for the gear, DO = 0.4643 in. 3) Root Diameter (DR), see Figure 4-54 DR = (1/P) × [(N – 7.2758) + 2.7475 × t × P] DR = (1/32) × [(12–7.2758) + 2.7475 × 0.0615 × 32] = 0.3166 in

4.10.8

Mating Spur Gears Tooth Form Comparison

To obtain the best performance of a low tensile strength thermoplastic gear, it is essential that the gears used in power drives be designed so that the teeth are as strong as possible. This can be done by modifying the geometry or the tooth form. A gear tooth is a cantilevered beam. The two mating spur gears should have equal circular tooth thicknesses and strength at their roots, where the root ﬁllet radius is as large as possible and tangent to the ﬂanks of the teeth.

Example 4-10 Two mating spur gears made of the same material, see Figure 4-60, have a diametrical pitch (P) of 32, the pinion has 13 teeth (NP), and the gear has 60 teeth (NG). Determine both circular tooth thicknesses of the gears for equal strength. The standard circular tooth thickness of both gears is 0.0491 in. The pinion teeth are weaker than the gear and are capable of transmitting only 60% of the load. To design equal strength for both gears it is important to arrive at the right circular tooth thickness. Figure 4-61 shows two modiﬁed tooth proﬁles of a 13-tooth pinion and a 60-tooth gear; both mating gears have a diametral pitch of 32. Because the pinion has less than 26 teeth, the PGT-4 tooth design can be used for the pinion. For the 60-tooth gear, the standard circular tooth thickness should be used. tP =

2.6751 − (0.0426 × N ) 2.6751 − (0.0426 ×13) = = 0.066 or 0.064 in P 32

tG =

t P × htG 0.064 × 0.07075 = = 0.048 in 0.09468 htP

305

4.10 Plastic Gearing Technology Designs Standard pinion 13 teeth

Standard gear 60 teeth 0.064 inch (tP)

htP= 3.03/P = 3.03/32 = 0.09468

Standard circular tooth thickness 3.1416 t= π = = 0.0491 inch 2 x 32 2P

0.0491 (tP)

0.0491 (tG)

PGT-4 pinion 13 teeth

Figure 4-60 Standard mating spur pinion and gear

The teeth used on both gears are designed to have balanced strength. By applying the equation found in Figure 4-56, the pinion minimum circular tooth thickness is calculated to be 0.066 in. To achieve equal strength, the pinion circular tooth thickness was increased from the standard of 0.0491 in to 0.064 in for the PGT-4 tooth form and the gear circular tooth thickness was reduced from the standard of 0.0491 in to 0.048 in, by multiplying the pinion circular tooth thickness with the whole depth gear/pinion ratio.

4.10.9

PGT Spur Mating Gears Strength Balance

The following Equations 4-11, 4-12, 4-13, 4-14, 4-15, and 4-16 are used to calculate the circular tooth thickness of a pair of mating spur gears to achieve balanced tooth strength. These equations are valid only for spur gears of PGT-1 tooth form design. The longer whole depth teeth (ht) of the PGT-2, PGT-3, and PGT-4 tooth form designs do not require the coarser pitch gears used in power drives. These equations provide answers for any ratio and all combinations of number of teeth, but for a power drive it is advisable for the pinions to have at least 12 teeth. Pinions with less than 12 teeth will have reduced outside diameters. Equations 4-11 and 4-12 give speciﬁc values for the circular tooth thickness required of both spur pinion and gear. Equations 4-15 and 4-16 are needed to calculate the circular tooth thickness of the gear at the very beginning of the design. These equations then give the circular tooth thickness of the pinion. Spur Pinions and Gears with Less than 35 Teeth, Circular Tooth Thickness tP =

2.3329 − 0.0219 × N P P

(4-11)

tG =

2.3329 − 0.0219 × N G P

(4-12)

Spur Pinion with Less than 35 Teeth, Gear with 35 Teeth or More, Circular Tooth Thicknesses tP =

2.3329 − 0.0219 × N P P

(4-13)

tG =

 N G  2.1922 − 0.0066 × N P + Inv φ2 − 0.01490438  P  N G − 2.0938 

(4-14)

htG= 2.20/P + 0.002 = 2.20/32 + 0.002 = 0.07075

0.048 inch (hG)

Standard gear 60 teeth

Figure 4-61 PGT-4 pinion and standard gear for equal strength

306

4 Thermoplastic Gearing Design Spur Pinion and Gear Both with 35 Teeth or More, Circular Tooth Thickness  N × (N G − 2.0938)   t G 0.0149 − Inv φG  + tP =  P    N P − 2.0938 P    NG  0.0149 − Inv φP  − NP    P

tG =

π 2P

(4-15)

(4-16)

Where: tP = Pinion circular tooth thickness (in) tG = Gear circular tooth thickness (in) NP = Pinion number of teeth NG = Gear number of teeth P = Diametral pitch  0.9396 × N P  φP = cos −1   N P − 2.0938   0.9396 × N G  φG = cos −1   N G − 2.0938 

Example 4-11 A pinion with 14 teeth (NP) drives a gear of 28 teeth (NG) made of the same material. The diametral pitch (P) is 20 and the PGT-1 tooth form is required. Determine the circular tooth thickness of both the pinion and gear for balanced tooth strength. Because both gears have less than 35 teeth, Equations 4-11 and 4-12 are used. tP =

2.3329 − 0.0219 × N P 2.3329 − 0.0219 × 14 = = 0.1013 in P 20

tG =

2.3329 − 0.0219 × N G 2.3329 − 0.0219 × 28 = = 0.086 in P 20

Circular tooth thickness of pinion (tP) = 0.1013 in Circular tooth thickness of gear (tG) = 0.086 in

Example 4-12 A pinion having 10 teeth (NP) drives a gear of 44 teeth (NG), both gears are made of the same material. The diametral pitch (P) is 18 and the PGT-1 tooth design is required. Determine the circular tooth thickness of both the pinion and gear for balanced tooth strength. Because the pinion has less than 35 teeth and the gear more than 35 teeth, Equations 4-13 and 4-14 are used.

307

4.10 Plastic Gearing Technology Designs

tP =

2.3329 − 0.0219 × N P 2.3329 − 0.0219 × 10 = = 0.1174 in P 18

tG =

 N G  2.1922 − 0.0066 × N P + Inv φ2 − 0.01490438   P  N G − 2.0938 

 0.9396 × N G  φG = cos −1  = cos −1  N G − 2.0938 

 0.9396 × 44    = 9.40° 44 − 2.0938 

Inv φG = 0.00194 tG =

44  2.1922 − 0.0066 × 10  + 0.00194 − 0.0149  = 0.0923 in   18  44 − 2.0938

Circular tooth thickness of pinion (tP) = 0.1174 in Circular tooth thickness of gear (tG) = 0.0923 in

Example 4-13 A pinion having 36 teeth (NP) drives a gear of 72 teeth (NG). The diametral pitch (P) is 24 and the PGT-1 tooth design is required. Determine the circular tooth thickness of both the pinion and gear for balanced tooth strength. Since both gears have more than 35 teeth, Equations 4-15 and 4-16 are used.  N × (N G − 2.0938)   t G 0.0149 − Inv φG  tP =  P  N +  − 2.0938 N P   G P  0.0149 − Inv φP  − NP    P  0.9396 × N G  φP = cos −1  = cos −1  N P − 2.0938 

 0.9396 × 36    = 4.60° 36 − 2.0938 

Inv φP = 0.0021182  0.9396 × N G   0.9396 × 72  φG = cos −1  = cos −1  = 14.60°  72 − 2.0938   N G − 2.0938 

Inv φG = 0.005642 Standard circular tooth thickness (tG) =

π 3.1416 = = 0.06545 in 2 P 2 × 24

 36 (72 − 2.0938)   0.065 0.015 − 0.0056   0.015 − 0.00212  tP =   − 36    72 +   − 36 2.0938 24 24   = 0.076 in

Circular tooth thickness of pinion (tP) = 0.076 in Circular tooth thickness of gear (tG) = 0.06545 in

308

4 Thermoplastic Gearing Design

4.10.10 PGT Close Mesh Center Distance Between Spur Gears When two standard spur gears are brought into close mesh, the distance between their centers is half the sum of their standard pitch diameters and is referred to as the standard center distance. But it is rare for a pair of standard gears to be the best combination for a given drive. A pair of involute gears will function at widely varying center distances. The center distance at which a pair of mating spur gears operates should be regarded as a variable that can be manipulated to achieve the best possible drive. The center distance at which a pair of PGT tooth form design gears are in close mesh is determined by using Equation 4-17. C=

(N1 + N 2 ) × 0.46984631 P × cos φ1

(4-17)

Where: C = Close mesh center distance (in) N1 = Number of teeth in the ﬁrst gear N2 = Number of teeth in the second gear P = Diametral pitch t1 = Circular tooth thickness of the ﬁrst gear (in) t2 = Circular tooth thickness of the second gear (in) P (t1 + t 2 ) − π φ1 = Angle whose involute is + 0.01490438 N1 + N 2 Example 4-14 Two mating gears have the basic speciﬁcations given by the table below. Determine the center distance at which they will be in close mesh. Gear parameters

First gear

Second gear

Number of teeth (N) Diametral pitch (P) Tooth form design Standard pitch diameter (DP) Standard circular tooth Thickness (t)

10 32 PGT-1 0.3125 0.0595

40 32 PGT-1 1.250 0.01965

N1 = 10, N2 = 40, P = 32, t1 = 0.0595, t2 = 0.01965.

P (t1 + t 2 ) π + 0.01490438 N1 + N 2 32 (0.0595 + 0.0196) − 3.1416 = + 0.01490438 = 0.00273 10 + 40

Inv φ1 =

φ1 = 5.25°

cos φ1 = 0.995805 (N1 + N 2 ) × 0.46984631 P × cos φ1 (10 + 40) × 0.46984631 = = 0.73722 in 32 × 0.995805

Close Mesh Center Distance C =

309

4.10 Plastic Gearing Technology Designs

4.10.11 Maximum Close Mesh Center Distance The close mesh center distance computed in Example 4-14 using Equation 417 assumes perfection in the gears and any type of bearings employed. It also assumes the gears and the housing in which they are mounted are made of stable materials. The gears’ and bearings’ run-out affects the tolerances of the assembly; in addition the gears and the housing will also change because of the environment, causing the close mesh center distance to vary from high to low tolerances. The minimum operating center distance speciﬁed for the housing must be equal to, or greater than, the maximum close mesh center distance, otherwise the gears may bind. The designer needs to know the tolerances speciﬁed for the gears, bearings and housing, the dimensional changes caused by the materials used for the fabrication of the gears and housing, and the nature of the environment. All these changes of dimensions must be taken into account during the gear design calculations to determine the maximum and minimum operating close mesh center distance. The sum of the tolerances in the teeth is called the tooth-to-tooth composite tolerance. The composite tolerance plus the run-out in the gear is called the total composite tolerance. This tolerance system was developed by the American Gear Manufacturers Association classifying a gear by AGMA Quality Number in accordance with the accuracy required. The system is presented, in detail, in the “Gear Handbook”, American Gear Manufacturers Association Publication 390.03. If speciﬁcations require that the gears in the previous Example 4-14 are to be to the accuracy of AGMA Quality Number Q7, from the “Gear Manual” it can be found that the ﬁrst gear has a maximum total composite tolerance of 0.0027 in and the second gear has a maximum total composite tolerance of 0.0031 in. The close mesh center distance is calculated to be 0.73722 in. If the tolerances in the gears are at the maximum allowed by the tolerances, these tolerances would cause the close mesh center distance to go from a high of 0.74012 in to a low of 0.73432 in. To allow for these maximum permissible tolerances, the operating close mesh center distance must not be less than 0.74012 in. To allow for the tolerances in each of two mating gears, the minimum operating center distance must exceed the calculated close mesh center distance by half the sum of the total composite tolerances of the gears. Because the coefﬁcients of linear thermal expansion of the thermoplastic materials are relatively high and some thermoplastics expand as they pick up moisture, these factors must also be taken into consideration in arriving at the minimum operating center distance. Finally, if the gears are mounted on shafts running in bearings, the maximum allowable run out of the bearings must be taken into account. Assuming that the gears are inspected in a relative humid environment and at a room temperature of 73 °F, the minimum operating center distance (c) must exceed the calculated close mesh center distance and the required increase in center distance (∆c) is expressed by the following Equation 4-18: ∆c =

  α × N1  Tol1 + Tol 2 α × N2 + C (T − 70)  1 + 2 − αH  2  N1 + N 2 N1 + N 2    ∆ DW1 × N1 ∆ DW2 × N 2   + +  + TIR N1 + N 2    N1 + N 2

(4-18)

310

4 Thermoplastic Gearing Design Where: ∆c = Required increase in center distance (in) Tol1 = First gear total composite tolerance (in) Tol2 = Second gear total composite tolerance (in) C = Close mesh center distance (in) T = Maximum application end use temperature (°F) α1 = Coefﬁcient of linear thermal expansion of ﬁrst gear (in/in/°F) α2 = Coefﬁcient of linear thermal expansion of second gear (in/in/°F) αH = Coefﬁcient of linear thermal expansion of housing (in/in/°F) N1 = Number of teeth in ﬁrst gear N2 = Number of teeth in second gear ∆DW1 = Dimensional change due to moisture absorption ﬁrst gear (in/in) ∆DW2 = Dimensional change due to moisture absorption second gear (in/in) TIR = Maximum allowable run-out of bearings (in) The coefﬁcients of linear thermal expansion and the dimensional change caused by the moisture absorption of the thermoplastic polymers are found in the technical manuals issued by the plastic resins suppliers. Unfortunately, information about expansion caused by moisture absorption is not so readily available. In the case of applications where the gears will not be exposed immediately to conditions of high humidity, the expansion of most thermoplastic materials is small and gradual. This expansion can be discounted, being offset by the equally small and gradual mold shrinkage that occurs as molding stresses are relieved. If the gears are to be molded of one of the more hygroscopic thermoplastic resins and if conditions are such that there could be a high percentage of moisture absorption for a long period of time and high temperatures, it is advisable to consult with the plastic supplier’s technical department to determine what allowance, in terms of in/in, should be made for the melt ﬂow and perpendicular to the ﬂow directions.

Example 4-15 The ﬁrst gear in Example 4-14 has 10 teeth and the second gear has 40 teeth. The gears need to have the accuracy required by AGMA Quality Number Q7. The maximum total composite tolerances are 0.0027 in and 0.0031 in, respectively. The ﬁrst gear is to be molded of an unreinforced nylon 6/6 having a coefﬁcient of linear thermal expansion of 5.0 × 10–5 in/in/°F. The second gear is molded of a high viscosity acetal homopolymer having a coefﬁcient of linear thermal expansion of 6.8 × 10–5 in/in/°F. Considering the environment, in which the gears will operate, it has been determined that expansion of the unreinforced nylon 6/6 due to water absorption could be 0.003 in/in and that of the acetal homopolymer 0.0005 in/in. The housing is made of glass reinforced nylon 6/6 having a coefﬁcient of linear thermal expansion of 1.3 × 10–5 in/in/°F. The gear shafts run in bearings concentric (TIR) to 0.0005 in. The maximum temperature to which the gears will be subjected is 150 °F. Determine the minimum operating center distance to be speciﬁed for the housing design. It has already been determined in Example 4-14, that the gears will be in close mesh at a center distance of 0.73722 in. The calculation employed to arrive at this center distance allowed for no tolerances in the gears, assuming they were conditioned at 50% relative humidity and at a temperature of 73 °F.

311

4.10 Plastic Gearing Technology Designs Increase in center distance ∆c: ∆c =

  α × N1  α × N2 Tol1 + Tol 2 + C (T − 70)  1 + 2 − αH  + + 2 N N N N  1  2 1 2   ∆ DW1 × N1 ∆ DW2 × N 2   + +  + TIR N1 + N 2    N1 + N 2

Tol1 = 0.0027 in Tol2 = 0.0031 in C = 0.73722 in T = 150 °F α1 = 5.0 · 10–5 = 0.00005 (in/in/°F) α2 = 6.8 · 10–5 = 0.000045 (in/in/°F) α H = 1.3 · 10–5 = 0.00001 (in/in/°F) N1 = 15 N2 = 60 ∆DW1 = 0.003 in/in ∆DW2 = 0.0005 in/in TIR = 0.0005 in ∆c =

0.0027 + 0.0031 2   0.00005 × 10 0.000068 × 40  + 0.737 (150 − 73)  + − 0.00001   + + 40 10 10 40   0.003 × 10 0.0005 × 40   + +  + 0.0005  40 + 10 10 + 40  

= 0.0029 + 0.00325 + 0.0005 = 0.00665 in Minimum center distance (C) = 0.73722 + 0.00665 = 0.74387 in The speciﬁed circular tooth thickness of the mating gears in Example 4-14 determined the close mesh center distance. When the minimum operating center distance is ﬁxed by requirements of the gear’s mechanism, it becomes necessary to establish the circular tooth thicknesses such that the close mesh center distance is less than the minimum operating decrease in the center distance by c. ∆c is again calculated using Equation 4-18, but the value given for C is now the minimum operating center distance. This will introduce a tolerance, because C should have the value of the unknown close mesh center distance, but the tolerance is so small it can be considered negligible. ∆c is then subtracted from the minimum operating center distance. The answer is the close mesh center distance used in Equation 4-18 to get the sum of the circular tooth thicknesses for a given close mesh center distance. t1 + t 2 =

(N1 + N 2 )(inv φ1 − 0.0149043) + π P

Where: t1 = Circular tooth thickness of ﬁrst gear (in) t2 = Circular tooth thickness of second gear (in) N1 = Number of teeth in ﬁrst gear

(4-19)

312

4 Thermoplastic Gearing Design

N2 = Number of teeth in second gear P = Diametral pitch C = Close mesh center distance (in)  (N + N 2 ) × 0.46984631  φ1 = cos −1  1  P×C  Example 4-16 The gears in Examples 4-14 and 4-15 are required to operate at a ﬁxed minimum operating center distance of 0.7315 in. Determine the sum of the circular tooth thicknesses of each gear. Substituting 0.7315 in for 0.73722 in as the value for close mesh center distance (C), ∆c =

0.0027 + 0.0031 2   0.00005 × 10 0.000068 × 40  + 0.7315 (150 − 73)  + − 0.00001  40 + 10  10 + 40   0.003 × 10 0.0005 × 40   + +  + 0.0005  40 + 10 10 + 40  

= 0.0029 + 0.003231 + 0.00005 = 0.00295 in Close mesh center distance (C) = 0.7315–0.00295 = 0.73445 in The circular tooth thicknesses must be such that, if the gears and bearings were perfect and the gears and housing made of stable materials, the gears would be in close mesh at 0.73445 in. The sum of circular tooth thicknesses for close mesh center distance is calculated by employing Equation 4-19. N1 = 10, N2 = 40, P = 32, C = 0.73445  (N + N 2 ) × 0.46984631  φ1 = cos −1  1  P×C   (10 + 40) × 0.46984631  = cos −1  = cos −1 (0.99957) = 1.65°  × 32 0.73445   Inv φ1 = 0.00051 (N1 + N 2 ) (inv φ1 − 0.0149043) + π P (10 + 40) (0.00051 − 0.0149043) + 3.1416 = = 0.0756 in 32

t1 + t 2 =

From Table 4-8, a gear having 10 teeth, a diametral pitch of 32, and the PGT-1 tooth form design is required to have a minimum circular tooth thickness of 1.9069 / 32 = 0.0595 in. Allowing for a discrepancy of +0.0010 / –0.0000 inches, the maximum becomes 0.0605 in, the maximum tooth thickness of the second gear is 0.0756–0.0605 = 0.0151 in. The standard circular tooth thickness of a gear having a diametral pitch of 32 is 0.0491 in. The reduction to 0.0151 in, although large, is acceptable, as the teeth will have no undercut.

4.10 Plastic Gearing Technology Designs

Example 4-17 The ﬁrst gear has 15 teeth (N1), the second gear has 18 teeth (N2), a diametral pitch (P) of 32, and the minimum operating center distance is ﬁxed at 0.5156 in. Determine the sum of circular tooth thicknesses of each gear. The total composite tolerance of the second gear is now 0.0027 in ∆c =

0.0027 + 0.0027 2   0.00005 × 15 0.000068 × 18  + 0.5156 (150 − 73)  + − 0.000013    + + 18 15 15 18   0.003 × 15 0.0005 × 18   + +  + 0.0005  18 + 15 15 + 18  

= 0.0027 + 0.00269 + 0.0005 = 0.00589 in Close mesh center distance (C) = 0.5156–0.00589 = 0.5097 in t1 + t 2 =

(15 + 18)(inv φ1 − 0.0149043) + 3.1416 32

 (15 + 18) × 0.46984631  −1 φ1 = cos −1   = cos (0.9508026) = 18.047° 32 × 0.5097  Inv φ1 = 0.01084731 t1 + t 2 =

(15 + 18) (0.01084731 − 0.0149043) + 3.1416 = 0.0940 in 32

According to Table 4-8, the circular tooth thickness of the ﬁrst gear with 15 teeth is 1.6939 / 32 = 0.0529 in. The circular tooth thickness of the second gear is 0.0940–0.0529 = 0.0410 in. But, from the same Table 4-8, a gear having 18 teeth must have a circular tooth thickness of no less than 1.5708 / 32 = 0.04908 in to avoid an undesirable amount of undercut. The circular tooth thickness of the ﬁrst gear could be reduced to (π / 2 P) or 3.1416 / (2 × 32) = 0.049 in and the second gear to 0.0940–0.049 = 0.0450 in. The gears would function, but their performance would not be satisfactory. A better design solution should be implemented, such as the use of a ﬁner diametral pitch for the mating gears as illustrated in Example 4-18.

Example 4-18 Let the diametral pitch (P) of the gears in Example 4-17 be changed from 32 to 32.8. Determine the sum of circular tooth thicknesses. Both gears now have a total composite tolerance of 0.0028 in. ∆c = 0.0028 + 0.00269 + 0.0005 = 0.00599 in

313

314

4 Thermoplastic Gearing Design

Close mesh center distance (C) = 0.5156–0.00599 = 0.5096 in

t1 + t 2 =

(15 + 18)(inv φ1 − 0.0149043) + 3.1416 32.8

 (15 + 18) × 0.46984631  −1 φ1 = cos −1   = cos (0.92743031) = 21.962° 32.8 × 0.5096  Inv φ1 = 0.01994651 t1 + t 2 =

(15 + 18) (0.01994651 − 0.0149043) + 3.1416 = 0.1009 in 32.8

From Table 4-8, the minimum circular tooth thickness of the ﬁrst gear, which has 15 teeth and a diametral pitch of 32.8 is 1.5708 / 32.8 = 0.0478 in, say, 0.0488 in maximum and 0.0468 in minimum. The sum of the circular tooth thicknesses is 0.1009 in; therefore, the circular tooth thickness of the second gear is 0.1009–0.0488 = 0.0521 in. From Table 4-8, the minimum circular tooth thicknesses of a 18-toothed gear and a diametral pitch of 32.8 is 1.5708 / 32.8 = 0.0478 in. The diametral pitch of 32.8 was arrived at by trial and error, using Equation 4-19. This example demonstrates the freedom in designing special PGT gear tooth form designs and manufacturing thermoplastic gears by injection molding. A diametral pitch of 32.8 is required, instead of 32 initially speciﬁed.

There is nothing to be gained economically by specifying a diametral pitch of 32 and much to be lost. Rigid adherence to what is considered the “standard” diametral pitch could cause performance problems of the mating gears. Investigate the best diametral pitch as illustrated to improve the operating performance of the gear system.

4.11

PGT Helical Thermoplastic Gearing

A helical gear differs from a spur gear in that the teeth, instead of being parallel to the axis of the shaft on which the gear is mounted, are formed on a spiral that winds around the axis. Included in the data of a helical gear are three additional requirements that are not speciﬁcations of a spur gear. These are the helix angle, the direction of the helix, either right hand or left hand, and the lead. Mating helical gears on parallel shafts have the same helix angle but opposite hands of the helix. The lead of a helical gear is the distance between a point on the tooth ﬂank and the 360° spiral path that ends in the axis plane. Figure 4-62 shows that the lead (L) is equal to the circumference of a circle containing the point multiplied by the cotangent of the helix angle (ψ).

315

4.11 PGT Helical Thermoplastic Gearing

ψ No rm al

Lead (L) =

pl

π x DP x cot ψ

an e

ψ Axial plane DP

πx

DP

Figure 4-62 Helical gear enveloped view

The helix angle of a helical gear is always speciﬁed as the helix angle at the standard pitch circle. The lead (L) is calculated by using Equation 4-20. L = π × DP × cot ψ

(4-20)

Where: L = Lead π = 3.1416 DP = Standard pitch diameter ψ = Helix angle (degrees) The basic principle of load transfer in a rotating helical gear is that it is done through the gear tooth. Because helical gears have their teeth lined-up in the direction of a helix angle, the load moves through the angled tooth face width. Therefore, the helical gears transfer the load by using two basic directional planes, one in the normal plane and the other in the axial plane. Figure 4-63 shows the helical gear top view, the two cross section planes, and their relationship to one another, together with their ruling equations and commonly used terminology. P = Pn × cos ψ DP =

N Ν = P Pn × cos ψ

Where: ψ = Helix angle P = Diametral pitch Pn = Normal diametral pitch DP = Standard pitch diameter N = Number of teeth pn p = Circular pitch = cos ψ pn = Normal circular pitch tan φn φ = tan −1 = Pressure angle cos ψ φn = Normal pressure angle

Helical gear top view

(4-21) (4-22)

ψ

φn pn

ane l pl a rm no ion t c e ss s Cro

p

φ

Cross section axial plane

Figure 4-63 Helical gear plane views and equations

316

4 Thermoplastic Gearing Design The tooth form of a helical gear is speciﬁed as the tooth form conforming to the basic rack in the normal plane. For all helical gears having helix angles of not more than 23° and operating on parallel shafts, the tooth form speciﬁed is the same as that for spur gears. This is not a requirement, but it permits the use of spur gear lines of actions to generate the teeth in helical gears. While the injection molded thermoplastic gears do not require machined hobs that are a necessity for the metal gears, there was no good reason to depart from what has become an accepted practice in the world of gear engineering. As with spur gears, a pair of mating helical gears must have an adequate contact area ratio. Helical gears have a helical overlap that serves as an additional contact area to transfer higher loads and smooth performance. Helical gear axial pitch (px) is the lead (L) divided by number of teeth (N). px = L / N px =

π Pn × sin ψ

(4-23) (4-24)

Where: px = Axial pitch L = Lead N = Number of teeth Pn = Normal diametral pitch ψ = Helix angle (degrees) A helical gear having a face width equal to its axial pitch has a helical overlap of 1.0. To obtain the beneﬁts from the use of helical gears, the helical overlap should be approximately 2.0. In other words, the helical gear face width should be twice the axial pitch, or as close to that dimension as other design considerations permit. It is this helical overlap, or face width contact area ratio, that makes the helical gears superior to spur gears for drives requiring the maximum torque to be transferred, smooth performance, and a quiet operation. Unlike spur gears, the helical gears cause thrust forces on the bearings carrying the shafts on which the gears are mounted. The amount of thrust force increases with the helix angle and high helix angles require special helical gear tooth forms. The injection molded thermoplastic helical gears of a helix angle should be kept between 13 and 23°. Given a choice, it is recommended that a helix angle of 18° be speciﬁed for the helical gears. No speciﬁc explanation can be offered for this recommendation, but from experience it seems that helical gears having a helix angle of 18° have demonstrated excellent performance in critical applications. Everything discussed previously about spur gears holds true for helical gears. The PGT-1 tooth form design is usually speciﬁed for helical gears, although there may be occasions when the longer tooth of PGT-2 tooth form design can be advantageous. The spur gear equations are modiﬁed to account for the helix angle. They could be used to design spur gears by giving the helix angle a value of zero. Given the normal circular tooth thickness of a helical gear, the outside, root diameters and minimum normal circular tooth thickness are obtained by using the following PGT helical gear equations in the tables below (Courtesy: ABA/PGT).

317

4.11 PGT Helical Thermoplastic Gearing Table 4-13 PGT-1 Tooth Helical Gear Design Equations

Outside diameter

DO =

 1  N − 2.3158  + (2.7475 × t n ) Pn  cos ψ 

Root diameter

DR =

 1  N − 6.9758  + (2.7475 × t n ) Pn  cos ψ 

Minimum normal Circular tooth Thickness

tn =

1 Pn

 N × (1 − cos2 φ  2.3329 −  2.7475 × cos ψ  

Table 4-14 PGT-2 Tooth Helical Gear Design Equations

Outside diameter

DO =

 1  N − 2.0158  + (2.7475 × t n ) Pn  cos ψ 

Root diameter

DR =

 1  N − 7.2758  + (2.7475 × t n ) Pn  cos ψ 

Minimum normal Circular tooth Thickness

tn =

1 Pn

 N × (1 − cos2 φ  2.4793 −  2.7475 × cos ψ  

Where: DO = Outside diameter (in) DR = Root diameter (in) Pn = Normal diametral pitch N = Number of teeth tn = Minimum normal circular tooth thickness (in) ψ = Helix angle (degrees)  0.36397023  φ = tan −1   cos ψ  Example 4-19 A helical gear has 38 teeth (N), a normal diametral pitch (Pn) of 32, a helix angle (ψ) of 18°, the PGT-1 tooth form design, and a normal circular tooth thickness (tn) of 0.0475 in. Determine the outside and root diameters. Pn = 32, N = 38, tn = 0.0475, ψ = 18° DO = =

DR = =

 1  N − 2.3158  + (2.7475 × t n )  Pn  cos ψ 1  38  − 2.3158  + (2.7475 × 0.0475) = 1.3067 in   32 0.95105652  1  N − 6.9758  + (2.7475 × t n )  Pn  cos ψ 1 32

38   − 6.9758  + (2.7475 × 0.0475) = 1.1611 in   0.95105652

318

4 Thermoplastic Gearing Design

Given the normal circular tooth thickness of a helical pinion having a small number of teeth, the maximum outside diameters that will provide an adequate top land is obtained by using the following equation. DO (Max.) =

N × cos φ Pn × cos ψ × 1.017 × cos φ1

Where: DO (Max.) = Maximum outside diameter (in) tn = Normal circular tooth thickness (in) ψ = Helix angle (degrees)  0.36397023  φ = tan −1   cos ψ  t × Pn + Inv φ Inv φ1 = n N

Example 4-20 A helical pinion has 10 teeth (N), a normal diametral pitch (Pn) of 32, a helix angle (ψ) of 18°, and a normal circular tooth thickness (tn) of 0.0666 in. Determine the maximum outside diameter. N = 10, Pn = 32, ψ = 18°, cos ψ = 0.95105652  0.36397023  = tan −1 φ = tan −1   cos ψ 

 0.36397023    = 20.9419° 0.95105652 

cos φ = 0.93394337 Inv φ = 0.01719592 Inv φ1 =

t n × Pn 0.0666 × 32 + Inv φ = + 0.01719592 = 0.23031 N 10

φ1 = 45.8404° cos φ1 = 0.697165 N × cos φ Pn × cos ψ × 1.017 × cos φ1 10 × 0.93394337 = = 0.4328 in 32 × 0.951056 × 1.017 × 0.662745

DO (Max.) =

Find the outside diameter using the other equation: DO = =

 1  N − 2.3158  + (2.7475 × t n )  Pn  cos ψ 1  10  − 2.3158  + (2.7475 × 0.0666) = 0.4391 in   32 0.95105652

The outside diameter of 0.4328 in should be speciﬁed for the helical gear, because it is the smaller outside diameter value obtained by the two equations.

319

4.11 PGT Helical Thermoplastic Gearing

Example 4-21 A helical gear has 16 teeth (N), a normal diametral pitch (Pn) of 24, a helix angle (ψ) of 18°, and the PGT-1 tooth form design. Determine the minimum normal circular tooth thickness required to avoid undercutting of the teeth. Pn = 24, N = 16, ψ = 18°, cos ψ = 0.95105652  0.36397023  = tan −1 φ = tan −1   cos ψ 

 0.36397023    = 20.9419° 0.95105652 

cos φ = 0.93394337 cos2φ = 0.87225022 Normal circular tooth thickness: tn = =

4.11.1

 N × (1 − cos2 φ  2.3329 −  2.7475 × cos ψ  

1 Pn

1  16 × (1 − 0.87225)   2.3329 −  = 0.0646 in 24 2.7475 × 0.951056 

PGT-1 Helical Mating Gears Strength Balance

The number of teeth, the normal diametral pitch, the pressure angle, and the helix angle of a pair of mating PGT-1 helical gears are used to calculate the normal circular tooth thicknesses. The gear material bending stress multiplied by the ratio between the tooth whole depth divided by the normal circular tooth thicknesses should be equal to the balanced strength of both helical gears. The following group of equations should be used, based on the number of teeth obtained by Equation 4-25. Number of teeth (N) =

2.0938 × cos ψ 1 − cos φ

(4-25)

Helical pinion and gear less than 30 teeth, normal circular tooth thickness t nP =

1 Pn

 0.36397023 × N P × (1 − cos φ)  2.3329 −  cos ψ  

(4-26)

t nG =

1 Pn

 0.36397023 × N G × (1 − cos φ)  2.3329 −  cos ψ  

(4-27)

Helical pinion has less than 30 teeth and gear has 30 teeth or more, to calculate the normal circular tooth thicknesses  0.36397023 × N P × (1 − cos φ)  2.3329 −  cos ψ  

t nP =

1 Pn

t nG =

NG Pn

  Pn × B × cos ψ  N − (2.0938 × cos ψ) + Inv φG − Inv φ  G 

(4-26)

(4-28)

320

4 Thermoplastic Gearing Design Helical pinion and gear have 30 teeth or more, to calculate the normal circular tooth thicknesses  N × N G − (2.0938 × cos ψ)   t nG Inv φ − Inv φP  t nP =  P +   Pn  N P − (2.0938 × cos ψ)   N G 

(4-29)

 Inv φ − Inv φP  − NP   Pn 

t nG =

B=

 NG  Pn × B × cos ψ + Inv φG − Inv φ  Pn  N G − (2.0938 × cos ψ) 

 N P × cos φ  Pn × t nP + Inv φ  Pn × cos ψ  N P 

  N P × cos φ φP = cos −1    N P − (2.0938 × cos ψ)    N G × cos φ φG = cos −1    N G − (2.0938 × cos ψ) 

Where: Pn = Normal diametral pitch B = Helical tooth thickness constant N = Number of teeth (determining which equation to use) ψ = Helix angle (degrees) NP = Number of teeth in pinion NG = Number of teeth in gear tnP = Normal circular tooth thickness of pinion (in) tnG = Normal circular tooth thickness of gear (in)  0.36397023  φ = Pressure angle = tan −1   cos ψ  Example 4-22 A helical pinion having 10 teeth (NP) drives the gear having 24 teeth (NG). The normal diametral pitch (Pn) is 24, and the helix angle (ψ) is 18°. The PGT-1 tooth form design is required. Determine the normal circular tooth thicknesses for a balanced tooth’s strength. Pn = 24, NP = 10, NG = 24, ψ = 18°, cos ψ = 0.95105652  0.36397023   0.36397023  φ = tan −1  = tan −1  = 20.9419°  0.951056532   cos ψ  cos φ = 0.93394337 N =

2.0938 × cos ψ 2.0938 × 0.95105652 = = 30.146 or 30 teeth 1 − cos φ 1 − 0.93394337

4.11 PGT Helical Thermoplastic Gearing

Because both helical pinion and gear have less than 30 teeth, Equations 4-26 and 4-27 are selected to calculate the normal circular tooth thicknesses of the pinion and the gear to obtain a balanced strength for both teeth.  0.36397023 × N P × (1 − cos φ)  2.3329 −  cos ψ   1  0.36397023 × 10 × (1 − 0.93394337)  = 2.3329 −  = 0.0866 in 24  0.95105652

t nP =

1 Pn

 0.36397023 × N G × (1 − cos φ)  2.3329 −  cos ψ   1  0.36397023 × 24 × (1 − 0.93394337)  = 2.3329 −  = 0.0719 in 24  0.95105652

t nG =

1 Pn

Normal circular tooth thickness of pinion = 0.0866 in Normal circular tooth thickness of gear = 0.0719 in

Example 4-23 A helical pinion having 16 teeth (NP) drives a gear having 36 teeth (NG). The normal diametral pitch (Pn) is 24, and the helix angle (ψ) is 18°. The PGT-1 tooth form design is required. Determine the normal circular tooth thicknesses for a balanced tooth’s strength. Because the pinion has less than 30 teeth and the gear has more than 30 teeth, select Equations 4-26 and 4-28. Pn = 24, NP = 16, NG = 36, cos φ = 0.93394, cos ψ = 0.951056, φ = 20.9419°, inv φ = 0.01719592  0.36397023 × N P × (1 − cos φ)  2.3329 −  cos ψ   1  0.36397023 × 16 × (1 − 0.93394337)  = 2.3329 −  = 0.0847 in 24  0.95105652

t nP =

1 Pn

  N G × cos φ φG = cos −1    N G − (2.0938 × cos ψ)  36 × 0.93394337   −1 = cos −1   = cos (0.988628) = 8.75° − × 36 (2.0938 0.95105652)   Inv φG = 0.000099842 B= =

 N P × cos φ  Pn × t nP + Inv φ   Pn × cos ψ  N P  16 × 0.93394337  24 × 0.0847  + 0.01719592  = 0.0901137   24 × 0.95105652  16

321

322

4 Thermoplastic Gearing Design

  Pn × B × cos ψ + inv φG − inv φ  − × N ψ (2.0938 cos )  G  36  24 × 0.0901137 × 0.95105652  = + 0.00009984 − 0.01719592 24  36 − (2.0938 × 0.95105652)  = 0.0694 in

t nG =

NG Pn

Normal circular tooth thickness of pinion = 0.0847 in Normal circular tooth thickness of gear = 0.0694 in

4.11.2

PGT-1 Helical Mating Gears Center Distance

Given the number of teeth for the mating helical pinion and gear, the diametral pitch, the helix angle, and the PGT-1 tooth form design, the center distance at which the gears are in close mesh is determined by Equation 4-30. Center distance = C =

(N P + N G ) × cos φ 2 × Pn × cos ψ × cos φP

(4-30)

Where: N = Number of teeth, determined by Equation 4-25 ψ = Helix angle (degrees) NP = Number of teeth in pinion NG = Number of teeth in gear tnP = Normal circular tooth thickness of pinion (in) tnG = Normal circular tooth thickness of gear (in)  0.36397023  φ = Pressure angle = tan −1   cos ψ  P (t + t nG ) − π φP = Angle whose involute is n nP + Inv φ NP + NG Example 4-24 Two mating PGT-1 helical gears have the basic speciﬁcations given below. Find the center distance at which they will be in close mesh. Speciﬁcations

Pinion

Gear

Number of teeth (N)

14

38

Normal diametral pitch (Pn)

32

32

Helix angle (ψ)

18°

18°

Tooth form design

PGT-1

PGT-1

Normal circular tooth thickness on standard pitch circle (tn)

0.0618

0.0372

NP = 14, NG = 36, Pn = 32, tnP = 0.0618, tnG = 0.0372, ψ = 18°, cos = 0.95105652

 0.36397023  = tan −1 φ = tan −1   cos ψ 

 0.36397023    = 20.9419° 0.95105652 

323

4.12 PGT Spur and Helical Gears Horsepower Rating cos φ = 0.93394337 Inv φ = 0.01719592 Inv φP = =

Pn (t nP + t nG ) − π + Inv φ NP + NG 32 (0.0618 + 0.0372) − 3.1416 + 0.01719592 = 0.0177 14 + 38

φP = 21.166 cos φP = 0.932538 (N P + N G ) cos φ (14 + 38) × 0.93394337 = 2 × Pn × cos ψ × cos φP 2 × 32 × 0.95105652 × 0.932538 = 0.8556 in

C=

Close mesh center distance = 0.8556 in

4.12

PGT Spur and Helical Gears Horsepower Rating

In this section, a more sophisticated equation will be formulated speciﬁcally for use in estimating the load carrying capacity of injection molded thermoplastic spur and helical gears, designed conforming to the Plastic Gearing Technology Tooth System. Because of the number of various thermoplastic materials developed for gear applications, it is suggested that the technical representative of the plastic supplier under consideration be consulted. It is desirable to obtain the most recent information on their materials’ load bearing strength capacity, coefﬁcient of friction, wear, mold shrinkage, and the injection molding characteristics of the thermoplastics. The fundamental horsepower Equation 4-31 is valid for spur and helical injection molded thermoplastic gears. HP =

DP × F × n × J × σ × K T × K L 126,000 × P × SF × K S

Where: HP = Horsepower DP = Operating pitch diameter of gear dP = Operating pitch diameter of pinion NG = Number of teeth of gear NP = Number of teeth of pinion F = Effective face width (in) n = Speed (rpm) J = PGT gear system geometry factor (Table 4-15) σ = Tensile stress of the thermoplastic material (psi) KT = Temperature factor T = Maximum application temperature of the gears (°F) KL = Life factor P = Diametral pitch SF = Service factor (Table 4-16)

(4-31)

324

4 Thermoplastic Gearing Design KS = Safety factor Number of Teeth of Gear (N G ) R = Ratio = Number of Teeth of Pinion (N P ) dP =

2×C R +1

(4-32)

DP =

2×C ×R R +1

(4-33)

4.12.1

PGT Gear Horsepower Equation Basic Parameters

Operating Pitch Diameter (DP , dP) The operating pitch diameter of a pinion and gear cannot be determined until it is meshing with its mate at a deﬁnite operating center distance. Knowing the operating center distance and the numbers of teeth in two mating gears, the operating pitch diameters are obtained by using Equations 4-32 and 4-33. The gear having the lesser number of teeth is referred to as the pinion and its mate as the gear. Effective face width (F) F = Effective face width that is in contact with the mating gear. Speed (n) n = Number of revolutions turned in one minute. Table 4-15 PGT Gear System Geometry Factor (J) (Courtesy: ABA/PGT)

Tooth design

PGT-1

PGT-2

Geometry factor (J)

0.75

0.65

PGT gear system geometry factor (J) The geometry factors (J) for PGT-1 and PGT-2 tooth form designs are given in Table 4-15. Tensile stress of the thermoplastic (σ) The tensile stress of the injection molding thermoplastic resin may be obtained from the resin supplier’s published properties information, either directly from the technical representative, or from the marketing technical organization of the speciﬁc injection molding thermoplastic chosen for the gear application. Temperature factor (KT ) The temperature factor allows for the decrease in the tensile stress of a thermoplastic material with an increase in temperature. KT = 1.0 – [(T – 73) × 0.003] Life factor (KL) The life factor adjusts the horsepower rating for the number of cycles required of the gear before it fails. KL = 1.0 – (log M / 5) Where: M = HP rating failure requirements =

n × Hours to Failure × 60 1,000,000

325

4.12 PGT Spur and Helical Gears Horsepower Rating Normal diametral pitch (P) P = Diametral pitch of a spur or helical gears. PGT gear system service factor (SF) The service factor takes into account the nature of the load on the mechanism driven by the gears. The service factor value is obtained from the table below. Table 4-16 PGT Gear System Service Factor (SF) (Courtesy: ABA/PGT)

Type of load

Occasional 1/2 hour/day

Intermittent 3 hours/day

8–10 hours/day

24 hours/ day

Steady Light shock Medium shock Heavy Shock

0.50 0.80 1.00 1.25

0.80 1.00 1.25 1.50

1.00 1.25 1.50 1.75

1.25 1.50 1.75 2.00

Safety factor (KS) The value given to the safety factor ranges from 1.0 to 3.0, sometimes even higher than that. As long as the gears have an adequate contact ratio and a high percentage of recess action, they are designed as speciﬁed for a power drive and have teeth of balanced strength, they have some degree of lubrication and the designer has complete conﬁdence in the data relating to the properties of the injection molding thermoplastic materials to be used for the gears, then there is no reason why the safety factor should be greater than 1.0. If a high degree of reliability is desired, the safety factor could be increased to 1.25 or, possibly, 1.50. On the other hand, if the designer has found it necessary to make compromises because of circumstances beyond his or her control, then the extent to which the gears have been weakened must be assessed and the safety factor increased accordingly. Prototype gears for testing can be produced from a single cavity MUD universal frame tool at a reasonable price. No amount of theorizing can substitute for testing prototype gears in the mechanism they will be required to drive and under the conditions to which they will be subjected in the ﬁeld. Example 4-25 A pair of helical gears is designed for a speed-reducing drive in a domestic appliance. The pinion and gear have 14 and 38 teeth, respectively and a normal diametral pitch of 32. The PGT-1 tooth design is speciﬁed and the teeth are designed to have balanced strength. The operating center distance is 0.8556 in, with a pinion and gear face width of 0.800 inches. The pinion rotates at a speed of 1,800 rpm. The drive operates intermittently for one or two hours per day and is subject to light shock loading. A life of 3,000 hours is required. The helical gears are grease-lubricated. The pinion is injection molded of a thermoplastic material having a tensile stress of 9,500 psi and the gear is made of a thermoplastic material having a tensile stress of 8,000 psi. Determine the horsepower rating for the helical pinion: HP =

DP × F × n × J × σ × K T × K L 126,000 × P × SF × K S

326

4 Thermoplastic Gearing Design

1) Operating pitch diameter (dP) C = 0.8556 , R = dP =

NG 38 = = 2.714 NP 14

2 × C 2 × 0.8556 = = 0.46 R +1 2.714 + 1

2) Effective face width (F) F = 0.800 in 3) Speed (n) n = 1,500 rpm 4) PGT gear system geometry factor (J) J = 0.75 (from Table 4-15) 5) Tensile stress of the thermoplastic (σ) σ = 9,500 psi 6) Temperature factor (KT) T = 175 °F K T = 1.0 − [(T − 73) × 0.003] = 1.0 − [(175 − 73) × 0.003] = 0.694 7) Life factor (KL) K L = 1.0 − (log M /5) M =

n × Hours to Failure × 60 1,000,000

n = 1,800 rpm Hours to failure = 3,000 M =

1,800 × 3,000 × 60 = 324 1,000,000

K L = 1.0 − log (324 /5) = 0.484 8) Normal diametral pitch (P) P = 32 9) PGT gear system service factor (SF) SF = 1.0 (from Table 4-16)

4.12 PGT Spur and Helical Gears Horsepower Rating

10) The factor of safety (KS) The pinion and gear are designed for a power drive and have balanced tooth strength. The contact ratio is 2.714 and there is 75% recess action. A degree of lubrication is provided. A safety factor of 1.2 would be adequate for good reliability. KS = 1.2 HP (Gear) = =

DP × F × n × J × σ × K T × K L 126,000 ⋅ P ⋅ SF ⋅ K S 0.46 × 0.80 × 1,800 × 0.75 × 9,500 × 0.694 × 0.484 = 0.327 126,000 × 32 × 1.0 × 1.2

Example 4-26 Determine the horsepower rating for the gear in Example 4-25. HP (Gear) =

DP × F × n × J × σ × K T × K L 126,000 × P × SF × K S

1) Operating pitch diameter (DP) DP =

2 × C × R 2 × 0.8556 × 2.714 = = 1.120 R +1 2.714 + 1

2) Effective face width (F) F = 0.800 in 3) Speed (n) n = 1,800/2.714 = 663.22 rpm 4) PGT gear system geometry factor (J) J = 0.75 (from Table 4-15) 5) Tensile stress of the thermoplastic (σ) σ = 8,000 psi 6) Temperature factor (KT) T = 175 °F K T = 1.0 − [(T − 73) × 0.003] = 1.0 − [(175 − 73) × 0.003] = 0.694 7) Life factor (KL) K L = 1.0 − (log M /5)

327

328

4 Thermoplastic Gearing Design

M =

n × Hours to Failure × 60 1,000,000

n = 663.22 rpm Hours to failure = 3,000 M =

663.22 × 3,000 × 60 = 119.38 1,000,000

K L = 1.0 − log (119.38/5) = 0.545 8) Normal diametral pitch (P) P = 32 9) PGT gear system service factor (SF) SF = 1.0 (from Table 4-16) 10) Safety Factor (KS) KS = 1.2 HP (Gear) =

DP × F × n × J × σ × K T × K L 126,000 × P × SF × K S

1.120 × 0.80 × 663.22 × 0.75 × 8,000 × 0.694 × 0.545 126,000 × 32 × 1.0 × 12 = 0.278

=

The horsepower rating for the helical pinion is 0.327, and it is higher than the rating for the helical gear (0.278). To establish the horsepower rating for the drive, select the lower calculated value of the helical pinion and gear. The calculated horsepower rating of the drive is 0.278.

4.13

PGT Spur and Helical Gear Specifications

The ﬁnal step in designing a gear is the preparation of a drawing which lists the gear documentation. The data must be speciﬁed in a way that there is no possibility of misinterpretation. This might appear to be self-evident, but all too often, ambiguity in writing gear speciﬁcations has resulted in costly and timeconsuming changes to be performed to the expensive gear mold. There are three types of data speciﬁcations on the gear documentation: • The ﬁrst group consists of data basic to the design of the gear; • The second group consists of data used in the injection molding and inspection • The third group consists of engineering reference data. Table 4-17 is based on the American Gear Manufacturers Association, Standards Department; this table may be used as a gear designer’s checklist for the various documentation that accompanies a gear drawing.

329

4.13 PGT Spur and Helical Gear Speciﬁcations

X

X

Diametral pitch

X

Normal diametral pitch

X

X

Basic speciﬁcations

4 5

X

Transverse pressure angle

6 X

7

Helix angle

X

8

Hand of helix

X

9

Standard pitch diameter

X

X

10

Tooth form

X

X

11

Addendum

X

12

Whole depth

X

13

Max. calculation of circular thickness on standard pitch circle

X

Max. calculation of normal circular thickness on standard pitch circle

Manufacturing and inspection

3 X

Normal pressure angle

Engineering reference

1 2

Transverse diametral pitch Pressure angle

Item number

Number of teeth

Optional

Gear design documentation

Helical gear

Type of data

Spur gear

Table 4-17 Spur and Helical Gears Designer Check List (Courtesy: American Gear Manufactures Association)

14 X

15

Gear testing radius

X

X

16

A.G.M.A. quality class

X

X

17

Maximum total composite error

X

X

18

Maximum tooth-to-tooth composite error

X

X

19

Testing pressure (oz) Master gear speciﬁcations

X X

X

Measure over two 0.xxx diameter pins (for setup only) Outside diameter (preferably shown on drawing of gear)

21 X

X

20

X

22 23

Maximum root diameter

X

24

Active proﬁle diameter

X

25

Surface roughness of active proﬁle

X

26

Mating gear part number

X

27

Number of teeth in mating gear

X

28

Minimum operating center distance

X

29

330

4 Thermoplastic Gearing Design Tables 4-18 and 4-19 are illustration tables for the data that should appear on the drawing of a PGT-1 tooth design pinion and spur gear. The last two speciﬁcation tables (Tables 4-20 and 4-21) are illustration tables for the data that should appear on the drawing of a PGT-1 tooth design helical pinion and gear. The gear documentation is presented in a format recommended by the American Gear Manufacturers Association. Some of the information is redundant; however, it is provided for the convenience of the personnel in manufacturing and inspection departments. The PGT tooth form design is best speciﬁed by including it on the dimensioned drawing of the basic gear. The basic gear documentation for the PGT system tooth form design is based on 1.0 diametral pitch (Tables 4-8, 4-9, 4-10, and 4-11), for other diametral pitch sizes, the given values should be divided by the required diametral pitch to obtain the circular tooth thickness, the outside diameter, and the root diameter. During the gear inspection, performance requires that the injection molded thermoplastic gear be brought into close mesh with a master gear of known accuracy in a center distance measuring instrument. The testing radius of the gear is the center distance, as measured, less half the pitch circle diameter of the master gear. As the gear is rotated with the master through 360°, the center distance and, as a consequence, the testing radius will vary from a high to a low value (see Figure 4-45). For the gear to be acceptable, the high and low values must be within the maximum and minimum tolerance limits speciﬁed. The speciﬁcations of the master gear are supplied by the gear mold maker; consequently, the master gear should not be purchased at the time the gear drawing is being prepared. In that event, the testing radius speciﬁed can be the value that would apply if the master were a theoretically perfect gear of a known pitch circle diameter, or if the gears were to be rotated in close mesh with a standard gear. Once the master gear is available, it is preferable to change the testing radius to conform to the speciﬁcations supplied and to specify the master gear to be used by a tool number. To determine the testing radius to be speciﬁed in the documentation, the close mesh center distance is calculated by employing the center distance (C) equation. Two calculations are made: one for the maximum and the other for the minimum calculated circular tooth thickness of the gear. To the maximum center distance is added half the total composite tolerance and from the minimum is subtracted half the total composite tolerance. These are the maximum and minimum values of the close mesh center distance that are obtained as the gear is rotated with the master gear using a center distance measuring instrument. To obtain the maximum and minimum values of the testing radius half the pitch circle diameter of the master is subtracted from the maximum and minimum values of the close mesh center distance. If the prototype plastic machined gears prove to be unsatisfactory in testing, gear changes are made by a few adjustments to the settings of the machine used to cut the teeth. But, for injection molding the ﬁrst samples of thermoplastic gear changes are considered only after the prototype or the production mold has been built. If the injection molded thermoplastic gear requires design modiﬁcations on the mold, these unexpected changes are both time-consuming and costly. It is essential, therefore, that the ﬁnal design of an injection molded thermoplastic gear be the result of close study and that the data appearing on the drawing be exact and speciﬁed, so that there is no possibility of data misinterpretation.

331

4.13 PGT Spur and Helical Gear Speciﬁcations Table 4-18 PGT Spur Pinion Specifications Example (Courtesy: Plastics Gearing Technology, Inc.)

Engineering references

Manufacturing and inspection

Basic speciﬁcations

Spur pinion documentation

Values

Number of teeth

15

Diametral pitch

76

Pressure angle

20°

Standard pitch diameter

0.1974

Tooth form

PGT-4

Addendum

0.0178

Whole depth

0.0399

Calculated normal circular tooth Thickness on standard pitch circle

0.0278 Max. 0.0268 Min.

Gear testing radius

0.1093 Max. 0.1059 Min.

AGMA quality number

Q7

Maximum total composite tolerance

0.0021

Max. tooth-to-tooth composite tolerance

0.0015

Master gear speciﬁcations

152 T, 0.0207 CTT

Testing pressure (oz)

5

Diameter of measuring pin

0.025

Measurement over two pins (setup only)

0.2475 Max. 0.2458 Min.

Outside diameter

0.2440 Max. 0.2410 Min.

Maximum root diameter

0.1728

Mating Gear Part Number

Gear

Number of teeth in mating gear

120

Operating center distance

0.8780 Max. 0.8750 Min.

332

4 Thermoplastic Gearing Design Table 4-19 PGT Spur Gear Specifications Example (Courtesy: Plastics Gearing Technology, Inc.)

Engineering references

Manufacturing and inspection

Basic Speciﬁcations

Spur gear documentation

Values

Number of teeth

120

Diametral pitch

76

Pressure angle

20°

Standard pitch diameter

1.5789

Tooth form

PGT-4

Addendum

0.0178

Whole depth

0.0399

Calculated normal circular tooth Thickness on standard pitch circle

0.0017 Max. 0.0007 Min.

Gear testing radius

0.7630 Max. 0.7590 Min.

AGMA quality number

Q7

Maximum total composite tolerance

0.0025

Max. tooth-to-tooth composite tolerance

0.0012

Master gear speciﬁcations

152 T, 0.0207 CTT

Testing pressure (oz)

5

Diameter of measuring pin

0.028

Measurement over two pins (setup only)

1.5777 Max. 1.5744 Min.

Outside diameter

1.5630 Max. 1.5580 Min.

Maximum root diameter

1.4826

Mating gear part number

Pinion

Number of teeth in mating gear

15

Operating center distance

0.8780 Max. 0.8750 Min.

333

4.13 PGT Spur and Helical Gear Speciﬁcations Table 4-20 PGT Helical Pinion Specifications Example (Courtesy: Plastics Gearing Technology, Inc.)

Engineering references

Manufacturing and inspection

Basic speciﬁcations

Helical pinion documentation

Values

Number of teeth

15

Normal diametral pitch

16

Normal pressure angle

20°

Helix angle

18°

Hand of helix angle

R.H.

Standard pitch diameter

0.9857

Tooth form

PGT-1

Addendum

0.0625

Whole depth

0.1456

Calculated normal circular tooth Thickness on standard pitch circle

0.1221 Max. 0.1201 Min.

Gear testing radius

0.5259 Max. 0.5191 Min.

AGMA quality number

Q7

Maximum total composite tolerance

0.0043

Max. tooth-to-tooth composite tolerance

0.0021

Master gear speciﬁcations

N = 30, tn = 0.0982

Testing pressure (oz)

15

Diameter of measuring pin

0.110

Measurement over two pins (setup only)

1.1835 Max. 1.1798 Min.

Lead

9.5310

Outside diameter

1.1770 Max. 1.1710 Min.

Maximum root diameter

0.8853

Mating gear part number

Gear

Number of teeth in mating gear

45

Operating center distance

2.0160 Max. 2.0110 Min.

334

4 Thermoplastic Gearing Design Table 4-21 PGT Helical Gear Specifications Example (Courtesy: Plastics Gearing Technology, Inc.)

Engineering references

Manufacturing and inspection

Basic speciﬁcations

Helical gear documentation

Values

Number of teeth

45

Normal diametral pitch

16

Normal pressure angle

20°

Helix angle

18°

Hand of helix angle

L.H.

Standard pitch diameter

2.9572

Tooth form

PGT-1

Addendum

0.0625

Whole depth

0.1456

Calculated normal circular tooth Thickness on standard pitch circle

0.0961 Max. 0.0941 Min.

Gear testing radius

1.4781 Max. 1.4707 Min.

AGMA quality number

Q7

Maximum total composite tolerance

0.0047

Max. tooth-to-tooth composite tolerance

0.0017

Master gear speciﬁcations

N = 30, tn = 0.0982

Testing pressure (oz)

15

Diameter of measuring pin

0.110

Measurement over two pins (setup only)

3.1095 Max. 3.1044 Min.

Lead

28.5930

Outside diameter

3.0770 Max. 3.0710 Min.

Maximum root diameter

2.7853

Mating gear part number

Pinion

Number of teeth in mating gear

15

Operating center distance

2.0160 Max. 2.0110 Min.

335

5

Plastic Journal Bearing Design

5.1

Introduction

A journal bearing is a simple device for providing support and radial positioning while permitting rotation of a shaft. It is the oldest bearing device known to man, yet it is as modern as the latest materials and technologies available today. A great variety of plastic materials can be used for journal bearings. These materials include a number of composites, such as acetal homopolymer, acetal with Teﬂon® powder, chopped ﬁber and fabric, unreinforced and internally lubricated or with molydisulﬁde nylon 6/6, nylon 66 with Teﬂon® powder, chapped ﬁber, fabric, polysulfone with glass ﬁber and Teﬂon®, polycarbonate with glass ﬁber and Teﬂon®, PTFE, sintered metals, wood, rubber, bronze, among others. Bearing types range from simple thermoplastic sleeve bearings to some exceedingly complex gas-lubricated high-speed rotor bearings. The journal bearing is inherently quiet in operation, because it has no moving parts. With proper selection, installation and lubrication, it does not fail suddenly. Wear, if any, is gradual and replacement of worn bearings can be scheduled when equipment is normally idle. These bearings are uniquely suited to conditions involving oscillating or longitudinal movement. Journal bearings are subject to their environments and must be properly selected and used with proper control of shaft tolerances, housings, mounting or installation procedures, lubrication, and so forth. Plastic journal bearings can be divided into two general groups, based on the technical requirements, which can be entirely different: • Injection molded, compression molded, and machined sleeve bearings that are pressed or mounted into metal or thermoplastic housings. The main requirements in these cases are, ﬁrst of all, excellent bearings and wear properties. Impact resistance and tensile strength as well as dimensional stability are considerations of lesser importance. • Composite journal bearings are exposed to wear, as a part of a larger unit, e.g., a gear box, spherical joint bearings, and others. In these applications, all mechanical and physical properties of the composite journal bearings are extremely important.

5.2

Materials Used for Journal Bearings

Numerous materials are employed to meet the needs of special applications. Although babbitt and bronze are the materials used in many applications, the technology is rapidly changing, particularly with the advent of new thermoplastics, sintered metals, and plastic composites. The selection of the most desirable material for a particular application can become a complex decision; however, there are certain applications in which a particular type of material excels.

336

5 Plastic Journal Bearing Design

5.2.1

Babbitt Journal Bearings

Babbitt journal metal bearings are universally accepted as providing reasonable capacity and dependable service, often under adverse conditions. Babbit is a relatively soft bearing material, therefore it minimizes the danger of scoring or damage of large expensive shafts or rotors. Babbitt can often be repaired quickly on the spot by rescraping, pouring of new metal, and so forth. The actual bearing operating temperature must not exceed 200 °F. Babbitt journal metal bearings are used for applications involving light to moderate loadings; bronze is required for heavier loadings or higher temperatures.

5.2.2

Bronze Journal Bearings

Bronze journal bearings are suitable for heavier loads with capacities ranging from 75 to 200% higher than Babbitt, depending on the speciﬁc ranges of loads and speeds. Bronze bearings withstand higher shock loads than Babbitt bearings, permit slightly higher speed of operation, and are used at operating temperatures up to 300 °F. Bronze is a harder material than babbitt and therefore has a greater tendency to score or damage journals if there are malfunctions. Field repair of a bronze bearing generally requires replacement of the bronze bearing bushing.

5.2.3

Sintered Porous Metal Journal Bearings

An increasingly popular bearing material is sintered porous metal, which is usually a bronze alloy, iron alloy, and stainless steel. Sintered porous metals provide an excellent design for boundary lubrication conditions. In the fabrication of this type of bearing, the powdered metal alloy is ﬁrst pressed in dies to a controlled density and then sintered at high temperature in a reducing atmosphere. Any subsequent machining of the part must be controlled with proper tooling and machining techniques so that the open pore structure is maintained. The sintered porous metal journal bearing material functions somewhat like a sponge, wherein the lubricant is retained in the voids. These voids normally make up about 20% of the volume of the sintered porous metal journal bearing. Oils used for impregnation should be non-gumming types, resistant to oxidation. In operation, the development of heat or pressure between the shaft and the plain bearing initiates a capillary action, bringing a measured quantity of oil to the surface for lubrication. This type of bearing will function until the supply of lubricant contained within the sintered porous metal journal bearing is exhausted. An additional supply of lubricant is provided in a wicking type material surrounding the sintered porous metal journal bearing, with provision for additional replenishment through the oil supply reservoir.

5.2.4

Plugged Bronze Journal Bearings

The plugged bronze journal bearing is a hybrid combining the characteristics of cast bronze and carbon-graphite. The basic bronze bearing is provided with a series of holes or grooves that are ﬁlled with plugs of a solid lubricant compound of carbon-graphite, metallic oxides, waxes, or organic salts. These individual solid lubricant reservoirs are uniformly spaced around the plugged bronze journal bearing to provide a continuous distribution of a uniform ﬁlm of low friction solid lubricant over the entire contact surface. This self lubricating plugged

5.2 Materials Used for Journal Bearings bronze journal bearing is generally suitable for temperatures up to 500 °F and is less susceptible to shock than the pure carbon-graphite bearing.

5.2.5

Carbon-Graphite Journal Bearings

Carbon-graphite is utilized for journal bearing applications, in particular at temperatures up to 700 °F or even higher, when the use of conventional bearings and lubricants is next to impossible. This journal bearing utilizes a carbongraphite bushing (liner). The carbon-graphite bearing is entirely self-lubricating; in essence, the bearing itself is the lubricant. In service, this solid lubricant is gradually consumed with shaft and bearing clearance gradually increasing as the solid lubricant is used. These types of carbon-graphite journal bearings have been applied in ovens, dryers, furnaces, and so forth, where loads are light and speeds low. They are often used in environmental applications because of the basic inertness of the carbon-graphite material. As an extreme example, the carbon-graphite journal bearings have been used successfully in pumps transporting molten salt at 1200 °F.

5.2.6

Cast-iron Journal Bearings

Cast-iron journal bearings are low in cost and very suitable for many moving shafts and oscillating or reciprocating arms supporting light loads. The lubricating characteristics of cast iron are attributed to the free graphite ﬂakes present in the alloy. With the use of cast-iron journal bearings, higher shaft clearance is usually provided. Any large wear particles or extraneous debris will not jam the clearance space and seize the bearing. These bearings have been applied in temperatures as high as 1000 °F under light loads and at slow speed intermittent operation.

5.2.7

Wooden Journal Bearings

Wooden journal bearings are still used in a remarkably large number of applications. Light duty machinery frequently employs small oil impregnated hard maple bearings. For heavy duty operations in water or other liquids, lignum vitae wooden bearings are employed. Native only to the Caribbean, lignum vitae is the hardest of all woods. Its speciﬁc gravity of 1.25 and closely interwoven grain structures give the wood a high resistance to wear, compression, and splitting. This material has a resin content of 30% by volume, providing a remarkable self lubricating quality. It is unaffected by salt water, mild acids, alkalis, oils, bleaching compounds, and liquid phosphorous and is often used in chemical processing and food processing industries. The operating temperatures for wooden journal bearing should not exceed 150 °F.

5.2.8

Rubber Journal Bearings

Rubber has been a surprisingly effective journal bearing material when immersed in water and other liquids. It is especially useful for light loads accompanied by abrasive conditions, such as sand and grit. For example, water lubricated rubber journal bearings for propeller shafts of ships operating in shallow sandy water have demonstrated better results concerning wear than bronze or wooden bearings. The rubber must never run dry, even when starting and the shaft must be particularly smooth. The inner surface of the rubber journal bearing is ﬂuted to provide passage for the coolant lubricant and to improve distribution.

337

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5 Plastic Journal Bearing Design

5.2.9

Self-Lubricated Thermoplastic Journal Bearings

Self-lubricated thermoplastic bearings offer a growing class of designs, having a low coefﬁcient of friction, reduced wear rates, resistance to impact and vibration, elimination of lubrication, and cleanliness. The different plastic bearing families are: The Teﬂon® fabric composite bearings with metal support backing, Vespel® bearings, and the thermoplastic bearings reinforced with Kevlar® chopped ﬁbers. The compounded thermoplastic resins commonly used for journal bearing applications are ﬂuorocarbons, nylon 6/6, acetal homopolymer, high density polyethylene, polypropylene, polysulfone, reinforced polycarbonate, and phenolic. These thermoplastic composites are engineered for self lubricated bearings and have different properties such as friction, strength, heat resistance, chemical resistance, wear, cost, and so forth. These materials are commonly used in journal bearing applications involving corrosives, abrasives, and lubrication problems. Requirements of reliability, serviceability, and cost in electro-mechanical data processing machines make optimum material selection, design, and fabrication a necessity. Because of these requirements, the use of self lubricating materials has become increasingly important. For several years, certain types of thermoplastic materials such as nylon 6/6 and acetal homopolymer have been used in self lubricating journal bearing applications because of the following desirable properties: • Low coefﬁcient of friction against steel • Good wear resistance • Quiet operation • Operable without lubrication • Easily fabricated at low cost • Low density • Extended design freedom Most thermoplastic journal bearings are self lubricated, or they are only initially lubricated. Occasionally, thermoplastic journal bearings are used with complete hydrodynamic lubrication. The procedure for designing a fully lubricated thermoplastic journal bearing is more dependent on the properties of the lubricant than upon the properties of the bearing material. This is particularly true for hydrodynamically lubricated metal bearings. The design information presented here is for self lubricated thermoplastic journal bearings. Although the data applies speciﬁcally to journal bearings, it appears that it can also be used conservatively for plane surface or thrust bearings. In recent years, new polymers and a vast variety of composite materials have been introduced, many supposedly with excellent bearing properties. Because their wear surface has such excellent lubricity, Teﬂon® fabric bearings never need lubrication. The lubrication is inherent in the bearing surface so there is no danger of these bearings drying out and thus causing shaft seizure and costly repairs. Various additive compounds such as graphite, molybdenum disulﬁde, aluminum, calcium and other stearates, Teﬂon®, TFE ﬂuorocarbon powder/ﬁbers, and

5.3 Hydrodynamics of Lubrication petroleum hydrocarbons have been used singly and in combination as low coefﬁcient of friction additives for thermoplastics to reduce friction and improve their bearing characteristics.

5.3

Hydrodynamics of Lubrication

The classic hydrodynamic lubrication is the most desired mode of operation for a journal bearing. The principle of the load-carrying lubricant ﬁlm is straight forward, although the mathematical equations become complex. As the shaft revolves within the bearing, the motion of the shaft forces the lubricant to ﬂow in the direction of rotation. In the area of highest loads, the shaft is displaced from the geometric center of the bearing by the applied loads. As the distance between shaft and bearing decreases at the load zone, a wedge-shaped clearance is created, ranging from the high clearance space opposite the load zone to the minimum clearance space at the load zone. The lubricant entrained by the shaft and carried into the diminishing clearance wedge is squeezed severely. The lubricant used with the journal bearing is essentially incompressible; a pressure is developed within the oil ﬁlm that is sufﬁcient to support the applied load. So long as the ﬂuid ﬁlm wedge is of sufﬁcient thickness to prevent metal to plastic contact between shaft and sleeve, the bearing is operating with perfect ﬂuid ﬁlm lubrication. The load carrying ability of the full ﬁlm lubricant wedge is largely dependent on lubricant viscosity and the relative velocity between shaft and bearing. With insufﬁcient velocity and/or viscosity, the lubricant wedge will not develop. Direct contact will occur between the shaft and journal bearing’s surface, with the separating ﬁlm of lubricant being wholly or partially excluded. Where operating speeds are insufﬁcient to produce hydrodynamic lubrication, the hydrostatic principle of lubrication is often employed. With hydrostatic lubrication the necessary pressure to support the applied load is provided from an external source. With the hydrostatic principle, a separating oil ﬁlm can be maintained even with a stationary bearing. Boundary lubrication, thin ﬁlm lubrication, or partial ﬁlm lubrication are synonymous terms applied to a prevalent condition of lubrication. These terms are descriptive of a lubrication regime in which the separating ﬁlm of lubricant exists over only part of the load bearing surface and part of the load is carried by direct contact of the shaft and bearing material. The characteristics of the lubricant and the sliding surfaces become the factors in determining overall friction of the bearing. With such lubrication, different journal bearing materials have different coefﬁcient of friction characteristics under the same operating conditions. Proper selection of shaft and journal bearing materials and lubricant combinations is necessary for successful operation in the boundary lubrication regime. The correct combination results in an absorbed layer of lubricant on the surfaces of both the shaft and the bearing. This layer of lubricant may be of the chemically active type or the molecular attraction polar type ﬁlm. These absorbed layers are extremely thin, often mono-molecular in nature. With overloading, improper material selection, or improper lubricant selection, the absorbed layers will be disrupted, resulting in extreme boundary lubrication. An extreme boundary lubrication condition, if permitted to continue, marks the beginning of the breakdown of the shaft or the bearing surface or both. The three basic conditions of a journal bearing/shaft interface as shown in Figure 5-1 include, ﬂuid ﬁlm, boundary, and extreme boundary lubrication.

339

340

5 Plastic Journal Bearing Design Journal bearings designed for full ﬁlm lubrication often pass through an extreme boundary lubrication phase during periods of starting and stopping. In some cases, such bearings are hydrostatically lubricated during starting and stopping periods.

FLUID FILM Velocity

Fluid film

Adsorbed boundary layer Mating surface completely separated by a fluid film of lubricant. BOUNDARY Velocity

Absorbed boundary layer Mating surface separated by two absorbed boundary

Journal bearings are generally designed for a speciﬁc type of lubricant, i.e., grease, oil, or solid lubricant. For journal bearings designed for oil lubrication, the viscosity of the oil is the most important factor for satisfactory lubrication. The oil viscosity governs the operating temperature of the bearing, the rate of ﬂow of lubricant through the bearing, and the ﬂuid ﬁlm thickness or load carrying capacity of the bearing. The optimum oil viscosity for a given bearing application can be a very difﬁcult value to establish. The optimum depends on journal bearing loading conditions, bearing design, bearing clearances, operating temperature, oil supply methods, and so forth. If oil viscosity is too low, full ﬁlm lubrication will not be attained and a boundary or extreme boundary lubrication regime will prevail. With an oil viscosity that is too high, the bearing will be unable to pump the oil or develop the oil wedge and the result again will be boundary or extreme boundary lubrication. Fortunately, in most applications, the range between these two viscosity levels is broad and the oil viscosity selection is not critical. For most applications of ring oiling bearings at normal ambient temperatures and operation within normal load and speed limits, the oil viscosity should range between 100 and 200 SUS at the estimated operating temperature. When operating conditions are severe and involve variable and somewhat unpredictable loading conditions, EP additive oils are beneﬁcial. These oil types will provide a thin molecular ﬁlm to prevent or minimize metal to plastic contact in boundary lubrication conditions.

layers of lubricant. EXTREME BOUNDARY Velocity

Metal to plastic contact Mating surface in direct metal to plastic contact at various high points.

Figure 5-1 Boundary conditions for lubrication

Many rigid journal bearings are designed for grease lubrication. Because of its apparent high viscosity, grease is retained in the journal bearing longer than oil, primarily because of a reduced side leakage effect. Grease, in many cases, will provide true hydrodynamic full ﬁlm lubrication and because of its high apparent viscosity and its consistency it will function more effectively in the boundary lubrication regime. This protects the journal bearings in start up and shut down operations or in a slow speed operation, where velocities are not sufﬁcient for developing a hydrodynamic ﬁlm. In many instances, grease fortiﬁed with solid lubricant additives has been proven beneﬁcial in extreme boundary lubrication conditions. Solid lubricants can be beneﬁcial as additives to oils or grease and they are also coming into increasing use as pure solid lubricants. Probably the best known types are graphite and molydisulﬁde. With the use of solid lubricants, the ruling principles are entirely different from those of the ﬂuid or grease lubricants. There is no lubricant wedge and the action is primarily a friction phenomenon, with the solid lubricant on the bearing surfaces providing a low coefﬁcient of friction characteristic. Solid lubricants are sometimes used in loose powder form and in this state they must have the ability to form a ﬁlm on the surface to be lubricated. This characteristic is exhibited by MOS2, graphite in moist air, zinc stearate, and other compounds. In the bonded ﬁlm or solid form, dry lubricants have been found effective when properly applied. The performance of the solid lubricant is limited by the wear life of the ﬁlm, which must be replenished as it wears away or erodes to extend life of the system. In the case of the solid graphite bushing, the bearing itself is the lubricant and is gradually consumed during the gradual shear and wear of the structure. The most promising area of application for increased development of solid lubricants is in the ﬁeld of high temperature lubrication and lubrication in vacuum.

5.3 Hydrodynamics of Lubrication Lubricant determinations are usually governed by the type of journal bearing selected, as most journal bearings are designed for a speciﬁc type of lubricant. The method of applying lubricants may vary from simple oil cups, grease cups, or ﬁttings to completely automatic systems. The preferred method relates to the nature of the application and the economics of providing the lubricants to journal bearings at the required rate. When determining whether the journal bearings need to be lubricated, the following points should be considered: • A one-time lubrication, consisting of an initial greasing or use of dry lubricant, generally reduces break in wear and improves overall wear resistance. • Lubrication of the journal bearings can increase the PV limit by reducing the coefﬁcient of friction and helping to remove wear debris. Circulation of the lubricant can further increase the PV limit by cooling the journal bearing. • Lubrication with a chemically compatible ﬂuid to wet the journal bearing surfaces will reduce both friction and wear rates. The amount of reduction increases with increasing ﬂuid ﬁlm thickness, which in turn increases with ﬂuid viscosity and surface velocity and decreases with increasing bearing pressure. Application geometry will also affect the reduction of friction. Even thin ﬁlm lubricants can reduce dry wear rates by a factor of 10 or more. Thick ﬁlms, which cause complete separation of the solid mating surfaces, can theoretically reduce wear to negligible proportions. • The frictional behavior of a journal bearing system using thin ﬁlm lubrication is determined by the properties of the bearing material as well as by the properties of the lubricant. Frictional behavior is determined exclusively by the lubricant properties with thick ﬁlm lubrication. • Unlubricated journal bearings should have surface grooves to carry wear debris out of the interface. In lubricated systems the grooves can help increase the supply of lubricant. The effect of grooving on the journal bearing pressure should be considered. • Because some types of resins do not wet, water is not an effective thin ﬁlm or boundary lubricant for these types of journal bearing materials. In fact, water can adversely affect the wear rate for these journal bearing materials. However, periodic contamination by casual water should not cause any problems. • Purging an unlubricated Vespel® journal bearing with nitrogen gas can reduce wear rates to less than 20% of the corresponding rate in air. In addition, operation in nitrogen can increase the wear transition temperature by at least 100 °F above the value in air. • For applications in dirty environments, sealing or purging should be considered to prevent journal bearing surface contamination. The lubrication makes a big difference in the friction between the two rubbing surfaces and the clearance of the journal bearing. While an acetal homopolymer bearing can be run without any lubrication, this should be done only when absolutely necessary. Even varying degrees of lubrication affect performance and clearances. Continuous lubrication will allow tighter tolerance limits and provide optimum performance. Occasional lubrication will give higher performance than unlubricated surfaces, but clearances must be the same as for no lubrication at all.

341

342

5 Plastic Journal Bearing Design Lubrication, even if applied only at installation, will give improvements over completely unlubricated performance limits. Increasing lubrication decreases the need for larger clearances and raises performance limits.

5.4

Journal Bearings Design for Lubrication

Experience has shown that the service life of a properly dimensioned and designed journal bearing can be increased to several times the normal span, if correct lubrication is provided to the journal bearing. In many cases, it is possible to prevent entry of dirt without excessive cost, thus ensuring effective and longlasting lubrication at the same time. Journal Bearing Lubricated Felt Ring Design Figure 5-2 shows two illustrations of integrally injection molded annular grooves into which felt rings soaked in oil can be placed. The axial annular grooves or pockets are either molded at the edge of the journal bearing housing wall or they could be subsequently machined after molding if needed. Journal Bearing Felt Snapped-On Ring Design Figure 5-3 shows a felt ring which is ﬁtted into an annular recess and secured by a snapped-on ring. The annular recess does not have to be machined after molding the snap ring and the journal bearing. Lubricated Axial Oil Wick with Debris Pocket Figure 5-4 shows how an axial groove for an oil wick on the journal bearing can provide very good results. The second axial groove is a pocket for accumulating the abraded particles once the effect of lubrication has disappeared and the bearing runs dry. This type of journal bearing design can solve many operation problems and it is used in many applications.

Lubricated annular felt ring

The journal bearing should be dimensioned for dry running and in addition should be provided with an initial lubrication. This does not allow a higher load (initial lubrication is only effective over a limited time period), but it does considerably prolong the effective service life. In case higher bearing temperatures are expected, selection of the lubricant must be carried out very carefully in order to avoid any chemical reaction. Lubricated Axial Groove Oil Wick with Seal Lubricated annular felt ring

Figure 5-5 shows the heavy journal bearing of an agricultural machine and the oil chamber with spin-welded end cup and labyrinth seals. The snap-ﬁtted oil plug can be removed to add oil and snap back on the oil chamber. Lubricated Oil Wick Connected to Rod Bearing Figure 5-6 shows a heavy-duty integral connecting rod bearing that was injection molded with an oil chamber and spring-welded end cap. The lubricated wick element provides lubrication to the journal bearing continuously.

Figure 5-2 Journal bearing lubricated felt ring design (Courtesy: Du Pont)

343

5.4 Journal Bearings Design for Lubrication

Snap fitted ring

Lubricated felt ring

Figure 5-3 Journal bearing lubricated felt snapped-on ring design (Courtesy: Du Pont)

Lubrication axial oil wick on groove

Debris pocket

Metal shaft

Journal plastic bearing

Figure 5-4 Lubricated axial oil wick with debris collection pocket (Courtesy: Du Pont)

Journal plastic bearing Oil wick Labyrinth seal Oil plug

Metal shaft

Lubrication groove

Spin weld end cup Oil chamber

Housing

Figure 5-5 Lubricated oil wick on axial groove with labyrinth seal

344

5 Plastic Journal Bearing Design Oil plug

Oil chamber

Spin welded end cap

Oil wick

Rod bearing housing Debris trap grooves

Figure 5-6 Oil wick lubricated integral connecting rod bearing

Journal plastic bearing Lubrication axial and thrust grooves

Metal shaft

Axial/thrust debris trap grooves

Figure 5-7 Axial and thrust grooves for journal bearing lubrication (Courtesy: Du Pont)

Axial and Thrust Grooves for Journal Bearing Lubrication Perpendicular round holes for lubrication/debris trap

Wear problems can inﬂuence journal bearing design decisions. One such inﬂuencing factor is provision of axial and thrust grooves as shown in Figure 5-7. Although this solution does not decrease the rate of wear, it increases considerably the service life of the journal bearing by eliminating the abrasive inﬂuence of the abraded particles from the shaft, journal bearing, and of dust. Journal Bearing with Round Holes for Lubrication/Debris Trap

Figure 5-8 Round holes for journal bearing lubrication/debris trap (Courtesy: Du Pont)

When the journal bearing has a very thin wall thickness, the use of axial grooves is not always possible. A minimum of six perpendicular round holes made as deep as technically feasible through the journal bearing wall thickness is recommended. The holes’ diameter should be approx. 10% of the journal bearing diameter. The perpendicular holes provide the same lubrication and debris trap effects as the grooves (see Figure 5-8). Naturally, the grooves or holes will be molded integrally, even if the bore is subsequently machined.

345

5.5 Journal Bearing Design Principles

5.5

Journal Bearing Design Principles

There are many methods for designing journal bearings. Bearing design may be as simple as merely ﬁnding a standard size, low cost journal bearing to ﬁt both the shaft and the housing without making an analysis. The designer then writes an instruction sheet, in which the user is told to squirt oil into the journal bearing occasionally. If the journal bearing fails, the user did not take care proper of it. On the other hand, the design of an aircraft engine journal bearing might require a very careful design study. Details concerning the material, construction, assembly, and lubrication will be studied very closely. A whole series of relations between the variables might be obtained to plot curves showing the effect of changing each variable. To obtain the maximum reliability, the load life curve would be studied closely and recommendations made for replacing after a speciﬁed number of engine hours.

5.5.1

Journal Bearing Nomenclature and Equations

Figure 5-9 shows a journal bearing assembly, basic equations, and variables required to calculate the journal bearing geometry depending on the end use conditions. The following information will assist in making design decisions:

L

• Journal bearing resins (properties, bearing technical information)

N V

• The number of units to be injection molded • Type of load and its characteristics, whether steady or alternating, predictable or unpredictable • Rotating speed and speed mode characteristics • Journal bearing, shaft and housing design drawings to be available, together with the information concerning making changes • The method of lubrication using force, splash, ring, or wick feed • Type of lubricant used and its operating temperature limits After reviewing the preceding information, the following tentative decisions must be made to complete the journal bearing design: • Journal bearing resin price and manufacturing molding cost • The method of analysis; whether to use the hydrodynamic theory, experimental methods, or some other method • Required assumptions to be used in the analysis, e.g., to avoid side oil leakage or operating maximum temperature • The desired journal bearing life expectancy • If the load is variable or unpredictable, analysis may be necessary in order to obtain a realistic load value on which to base the design • The shaft and journal bearing dimensions, DB, dB, DS, c, t, L • Tolerances used in the journal bearing design analysis.

W

DS dB

c

DB t

Figure 5-9 Journal bearing design assembly and equations

346

5 Plastic Journal Bearing Design V = π × dB × N

(5-1)

W dB × L

(5-2)

P=

PV = Pressure velocity limits (psi-ft/min) DB = Bearing outside diameter (in) dB = Bearing inside diameter (in) DS = Shaft outside diameter (in) c = Clearance bearing/shaft (in) L = Length of bearing (in) t = Bearing wall thickness (in) N = Rotating speed (rpm) V = Surface velocity (ft/min) W = Static load (lb) P = Pressure bearing (psi) The coefﬁcients of friction, power loss, minimum lubricant ﬁlm thickness, oil ﬂow, and temperature rises can be calculated, using any of the procedures illustrated in this chapter or any other procedure selected by the designer as appropriate to the particular problem. When the analysis is completed, the tentative decisions are adjusted according to the results; then the analysis is repeated again, until the independent variables have values that are satisfactory to the designer. For a simple injection molded thermoplastic journal bearing design, a single set of calculations may be sufﬁcient. The designer uses the results of the analysis to alter his preliminary decisions and ﬁxes the design at that point. This design procedure is satisfactory when the designer has experience on which to base these alterations. However, in important applications, the experienced designer will usually carry the design study to much greater detail. Nothing is quite so basic to the design of a thermoplastic journal bearing as the proper running clearance between the shaft outside diameter and the journal bearing inside diameter. These clearances are much larger for thermoplastic journal bearings than for metal bearings. However, such clearances are necessary and beneﬁcial in many ways. Resilience and vibration absorption of the thermoplastic journal bearing aid in overcoming possible disadvantages of large clearances. For thermoplastic journal bearings, generous clearances contribute to improved service life, particularly in the case of unlubricated bearings. Clearances of as much as 0.015 in/in of shaft outside diameter are recommended for thermoplastic journal bearing designs. Basic clearance for a completely unlubricated thermoplastic journal bearing should not be less than 0.005 in/in of shaft outside diameter. The basic clearance for a continuously lubricated thermoplastic journal bearing can be as little as 0.004–0.007 in/in of shaft outside diameter. For thermoplastic journal bearing diameters up to 3.00 in, the clearance can range from 0.008– 0.015 in/in of shaft outside diameter. Several design and service factors inﬂuence the clearance requirements. The applicability of each factor should be considered before the ﬁnal design.

347

5.5 Journal Bearing Design Principles

5.5.2

Thermoplastic Journal Bearing Axial Wall Thickness

Sheet metal housing

D

The thermoplastic journal bearing’s wall thickness should be as thin as possible, offering the following advantages: • Improving the dissipation of frictional heat • Reducing the journal bearing running clearance variations resulting from thermal and moisture related dimensional changes

Plastic bearing

D + 3%

3 lobes bearing

• Reducing the journal bearing distortion under high loading Thermoplastic journal bearing wall thicknesses range from 0.040–0.125 in for most typical applications.

Plastic bearing

5.5.3

Figure 5-10 Journal bearing, three-lobes assembly

Mounting Thermoplastic Journal Bearings

The following illustrations show a few possible solutions for securing thermoplastic journal bearings in sheet metal housings. In the case of low loads and speeds, the thermoplastic journal bearing can simply be snapped into a hole in the metal sheet and retained by an undercut as shown in Figure 5-10. However, the edges of the hole must not be sharp, because this would damage or shear off the undercut. Also, stamping burrs could make a proper snap-in impossible. The thermoplastic journal bearing can have an annotated groove around its periphery. From the point of view of molding technology, the solution shown in Figure 5-10 is preferred. Here the thermoplastic journal bearing is molded with three lobes equally spaced, where the shoulder is interrupted. In this way, simple axial ejection from the mold without undercuts is possible. Figure 5-11 shows a better solution, in which the stamped aperture is at the same time provided with a deep drawn collar providing better guidance of the thermoplastic journal bearing, especially if it is longer. This mounting design requires three or four snap ﬁttings with a slit in the middle; the undercut can be increased to 5%. Figure 5-12 shows that riveted-on bearings can become necessary, especially in case of wide operating temperature ranges, when loosening of the part is not acceptable. Metal rivets can be formed by ultrasonic riveting. Figure 5-13 shows a direct encapsulation of a journal bearing on a sheet metal housing. It is only economical if a number of other components are needed for assembly and can be molded-on at the same time. Using a vertical molding machine and a three-plate mold with three or four pin-point gates to control run-out dimensions, it is possible to encapsulate close tolerance sheet metal inserts. Thermoplastic journal bearings can be installed either by press-ﬁtting or with commercial adhesives. To press-ﬁt a thermoplastic journal bearing into a metal housing, the suggested practice is to use a low interference ﬁt. After it is pressed into place, the bore of the bearing will be reduced by 90% of the calculated diametral interference, which will result in a small compressive load in the journal bearing wall. A typical interference ﬁt is 0.5%; however, press ﬁt interference should be adjusted to the needs of the application.

D

Plastic bearing Sheet metal Stamped housing Snap fitting

D + 5%

Figure 5-11 Snap-ﬁtted journal bearing assembly

Sheet metal housing

Plastic bearing

Ultrasonic riveting

Figure 5-12 Riveted plastic bearing assembly

Sheet metal housing

Encapsulated plastic bearing

Figure 5-13 Encapsulated plastic bearing

348

5 Plastic Journal Bearing Design Waved split journal bearing

5.6

Split Bushing Thermoplastic Journal Bearings

Split bushing thermoplastic journal bearings, as shown in Figure 5-14, are mostly used for reduced loads, low speeds, and for an oscillating movement.

Spiral split journal bearing

Split bushing thermoplastic journal bearings are less susceptible to thermal and humidity conditions, because their clearances are only inﬂuenced by changes in wall thickness. In addition, assembly is simpler because no press-ﬁt is used. Split bushing thermoplastic journal bearings must be secured against rotation in the housing. Their use generally means a simpliﬁcation and is especially recommended for all manually operated and intermittently used shafts, such as rotating roller cranks, window actuators, hinges, locks, wheel suspensions, and similar applications.

5.7

Self-Centering Thermoplastic Journal Bearings

Self-centering thermoplastic journal bearings are easy to integrate, with their end caps having a snap-ﬁt joint or several other functions without higher costs. The designer has a wide variety of new design possibilities, which allow ingenious and simple solutions. Figure 5-15 shows a mounting ﬂange of a small motor with a ﬂexible suspension (self-centering) of the bearing to a limited extent. Slot split journal bearing

Figure 5-16 shows a small elastically suspended (self-centering) journal bearing, encapsulated into a sheet metal housing for the rotation of a metal shaft. Figure 5-17 shows a Thomson Spacer Nyliner® bearing, designed to permit the utilization of the advantages of unreinforced and internal lubricated nylon 6/6. The unique cone self-centering design eliminates the need for wide clearances to prevent seizures caused by thermal expansion. Figure 5-18 shows a connecting rod ball joint bearing housing with a snap-ﬁt securing ring, made of unreinforced nylon 6/6. The ball joint bearing is made of acetal homopolymer, ensuring low wear and friction even without lubrication. Figure 5-19 is similar in design to Figure 5-18, but with a spin welded securing ring for high axial loads.

Straight split journal bearing

Core Self-centering suspension Metal shaft

Snap-fit Mounting end flange

Figure 5-14 Split bushing thermoplastic journal bearings

Figure 5-15 Flexible suspension thermoplastic end ﬂange bearing

349

5.7 Self-Centering Thermoplastic Journal Bearings

Encapsulated sheet metal housing

Metal shaft Self-centering suspension

Figure 5-16 Elastic suspended thermoplastic end journal bearing

"C" lock ring Snap-fit ring, nylon 6/6

Metal shaft Ball joint bearing housing, nylon 6/6 Acetal, ball joint bearing

Figure 5-18 Snap-Fit connecting rod end ball joint bearing

"C" lock ring

Spin welded ring, nylon 6/6 Sphere retainer, nylon 6/6

Acetal, ball joint bearing

Ball joint bearing housing, nylon 6/6 Metal shaft

Figure 5-19 Spin welded connecting rod end ball joint bearing

Figure 5-17 Thomson spacer Nyliner® bearing (Courtesy: Thomson)

350

5 Plastic Journal Bearing Design Equal load distribution by shaft

Clearance Load W

Load W

Rotating Stationary journal bearing

shaft Rotating metal shaft

Rotating shaft

C Integrally cantilever thermoplastic bearing & housing

C

Figure 5-20 Stationary bearing load carrying contact surface

Uneven load reaction of bearing

Clearance

Bearing deflection

Fixed

Rotating journal bearing

shaft

Fixed shaft

Reaction R R

R Reaction Reaction

C

C

Figure 5-21 Rotating bearing load carrying contact surface Figure 5-22 Uneven load reaction across the length of a bearing

5.8

Journal Bearing Load Carrying Contact Surface (C)

Rib W

W

T Rotating shaft

Integrally ribbed cantilever thermoplastic bearing & housing

In the case of a stationary journal bearing and a rotating shaft, the load carrying contact surface will increase with increasing wear. Should the clearance between the shaft outside diameter and inside diameter of the bearing be too high initially, the speciﬁc surface pressure can be several times that of the calculated theoretical value. Wear in such a case will be rapid, as shown in Figures 5-20. In the case of a stationary shaft and a rotating journal bearing (for example a roller), the load carrying contact surface will decrease with increasing wear and running conditions will deteriorate, as shown in Figures 5-21. It can be seen that good distribution of the load requires as low an initial clearance as possible.

Rib (60% T)

5.9

Figure 5-23 Equal spaced reinforced ribs around bearing

Load Reaction Across the Length of Thermoplastic Bearing

Figure 5-22 shows a uniform shaft load transferred to a cantilever ﬂexible injection molded thermoplastic journal bearing, causing an uneven load reaction along the length of the integral journal bearing. Unequal distribution of the reaction load can cause high speciﬁc spots (edge loads), leading to the destruction of the journal bearing in a short time. As the safety factor is generally small for the thermoplastic journal bearings, this point should be given special attention.

351

5.10 Injection Molded Journal Bearings Process Defects Irregularly molded skin, warpage, inexact geometries of the bore as well as faulty assembly are the main reasons for an uneven distribution of the load. Because thermoplastics are elastic materials, the speciﬁc reaction loads can vary considerably in integrally injection molded thermoplastic journal bearings. Figure 5-23 shows an integral cantilever thermoplastic journal bearing supported by ribs and Figure 5-24 shows a thermoplastic journal bearing that is externally rib-supported. The thin ribs around the bearing distribute the load uniformly to the housing. This technique controls warpage and alignment, but a post molding machining operation is the best solution for critical run-out dimensional control. Figure 5-24 Thermoplastic housing with external ribbed bearing

5.10

Injection Molded Journal Bearings Process Defects

An incorrectly designed bearing will be impossible to mold correctly, even by the best molding processors. It is therefore the task of the designer to design all parts with the behavior of thermoplastics during injection molding in mind, in order to avoid warpage and to stay within the required tolerances. A single edge-gated journal bearing having a thick and broad shoulder will always show distortion and warpage, as shown in Figure 5-25. When the journal bearing is single edge-gated at the ﬂange (Figure 5-25), a more or less pronounced ovality is to be expected, especially if the ﬁrst stage injection speed and pressure are too high, or if the molding material has excessive mold shrinkage characteristics. If no subsequent machining of the bore has been provided for, the load at the bearing edge will always be higher than expected. Some thick walled or thick shoulder journal bearings, as shown in Figures 5-26 and 5-27, are the results of molding process problems. These molded bearings have shrinkage, warpage, and sink mark defects in the thicker cross sections, causing dimensional control problems; therefore these design techniques should be avoided.

Thick wall

A correct journal bearing design geometry with the appropriate type, size, and location of the gate, as shown in Figure 5-28, meets the following conditions:

Molding process dimensional defect problems

Molding problems

Single edge gate

Smaller diameter

Oval elongated circle

Shrinkage, warpage

Figure 5-25 Gate process effects on molded journal bearing

Sink marks

Figure 5-26 Molded effects on thick-walled journal bearing

352

5 Plastic Journal Bearing Design • The wall thickness is about 10% of the shaft diameter. Under no circumstances should it be more than 0.250 in.

Thick shoulder

• Using a sprue diaphragm or spider gate, three-plate molds with three or more individual pin point gates or a hot runnerless mold with several gate drops is recommended. Should the bearing bore be subsequently machined, the gates should be placed on the inside of the bore so that they can be removed in the same operation. • A continuous and uniform bearing ﬂange (shoulder) should be replaced by individual small lugs that will not cause deformation of the bore. Molding problems

The remarks concerning incorrect molding are of course just as valid for journal bearings that are integrally molded with metal plate inserts or housings. Journal bearings and other components integrally molded of self-lubricating resins should also be molded at process conditions correctly selected for these thermoplastic resins, otherwise serious molding defects may appear on the molded journal bearing:

Smaller diameter

• Orange peel defects on the bearing surface (internal and external)

Shrinkage, warpage

• Conical, oval, or sunken journal bearing inside diameter • An excessive amount of post-mold shrinkage; the journal bearing inside diameter may become smaller even if subsequently bored by machining

Figure 5-27 Molded effects on thick shoulder journal bearing

• Tolerances may be too wide.

T = 0.040 - 0.125 inch Sprue

As the geometrical accuracy is decisive for the behavior of a journal bearing, its bore should always be machined as long as the economics of the part allows it. With this measure, the above sources of failure introduced by faulty bearing, gate, mold designs, and processing conditions can be corrected to produce high quality control journal bearings. The post-mold shrinkage can also be controlled by annealing the journal bearings in an oil bath at a temperature just below the melting point of the resin.

Sprue puller

Gate depth = (40 - 60%) T Gate length = 0.040 inch Spider gate (3 eq. spaces)

5.11

Factors Affecting Journal Bearing Performance

Throughout this chapter, references are made to various conditions of lubrication. The descriptions of these terms are reviewed in the following, because they may be different for thermoplastic and metal journal bearings: • No Lubrication or completely unlubricated refers to a condition in which both the journal bearing and the shaft are wiped dry with a solvent. After solvent cleaning, the shaft and bearing are handled with gloves to avoid contamination with skin oils. No lubricant is added after installation or during the entire running operation.

Gate width = 75% T

Figure 5-28 Segmented ﬂanges bearing with sprue spider gate

• Initial Lubrication or lubricated at installation refers to the small amount of oil added when the journal bearing is assembled, but none thereafter. For example, three to ﬁve drops of oil for 1.00 in outside diameter shafts are a typical amount added.

5.12 Factors Affecting Journal Bearing Dimensions • Repeated Lubrication refers to a condition that goes beyond initial lubrication in that a few drops of oil are added infrequently during operation. This can be compared to a condition of occasional maintenance. • Continuous Lubrication refers to a constant supply of lubricant. Under this condition the coefﬁcient of friction and the wear rate are the lowest. This condition can approach boundary, hydrodynamic, or even force lubrication. Under continuous lubrication, the PV values are the highest. When testing a journal bearing, the use of duplicate bearings for each run is recommended rather than using a single bearing for all test conditions. The test can then best simulate ﬁeld conditions. This practice will eliminate unavoidable wear problems that might not be possible to determine in the ﬁeld. Because gradual increases of load and speed will raise the allowable PV value limit, failure in the ﬁeld may occur, if such a gradual initial wear-in period is not considered. In a journal bearing application, many factors are interrelated so that it is difﬁcult to separate them and distinguish between individual causes and results. The frictional properties of the self-lubricated thermoplastic journal bearings are inﬂuenced by load, speed, and lubrication. The frictional properties inﬂuence the wear rate, ambient running temperature, and performance such as drag and power loss. The effect of reduced clearance or increased angular contact is most evident in journal bearings, where the ends have bowed in slightly to produce a “barrel” effect. This effect can also be caused by stress relief of the molded thermoplastic bearing. If localized wear or if seizing occurs after the bearing has run successfully, these previous conditions may be the cause. The test samples should be inspected for roundness, and preferably be annealed and wiped clean. In most cases, correctly injection molded journal bearings will not exhibit these deﬁciencies.

5.12

Factors Affecting Journal Bearing Dimensions

Post-mold shrinkage, temperature, moisture, and load over time (creep) are frequently overlooked in journal bearing product design. The causes of dimensional changes are common to all materials, even if the values are small. The basic engineering principles are the same for self-lubricated resins as with any other material. Knowing the design properties of these materials enables the product designer to predict the effects caused by environment and load over time. The operating limits of a thermoplastic journal bearing are more sensitive to velocity than to bearing pressure. However, there are compressive stress limits that should be considered in a well designed bearing. The proportional stress limit (elastic range) at a tested strain and temperature is used for design analysis. These proportional stress limits are only a small percentage of the compressive strengths of the materials and are purposely limited to low values to prevent excessive deformation at high temperatures. For cool or light duty journal bearings, deformation is negligible. However, for journal bearings designed to operate near the limiting PV conditions, the proportional stress limits should be increased only after cautious experimentation.

353

354

5 Plastic Journal Bearing Design The physical property reductions under load and time (creep) are characteristics of thermoplastic materials. Their behavior is similar to that experienced with metals at high temperatures. For practical design purposes, the deformation at less than 1% strain after a one year period reaches a magnitude approximately equal to that of the initial strain. To estimate total deformation, the isochronous stress-strain curves at various temperatures are used for this analysis (see Chapter 2).

5.12.1

Length-to-Inside Diameter Ratio of Journal Bearings

A length-to-inside diameter ratio of 1 : 1 is recommended for journal bearing designs. For applications where the length exceeds the inside diameter, the journal bearing suffers localized spot wear and the spot melting of the bearing increases, probably caused by localized dimensional inaccuracies. In cases where the inside diameter exceeds the length, heat transfer usually becomes less effective (as if the wall thickness was too thick), causing higher wear of the journal bearing.

5.12.2

Types of Service and Motion of Journal Bearings

The clearance between the inside diameter of the journal bearing and the outside diameter of the shaft allow it to perform under the following conditions: • In intermittent operation • In frequent start-stop motion • In reciprocal motion • In linear or helical sliding motion • In occasional peak loading conditions • In air, water, or oil cooled operation Thermoplastic journal bearings submitted to these types of services or motion usually perform above the minimum PV value limit for unlubricated journal bearings, depending on the product design, the dimensional tolerances, the type of thermoplastic material used, and the manufacturing process.

5.12.3

Thermoplastic Journal Bearing Annealing Effects

Because thermoplastic molded journal bearings have been transformed from the stage of a polymer melt to a solid form by cooling, some residual stresses may be frozen-in. Over time, these stresses may relieve themselves. In a journal bearing, this is usually evidenced by reduction in clearance and subsequent seizing up of the journal bearing. The frozen-in stresses in the molded journal bearing can be relieved by applying heat. This process is referred to as annealing; a hot air oven or a hot oil bath are used (preferred). Annealing in oil also provides lubrication to the journal bearings. The annealing temperature should be higher than the expected service or ambient running temperature. The recommended annealing temperatures should be 50 °F (±3 °F) lower than the melting point of the resins. Care should be taken that the oil temperature does not exceed the recommended

355

5.12 Factors Affecting Journal Bearing Dimensions

Annealing of the self-lubricated journal bearings should be carried out without air, preferably by immersion in a suitable annealing oil. The oil should be stable at the annealing temperature; it should not attack chemically, nor give off noxious vapors. Reﬁned mineral oils (“Primol D”, “Uniﬂo”, or “Nujol”) have been found to be satisfactory for this process. Annealed journal bearings should be removed from the oil bath and cooled in air to room temperature, preferably in a draft free container. Accelerated cooling should be avoided, as it might introduce new thermal stress into the skin of the hot thermoplastic journal bearings. Post-molding annealing operations control the journal bearing’s dimensional variations and lubrication, assuming the correct processing conditions and design have been performed. However, post-molding annealing is an additional operation and should be avoided whenever feasible.

30 25

Annealing time, (minutes)

temperatures for longer than 1 minute for each 0.020 in of wall thickness, as it may cause deterioration of the physical properties of the journal bearing or cause warpage.

20 15 10 5 0 0

0.20

0.40

0.60

0.80

1.00

1.20

Part wall thickness, (inch)

Figure 5-29 Acetal homopolymer annealing time vs. wall thickness (Courtesy: Du Pont)

Figure 5-29 shows the annealing characteristics of acetal homopolymer resin.

Acetal Homopolymer Moisture Absorption Effects

The strain recovery and creep resistance of self-lubricated thermoplastic materials are excellent, particularly under high temperature and humidity conditions. The loads allowable for operation should be below the stress level causing deformation over time (creep). Therefore, self-lubricated thermoplastic resins seldom need design allowances to compensate for the creep variations.

Moisture content, (%)

1.6 1.2

se er

d

in

wa

te

r 1.3 1 .2

Figure 5-31 shows the time it takes for a part made from acetal homopolymer to change its dimensions based on the percent of moisture absorption at various temperatures. It can be seen that only prolonged service time while immersed in hot water (moisture and temperature) or similar conditions will cause dimensional changes.

2.0

1.10 th 1. 9 ng Le 0. 0.8 7 ase 0. cre In % 0.5 0.6 0.4 3 0. 0.2 1 0. 0 0.

Self-lubricated thermoplastic materials are affected by moisture from the environment. Although acetal homopolymer shows excellent stability over a wide range of humidity conditions, the dimensional changes of parts made from acetal homopolymer caused by service moisture content should be included in the journal bearing designs. Figure 5-30 shows a graph used to calculate approximate dimensional changes caused by temperature and moisture absorption of the acetal homopolymer resin.

Im

0.8 0.4

m

100% R.H.

50% R.H.

0 0

40

80

120

160

200

Temperature, (˚F.)

Figure 5-30 Acetal increase in length due to moisture vs. temperature (Courtesy: Du Pont)

2.0 Immersion in water (212˚ F.)

Moisture absorbed, (%)

5.12.4

1.6

5.12.5

TFE and Nylon 6/6 Moisture Absorption Effects

Unreinforced Teﬂon® TFE resin does not absorb moisture, therefore, it does not change dimensions with moisture variations. However, TFE reinforced composites may be subject to dimensional changes, depending on the type and percentage of reinforcement compounded with the basic resins. Unreinforced nylon 6/6 resins absorb moisture from the atmosphere and working environment. When the journal bearing clearance is 0.015 in/in, the moisture variations that cause dimensional changes in an air environment are not a problem. However, when the clearances are below 0.012 in/in, or the bearings are immersed in water, environmental effects should be considered in the design.

Immersion in water (140˚ F.)

1.2 Immersion in water (73˚ F.)

0.8 83% R.H. (73˚ F.)

0.4 50% R.H. (73˚ F.) 12% R.H. (73˚ F.) 0 0

10

20

30

40

50

60

Time, (days)

Figure 5-31 Acetal homopolymer moisture absorption vs. time at various temperatures and relative humidities (Courtesy: Du Pont)

356

5 Plastic Journal Bearing Design Nylon 6/6 journal bearings dry as molded normally contain less than 0.27% moisture by weight. Exposed to air at 73 °F and 50% relative humidity, the molded bearings will reach an equilibrium moisture content of 2.50%. The rate of moisture absorption and the resulting rate of dimensional change are relatively low. Preconditioning the journal bearings in boiling water or heating them in a salt solution of potassium acetate in water minimizes dimensional changes of nylon 6/6 in service. This procedure brings nylon 6/6 to an equilibrium moisture content at a faster rate and adjusts the journal bearings to the equilibrium dimensions of the anticipated service conditions.

5.12.6

Temperature Effects on Thermoplastic Journal Bearings

Temperature enters into the design of clearances for two reasons: all plastics have a coefﬁcient of linear thermal expansion roughly ten fold that of metals and plastics dissipate heat very slowly compared with metals. One of the ﬁrst steps in the product design should be to determine the temperature range over which the journal bearing must operate, using the upper and lower expected environmental temperatures at maximum velocity. Without the use of thermocouples, it is difﬁcult to measure actual equilibrium running temperature of the journal bearings. However, if the journal bearing is at all within the suggested pressure-velocity (PV) value limit, the equilibrium running temperature should not be critical. Thermal expansion must be considered for both lubricated and unlubricated journal bearings. Some designs may offer a partial solution but, basically, thermal expansion must be included until tests show that more latitude in clearances can be employed because of cooling or good heat dissipation through the shaft and housing. The coefﬁcient of linear thermal expansion of acetal homopolymer is 4.5 × 10–5 per °F. The greater the temperature range over which a journal bearing of acetal homopolymer must work, the greater the expansion involved. The journal bearings should have a minimum clearance of 0.005 in/in at the lowest service temperature at which operation is expected. Thermal expansion will work two ways: over the inside diameter and over the length of the journal bearing. The length should be free, unrestricted at one end to allow for thermal expansions. If the journal bearing is restricted at both ends and localized wear or spot melting occurs during testing, this is caused by an end compressive effect. Either longitudinal or diametral clearance should be enlarged. The low heat conductivity of self-lubricated journal bearings is the reason for quick heat build-up and possible spot melting or overall melting under excessive loads or small clearances. Therefore, the wall thickness of a journal bearing should be as thin as the design allows. The thinner the journal bearing wall, the better the heat transfer through the housing and the metal shaft.

357

5.12 Factors Affecting Journal Bearing Dimensions

5.12.7

Thermal Effects on Thermoplastic Journal Bearing Clearances

Thermal changes can cause a self-lubricated thermoplastic journal bearing to seize and consequently lead to its immediate destruction. If considerable changes of temperature are expected, it will be necessary to check their inﬂuence on the clearance of the journal bearing and to carry out suitable design modiﬁcations as necessary. For a journal bearing made of self-lubricated thermoplastic incoporated in a larger structure, as shown in Figure 5-32, the test data were developed using an acetal homopolymer journal bearing to determine the increased clearance with increasing temperature. The clearance should be speciﬁed and tested in relation to the lowest ambient temperature expected for the application. For example, the acetal homopolymer journal bearing inside diameter of 1.200 in will increase by 0.0024 in given an increase in temperature of 76 °F. Figure 5-33 shows a journal bearing made of self-lubricated acetal homopolymer pressed into a metal housing. The journal bearing clearance will decrease with increasing temperature as the wall expands toward the interior or inside diameter of the journal bearing. These dimensional changes can be neglected when using normal wall thicknesses. For example, given a wall thickness of 0.120 in and an increase in temperature of 105 °F, the clearance will decrease by only 0.00075 in. Should frequent temperature changes over a wide temperature range be expected, a journal bearing made of acetal homopolymer should not be secured by a press ﬁt alone, as over time all of the retaining force will disappear. The temperature rise produced by friction in an acetal homopolymer journal bearing will depend on the coefﬁcient of friction and the pressure-velocity (PV) value limit, therefore the heat conduction cannot be predicted with certainty. Therefore, it will be advantageous to determine the temperature of the journal bearing experimentally, for high loads and insufﬁcient cooling applications.

176

0.024

"d" dia.

Inside diameter increase, (inch)

140 0.016

105 86 76 68

0.008

60 0.004

50 0.0024 0.0016

0.0008

Integrally journal bearing and housing 0.0003

0.40

0.80

1.20

1.60

2.40

Inside diameter, "d" (inch)

Figure 5-32 Inside diameter increase vs. size and temperature changes (Courtesy: Du Pont)

3.15

4.00

Temperature increase, (°F.)

0.04

358

5 Plastic Journal Bearing Design

"d" dia.

Clearance "d" dia. Decrease, (inch)

T

Metal housing

140

0.0016 105 90

0.0008

70 60

0.0004 50

0.00024

Temperature increase, (°F.)

176

0.0024

Bearing pressed in metal housing

0.00016

0.00008

0.00004

Acetal bearing

0.02

0.04

0.06

0.08

0.12

0.20

Wall thickness, "T" (inch)

Figure 5-33 Clearance decrease vs. thickness and temperature changes (Courtesy: Du Pont)

For acetal homopolymer journal bearings, the maximum operating surface temperature caused by friction and external environmental conditions should not exceed 158–176 °F. For journal bearings made of unreinforced and internally lubricated nylon 6/6 resins, the journal bearing maximum operating surface temperature should not exceed 176–212 °F. If a good journal bearing performance and long service life are required for acetal homopolymer and unreinforced/internally lubricated nylon 6/6 resins, the maximum recommended end use temperatures for both types of compounded polymers should be included in the journal bearing designs.

5.13

The two most important criteria in the design of thermoplastic journal bearings are PV limits and wear. The PV limit of a thermoplastic in a given environment tells the designer if the material will work in the application. To be successful, the designer must have a basis for the prediction of the wear rate; otherwise, a prototype test for each new thermoplastic journal bearing is required.

500 Acetal homopolymer Mating material 1040 steel Finishing 16 RMS Hardness 22 R C Temperature (73˚ F.) No lubrication

Pressure, (psi)

400

300

The PV limit is the product of pressure (psi) and velocity (fpm). Figure 5-34 shows the PV limits curve for acetal homopolymer resin. The connecting points of limiting pressure and limiting velocity on pressure versus velocity coordinates describe the PV limits curve. This graph is developed for a speciﬁc type of resin and is useful to ﬁgure out the limits of the material when either the pressure or velocity is known.

200

Working area

100

Journal Bearing Pressure-Velocity (PV) Limits

PV limit 0 0

100

200

300

400

500

Velocity,(fpm)

Figure 5-34 Acetal homopolymer “PV” limit curve (Courtesy: Du Pont)

Any compounded plastic bearing resin sliding against another material surface without the assistance of a lubricant (oil, grease, etc.) at a given ambient temperature has a PV limit. The PV limit is caused by the plastic’s surface frictional temperature reaching or exceeding a critical PV value.

359

5.13 Journal-Bearing Pressure Velocity (PV) Limits Because the frictional heating caused by the sliding motion depends on the pressure, velocity, and coefﬁcient of friction, the surface temperature will depend on the same variables if heat loss parameters are unchanged. Measurement of surface temperature at various combinations of pressure and velocity would appear to be an excellent basis for determining the PV limit of a plastic. However, temperature affects compressive strain and compressive set under loads, so the maximum surface temperature at which a plastic will operate satisfactorily at high pressure and low velocity is not necessarily the same for low pressure and high velocity. The relation between a polymer’s surface temperature and its PV limit is an important notion, because it points out several factors in sliding element design. A plastic’s PV limit is decreased by increased ambient temperature and increased by decreased ambient temperature. This means that the PV limit of a plastic must approach zero as the ambient temperature approaches the critical temperature. Conversely, a plastic’s PV limit can be increased tremendously through cooling. Consideration should be given to providing an adequate heat transfer in all plastic journal bearing applications by using a thin wall thickness or by using a plastic compounded resin with conductive additives or reinforcements to improve the poor heat transfer properties of polymers. The plastic’s surface temperature is obviously important. However, a detailed study is not considered practical, because the journal bearing surface temperatures are difﬁcult to measure accurately and the use of this approach requires the designer to predict the surface temperature of each application.

50

Acetal homopolymer

5.13.1

Methods to Determine the PV Limits of Plastics

• Pressure Stepping Test With this method, a series of tests is conducted at several constant velocities. In each test, the contact pressure is increased in small increments, measuring the sample’s temperature and coefﬁcient of friction, until a pressure is reached at which temperature and coefﬁcient of friction equilibrium cannot be obtained or there are other signs of failure. Through proper selection of initial pressure and pressure increments, the duration of each test is reduced so that the effect of wear will be reduced. The highest pressure at which satisfactory operation is obtained multiplied by the test velocity gives the PV limit for the particular plastic at ambient temperature. • Wear Test Series The second method is to conduct wear tests at a series of pressures at several velocities. The results of each series are plotted as wear rate (or wear factor) versus pressure. The pressure at which this curve changes slope radically is the limiting pressure for the test velocity and ambient temperature. Examples of the wear test series curves for acetal homopolymer and TFE are shown in Figure 5-35.

5.13.2

Journal Bearing Coefficient of Friction

The coefﬁcient of friction is of prime concern to the bearing product designer. Yet, predicting accurately a speciﬁc coefﬁcient of friction for a plastic journal bearing is impossible. However, enough data have been developed to show the approximate levels of coefﬁcient of friction under most bearing conditions.

Wear factor, "K" (x 10-10)

40

Carbon filled TFE

30

20

10

0 0

40

80

120

160

Pressure, (psi) Mating material 1040 steel Finishing 16 RMS Hardness 22 RC Rubbing velocity 100 fpm Temperature 73˚ F. No lubrication

Figure 5-35 Acetal and TFE “PV” limits, wear test series curves (Courtesy: Du Pont)

360

5 Plastic Journal Bearing Design Teﬂon® has an exceptionally low coefﬁcient of friction at speeds lower than 20 fpm and at loads higher than 25 lbs. Under such conditions, a coefﬁcient of friction as low as 0.15 can be attained with thermoplastic compounds modiﬁed with a low percentage of coefﬁcient of friction additives. Unreinforced Teﬂon® TFE polymers running against steel have the lowest coefﬁcients of friction of all materials – as low as 0.04. The static coefﬁcient of friction is less than the dynamic coefﬁcient of friction. There are no slip-sticks, only smooth operation and easy breakaway. Unreinforced and internally lubricated nylon 6/6 has a coefﬁcient of friction of 0.20–0.40. Acetal homopolymer has a lower coefﬁcient of friction, between 0.15 and 0.35. The static and dynamic coefﬁcients of friction of acetal homopolymer are equal. Therefore, these journal bearings have low starting friction. Journal bearing tests have shown that there is no linear relationship between the wear factor and the coefﬁcient of friction. The smaller the clearance and the higher the coefﬁcient of friction, the higher the wear rate. A high coefﬁcient of friction causes frictional heat build-up fast enough to cause partial or complete melting of the inside surface of the journal bearing. However, a small clearance may increase the angular contact of the journal bearing with the shaft, in which case the coefﬁcients of friction will be higher.

5.13.3

Journal Bearing Failures Due to Small Clearances

A thermoplastic journal bearing subjected to excessive loads and speeds or small clearances with the shaft, fails because of a rapid heat build-up beyond the melting point. Figure 5-36 shows a small-clearance acetal homopolymer journal bearing. The internal surface temperature exceeded the melting point of the resin, causing surface imperfections due to the frictional heat build up during the test.

Figure 5-36 Overheated acetal bearing surface caused by small clearance (Courtesy: Du Pont)

Most thermoplastic journal bearing failures show that excessive loads and high speeds cause the wear rate to increase in a linear proportion to the increase in loads and speeds. If wear occurs under allowable loads and speeds, this is caused by a small clearance between the bearing inside diameter and the shaft. The excessive pressure-velocity or small clearance is the primary cause for frictional heat build-up. This condition is far above the maximum limit for the material, causing rapid wear and failure of the thermoplastic journal bearing. Nevertheless, even under normal pressure-velocity conditions, well within allowable PV limits, unsatisfactory wear rates can occur due to a small bearing clearance, the lack of lubrication, or the presence of metal debris, dust, rust, sand, and so forth. The cause of higher than expected journal bearing wear can also be the improper metal shaft mating surface (hardness and ﬁnishing). The following examples describe some combinations used for journal bearing resins and shaft materials without using a lubricant: • Acetal Homopolymer Bearing to Steel Shaft Acetal homopolymer bearings on steel shafts provide good performance, the lowest coefﬁcients of friction, and the lowest wear rates.

5.13 Journal-Bearing Pressure Velocity (PV) Limits • Acetal Homopolymer Bearing to Aluminum Shaft The coefﬁcient of friction is higher than steel. Aluminum shaft corrosion occurs in salt water and severely increases the wear of the acetal homopolymer. Aluminum has the tendency to ﬂake off, introducing minute particles in the bearing interface area, rapidly increasing wear. • Acetal Homopolymer Bearing to Acetal Homopolymer Shaft This combination is not the most desirable. When two equally low heat conducting materials are mated, the heat build-up is more severe. Under very light load and reduced speed conditions, this combination does well, but it produces squeaks, because high loads in an unlubricated operation produce tiny loose wear particles. • Acetal Homopolymer Bearing to Nylon 6/6 Shaft This combination offers good performance, because the frictional similarity and chemical dissimilarity of these two materials offer ideal design characteristics when a bearing of acetal homopolymer runs against a high strength nylon 6/6 shaft.

5.13.4

Definition of Different Types of Wear

The theory of solid friction and wear has been developed in great detail by Tabor, Bowden, Rabinowicz. Most theoretical studies on friction and wear have been conducted for combinations of metals and plastic materials. Adhesive Wear Whenever two materials are brought into contact, attractive forces tend to hold them together. When the surfaces are separated, the break does not occur at the original interface; rather, a cohesive failure may occur in one material. This transferred material may later transfer back to its original surface, forming loose wear particles. Adhesive wear generally produces ﬁnely divided wear particles and may transfer loose wear material to the mating surface. The resulting surface is quite smooth. Adhesive wear usually has a low wear rate; however, rapid wear and transfer from a polymer surface to a metal shaft have been observed for severe operating conditions. Abrasive Wear Abrasive wear occurs when a rough hard surface, or a soft surface containing hard particles, slides on a softer surface. These conditions generally produce a series of grooves in the softer surface and the displaced material usually forms loose wear particles. The same type of wear occurs when two smooth surfaces slide against each other in the presence of abrasive particles. Abrasive wear normally produces larger wear particles than adhesive wear, occurs at a faster rate, and produces a relatively rough wear surface. This type of wear is typical when a bearing surface becomes contaminated with abrasive particles, either from the environment or from an accumulation of wear particles. Corrosive Wear Corrosive wear results when sliding occurs in a corrosive atmosphere. Many materials form a surface ﬁlm that slows or stops corrosion. However, if this ﬁlm is continuously removed due to wear, corrosion of the base material will continue. Since many polymers are resistant to common corrosive chemicals, plastic bearings are often speciﬁed for corrosive environments.

361

362

5 Plastic Journal Bearing Design Thermoplastic journal bearings should be broken in for best wear performance. Fifteen minutes at one half the service conditions are usually more than satisfactory for nylon 6/6 or acetal homopolymer journal bearings. With journal bearings made of Teﬂon® TFE ﬂuorocarbon, break-in time is even less. If a ﬁfteen minute break-in period is impractical for the assembly process, reduce the design PV limit to a lower level of signiﬁcance in the break-in period. The break-in period can be eliminated either by applying a few drops of oil (initial lubrication) between the mating surface of the journal bearing inside diameter and the shaft outside diameter, or by annealing the journal bearings in oil at temperatures and time recommended by the resin supplier, without removing the oil from the inside surface of the journal bearing.

5.14

Mating Material Hardness and Surface Finishing

The wear performance of thermoplastic journal bearings can be substantially affected by the hardness of the mating material and its surface ﬁnish. The wear rate can be reduced by increasing the hardness and decreasing the mating surface ﬁnish. A ﬁne polish ﬁnishing operation of the steel shaft should be in the same direction as the journal bearing motion on the mating surface. Thrust bearing tester data without lubrication

Wear factor, (K)

1.2

1.0

0.8

0.6 20

30

40

50

60

Hardness, RC

0

10

20

30

Finishing, micro-inch (RMS)

Figure 5-37 Metal shaft hardness and ﬁnishing vs. plastic wear factor

40

Because lubricants are not needed in plastic bearing applications, metal shaft corrosion can be a problem. Stainless steel or other noncorrosive materials with proper hardness and ﬁnishing are recommended. If common steel alloys are used, it is recommended that they be chrome-plated. Where design limits permit, hard anodized or aluminum with a TFE hard surface spray coating of the shaft may be used. Aluminum and zinc are not good mating surfaces for plastic bearings because the softness of these materials can lead to rapid wear. Die cast aluminum with high silica content is very abrasive to plastics and the porosity of chrome plating may cause greater wear than a ﬁne polished steel surface. The metal surface of the shaft should be as hard and smooth as is practical. The shaft surface ﬁnishing for plastic bearings is not as critical as the one required for some metal bearings. During the Teﬂon® TFE fabric bearing break-in, it is believed that the shaft becomes coated with a thin ﬁlm of Teﬂon®. Sliding occurs primarily between the Teﬂon® coated shaft and the Teﬂon® TFE fabric. However, to prevent excessive wear, the surface ﬁnishing on the shaft should be better than 32 micro inches (RMS), a ﬁnish of 16 micro inches (RMS) being preferred. Plastic shafts are not a good mating material for plastic bearings; if used, the application should be limited to low PV limits. The softness of a plastic bearing mating surface can lead to high wear. Because plastics are relatively poor thermal conductors, the plastic-to-plastic bearing interfaces will run hotter than plastic-to-metal interfaces. The metal-to-plastic bearing interfaces have a higher PV limit than the plastic-to-plastic bearing interfaces. Plastic materials are poor heat conductors and it is necessary to provide means to carry away the frictional heat generated at the bearing/shaft interface. Figure 5-37 shows the effects of steel shaft hardness and ﬁnishing vs. wear factor for an unlubricated thrust plastic bearing.

5.15 Self-Lubricated Thermoplastic Journal Bearings

5.15

Self-Lubricated Thermoplastic Journal Bearings

One of the characteristics of self-lubricated thermoplastic materials is the low heat conductivity. Most thermoplastic materials have approximately the same thermal conductivity, although low friction additives and reinforcements are compounded to change this property of thermoplastic materials. Typical reinforcing materials are ﬁber glass and aramid ﬁbers. The loading levels vary from 10–40% by volume. Low coefﬁcient of friction additives are compounded with the matrix resins. Some of these additives are graphite, molybdenum disulﬁde, aluminum and calcium stearate, Teﬂon® TFE ﬂuorocarbon powder or ﬁber, silicone, carbon and aramid ﬁbers. Designing plastic journal bearings is similar to the procedures used for the metal bearings. The same basic engineering theories and equations are applicable. The functional properties of these thermoplastic self-lubricated bearing materials include the following: • When completely lubricated, the thermoplastic self-lubricated bearings perform similarly or better than metal bearings • They can be used with little or no lubrication at all • They have very low coefﬁcients of friction (0.30 for unlubricated and 0.04 for the thermoplastics self-lubricated) • The self-lubricated thermoplastics show no slip-stick behavior because there is no noticeable difference between the static and dynamic coefﬁcients of friction against steel surfaces Self-lubricated thermoplastic materials have other desirable engineering characteristics. They exhibit low creep and low deformation under loads even at high humidity and temperatures. With their high strength, they have enough resilience for press-ﬁtting, snap-ﬁtting, good impact strength, and a tendency to dampen vibrations produced by the journal bearing assembly. Self-lubricated thermoplastic journal bearings should not run hotter than the melting point. For example, acetal homopolymer journal bearings should not be used continuously above 176 °F and intermittently up to 200 °F. There is no accurate way of calculating the performance of a bearing, even if all the design parameters are known; the mating surface operating temperature of the bearing and shaft is difﬁcult to measure. However, an approximation can be reached which shows if a bearing will do well in an application. The PV limit is the product of the pressure load (psi) on the projected bearing area (in2) and the running velocity (fpm). Since the velocity, the total pressure load, and the shaft diameter are known, the PV limit value can be calculated. From this value, the design physical dimensions and tolerances for the journal bearing and shaft needed for a speciﬁc application can be determined. Most journal bearing applications are subjected occasionally to unexpected loads and vibration. It is considered good practice to limit the PV value to only 75% of the rated limit. For example, designing an acetal homopolymer bearing to operate at a velocity between 100 and 200 fpm, it is recommended to select a PV limit of 3,000 (Table 5-1) multiplied by 0.75 = 2,250 psi-fpm.

363

364

5 Plastic Journal Bearing Design The pressure versus velocity graphs of unreinforced nylon 6/6 and acetal homopolymer resins as determined by the limiting PV equations are shown in Figure 5-38.

70

Acetal @ 0.005 in./in. clearance PV1.20 = 7.600 psi-fpm

Bearing pressure, (psi)

60 50

Nylon 6/6 @ 0.015 in/in. clearance PV1.47 = 40,000 psi-fpm

40

Self-lubricated bar stock plastic materials can be machined to make prototype journal bearings that can be used for testing under the required end use conditions. The calculations should only be used to decide the design feasibility and to conﬁrm the results found in the plastic bearing prototype tests.

30 20 10 0 0

100

200

300

400

500

600

700

Velocity, (fpm)

Figure 5-38 Pressure-velocity graph per limiting PV equations (Courtesy: Du Pont)

800

Self-lubricated plastic materials for bearing applications include: phenolic, acetal homopolymer, nylon 6/6, ﬂuorocarbon, polycarbonate, polysulfone, polypropylene, and ultrahigh molecular weight polyethylene. These materials are used with heat treated (20–45 RC) and polished (16–32 RMS) steel shafts. For some speciﬁc applications, plastics such as phenolic, polycarbonate, polysulfone, acetal homopolymer, and nylon 6/6 are compounded with low coefﬁcient of friction additives used to decrease friction, such as ﬂuorocarbon (10–25% by weight), silicone lubricant (2% by weight), graphite, molybdenum disulﬁde, and stearates. Aramid ﬁbers and minerals are also used to increase strength, dimensional stability, and to decrease shrinkage. The design objectives for bearings are usually a long life, high efﬁciency, or both. Wear and friction are the major concerns in the design of bearings. The factors that control the performance of a bearing are pressure, velocity, lubrication, temperature, surface hardness, and ﬁnishing of the shaft. Nylon 6/6 and Teﬂon® TFE ﬂuorocarbon bearings have a tendency to cold-ﬂow under moderate pressure loads. The product of pressure and velocity (PV) is the power rating of the bearing per unit area. The product of PV and the coefﬁcient of friction is the energy dissipation (or the rate of heat generation). The wear rate is the product of the wear factor and PV. Increasing the bearing material hardness with reinforcements or metal support backing (composite bearings) can reduce the coefﬁcient of friction and the wear factor of a journal bearing. Self-lubricated thermoplastics are sensitive to heat. When the temperature rises above the melting point, the wear rate increases considerably. The best method of heat removal is by using oil or water through the metal shaft. The wear factor means the volume rate of material loss in a time period (in3/h). The bearing pressure is the load divided by the projected area of the bearing (psi). Table 5-1 shows the maximum operational conditions of several bearing materials.

Table 5-1 Plastic Bearing Materials Operating Limits

Bearing Material

Maximum pressure (psi)

Maximum temp. (°F)

Maximum velocity (fpm)

Maximum PV limit (psi-fpm)

Phenolic

6,000

200

2,500

15,000

Unreinforced nylon 6/6

1,000

200

1,000

3,000

500 2,500 60,000

500 500 350

50 1,000 150

1,000 10,000 25,000

Polycarbonate

1,000

220

1,000

3,000

Acetal homopolymer

1,000

180

1,000

3,000

Fluorocarbon Fluorocarbon (reinforced) Fluorocarbon (fabric)

365

5.15 Self-Lubricated Thermoplastic Journal Bearings Table 5-2 Coefﬁcient of Friction and Wear Factor

Bearing material

Shaft

Wear factor (K 10–10)

Coefﬁcient of friction Static

Dynamic

10,200 65 45

0.50 0.30 0.25

0.45 0.25 0.20

1,150 800

0.60 0.35

0.46 0.30

Acetal homopolymer Acetal homopolymer Acetal homopolymer (20% TFE)

Acetal Steel Steel

Unreinforced nylon 6/6 Unreinforced nylon 6/6

Nylon 6/6 Steel

Nylon 6/6 (20% TFE) Nylon 6/6 (30% glass, 15% TFE)

Steel Steel

30 45

0.25 0.35

0.18 0.26

Polysulfone (30% glass, 15% TFE)

Steel

65

0.40

0.28

Polycarbonate Polycarbonate (30% glass, 15% TFE)

Steel Steel

2,500 70

0.60 –

0.45 –

Table 5-2 shows the wear factor and coefﬁcient of friction (static and dynamic) of several thermoplastic materials. The wear factor and coefﬁcient of friction properties are not directly related. For example, some brake packing materials produce very high friction but have low wear, while a low coefﬁcient of friction material may wear rapidly. Wear factors of 200 or less are considered very good for most design applications. For bearings operating at a very low speed, or intermittent durations, wear is not a problem. Examples of thermoplastic bearings in such applications are: nylon 6/6 sliding pads and rollers used for furniture drawers, acetal homopolymer bearings used in door hinges, electrical appliances, and light duty wheels, ﬂuorocarbon sliding pads used between overhead highways and their supporting pillars to allow for thermal expansion, ﬂuorocarbon bearings used in automobile steering linkages and food processing equipment. Fluorocarbon fabric composite bearings have the highest load, temperature, and PV limit ratings, while phenolic bearings have the highest speed ratings. Phenolic bearings are used in marine propeller shafts and hydroelectric turbine shafts. Thermoplastic Journal Bearing Clearances Clearances for thermoplastic journal bearings are greater than those for metals to compensate for dimensional changes. More than half of all bearing problems are the results of insufﬁcient clearances. Increasing the clearances achieved the required performance in many applications. Their resilience and vibration absorption help these materials in overcoming the disadvantages caused by larger clearances. The recommended clearances for best performance of thermoplastic journal bearings operating in air at room temperature should be 0.005 in/in minimum, preferring a larger clearance up to 0.015 in/in. In very few cases, it is advisable to specify a diametral clearance of less than 0.004 in/in, even for small outside diameter shafts and plastic bearing applications.

366

5 Plastic Journal Bearing Design

5.15.1

Vespel® Polyimide Bearings

Vespel® bearings made of polyimide polymers do well with or without lubrication under operating conditions that destroy most other types of journal bearings and cause severe wear problems for metal journal bearings. Vespel® bearings reduce abrasion, corrosion, adhesion, fatigue, and wear problems that are typical for the conventional metal bearings, especially when they are operating without lubricants. The performances of some journal bearings in a given application depend on the following conditions: • The operating environment • Load or pressure on the bearing surface • Sliding velocity of the mating surface compared with the bearing • Hardness and ﬁnish of the mating surface • Frictional behaviour of the bearing material • Bearing thickness and the material’s ability to dissipate frictional heat • Coefﬁcient of linear thermal expansion of the bearing material Vespel® journal bearings have a higher pressure-velocity (PV) limit than most high performance engineering thermoplastic materials. Vespel® bearings are operated over a wide range of temperatures. They have excellent stress properties, outstanding creep resistance, abrasion resistance, and low coefﬁcients of friction. They exhibit excellent chemical resistance and have performed successfully in the following adverse environments: • Air and inert gasses at 700 °F • Gamma and electron beam radiation • High vacuum (10–10 torr) W

dS

• Liquid hydrogen and refrigerant (R-12, 134A)

Rotating shaft

N

• Hydraulic ﬂuids and jet fuels

When a bearing is subjected to increasing PV, it will eventually fail (PV limits). The failure point is usually manifested by an abrupt increase in the wear rate of the bearing material.

Journal bearing

As long as the mechanical strength of the journal bearing material is not exceeded, the temperature of the journal bearing surface is generally the most important factor in determining the PV limit. The parameters affecting a journal bearing operating surface temperature are: the coefﬁcient of friction and thermal conductivity of the journal bearing material, lubrication, ambient temperature, running clearance, hardness, and surface ﬁnish of the mating steel shaft.

DB Projected area (dB x L)

L dB

Figure 5-39 Journal bearing conﬁguration model (Courtesy: Du Pont)

Analysis of the plastic material requires calculating whether the PV in the application, derived from the journal bearing conﬁguration model as shown in Figure 5-39, is below 75% (safety factor) of the PV limit of the plastic journal bearing material. It is usually prudent to allow a generous safety margin in determining PV limits, because the real operating conditions for journal bearings are often more rigorous than the calculated or experimental conditions.

367

5.15 Self-Lubricated Thermoplastic Journal Bearings

5.15.2

Journal Bearing Pressure Equation

The journal bearing pressure P is determined by dividing the static load W per the projected surface area A that the journal bearing must withstand in operation.

P=

Static Load W = Projected Area dB × L

(5-3)

Where: P = Pressure bearing (psi) W = Static load (lb) DB = Journal bearing outside diameter (in) dB = Journal bearing inside diameter (in) DS = Shaft outside diameter (in) L = Journal bearing length (in) A = Journal bearing projected area = dB × L (in2) N = Shaft rotation speed (rpm) Thrust Bearing Pressure Equation The thrust bearing pressure P is determined by dividing the static load W per the projected surface area A that the thrust bearing must withstand in operation.

P=

Static Load 4×W = Projected Area π (D 2 − d 2 )

(5-4)

Where: P = Thrust pressure (psi) W = Static load (lb) d = Thrust bearing inside diameter (in) DM = Thrust bearing average diameter (in) D = Thrust bearing outside diameter (in) A = Thrust bearing surface area (in2) N = Rotating thrust plate speed (rpm)

Rotating thrust plate

Thrust bearing N

W

For either Vespel® bearing conﬁguration, the pressure (P) should not exceed the allowable values at room temperature, as shown in Table 5-3: D

Table 5-3 Vespel® Allowable Static Bearing Pressure (Courtesy: Du Pont)

Material

Vespel® SP-1

Vespel® SP-21

Vespel® SP-22

Vespel® SP-211

Fabrication

Mach

Direct

Mach

Direct

Mach

Direct

Mach

Direct

Pressure, psi

7,400

4,800

6,600

4,900

6,000

3,700

5,400

4,000

DM d

Table 5-4 Bearing Velocity (V) Equations

Type of service

Journal bearing

Thrust bearing

Continuous rotation

V = π (DS × N )

V = π (DM × N )

Oscillatory motion

V = π (DS × N ) (θ° /180)

V = π (DM × N ) (θ° /180)

Thrust bearing surface area

π A=

4

( D 2 - d 2)

Figure 5-40 Thrust bearing conﬁguration model (Courtesy: Du Pont)

368

5 Plastic Journal Bearing Design Where: V = Bearing surface velocity (in/min) N = Bearing speed of rotation (rpm) DM = Thrust bearing average diameter = (D + d) / 2 (in) D = Thrust bearing outside diameter (in) d = Thrust bearing inside diameter (in) DS = Journal bearing shaft outside diameter (in) θ° = Bearing angle of oscillation (degrees) Table 5-5 Plastic Materials PV Limit and Temperature (Courtesy: Du Pont)

Material

Reinforcements

PV Limit (psi-fpm)

Max. Temp. (°F)

Vespel® SP-21

15% Graphite

300,000

740

Vespel® SP-22

40% Graphite

300,000

740

Vespel® SP-211

15% Graphite + 10% TFE

100,000

500

PTFE PTFE PTFE PTFE

Unreinforced 15–25% Glass ﬁber 25% Carbon ﬁber 60% Bronze

1,800 12,500 20,000 18,500

500 500 500 500

Nylon 6/6

Unreinforced

3,000

215

Acetal homo. Acetal homo.

Unreinforced 20% TFE ﬁber

3,000 7,000

176 176

Pressure-Velocity (PV) Equation PV (psi-fpm) = P (psi) × V (in/min) / 12

5.15.3

(5-5)

Vespel® Wear Factor Effects Caused by Temperature

The PV limit is a very useful parameter in determining the suitability of a material for a bearing application. However, the contact pressure and the sliding velocity alone do not adequately describe the characteristics of the bearing materials. The operating temperature, bearing geometry, and mating material surface also play a signiﬁcant role in the wear behavior of the bearings. The temperature is generally the most important parameter, because it not only affects the coefﬁcient of friction but also determines the usable combinations of pressure and sliding velocity, or the PV limit. The wear factors of Vespel® bearings can be moderate even at high PV values, if sufﬁcient cooling is provided to the bearing. Wear can be severe at any PV value if the operating temperature is too high. The wear factor of a Vespel® bearing operating at temperatures below its limit can be predicted analytically or experimentally. The wear factor is derived from an equation relating the volume of material removed by wear in a given service time per unit of load and surface velocity. Wear volume (in3) = ν = K × W × V × ST Where: K = Wear factor (in3-min./ft-lb-h) W = Static load (lb)

(5-6)

369

5.15 Self-Lubricated Thermoplastic Journal Bearings ST = Service time (h) V = Bearing surface velocity (in/min) For ﬂat surfaces, Equation 5-6 is modiﬁed: Wear depth (in) = χ = K × P × V × ST

(5-7)

Where: P = Contact pressure (psi) K = Wear factor (in3-min./ft-lb-h) ST = Service time (h) V = Bearing surface velocity (in/min)

Vespel® Wear Transition Temperature

Figure 5-41 shows the Vespel® SP-21 bearing wear factor versus surface temperature in air. The wear factor is essentially constant over a wide range of operating conditions, because the bearing surface temperature does not exceed the wear transition temperature.

5.15.5

Frictional Behavior of Vespel®

Temperature, pressure, and velocity all affect the dynamic coefﬁcient of friction of plastic materials. Typical coefﬁcients of friction at various PV conditions for several self-lubricated Vespel® polyimide compounds are shown in Table 5-6 (Courtesy: Du Pont). The coefﬁcients of friction for Vespel® SP-21 and Vespel® SP-211 undergo a transition at a temperature of 300 °F, similar to that of unreinforced nylon 6/6, as shown in Figure 5-42. Below this temperature the frictional forces drop sharply in the range between 380 and 700 °F. The coefﬁcients of friction of both Vespel® materials are independent of temperature. The friction transition is not associated with wear transition.

Table 5-6 Vespel® Coefﬁcient of Friction (Courtesy: Du Pont)

PV conditions Static P = 50 psi, V = 500 fpm P = 100 psi, V = 100 fpm P = 100 psi, V = 300 fpm P = 100 psi, V = 1,000 fpm P = 1,000 psi, V = 315 fpm

SP-21

SP-22

SP-211

0.12

0.15

0.30

0.11 0.10 0.09 0.07 0.04

0.12 – 0.10 0.09 –

0.12 0.11 – 0.08 –

Steel thrust bearing tester in air PV values between 1.000 and 500.000 (psi-fpm) Wear factor, "K" x 10-10 (in3 -min/ft-lb-hr)

The wear rate of a plastic material operating in air is proportional to the product of pressure and velocity (PV), if the surface temperature does not exceed a critical value called “Wear Transition Temperature”. Above the wear transition temperature, wear increases dramatically. For Vespel® SP-21 material, the wear transition temperature ranges between 900 and 1,000 °F in a vacuum or gasses, and in air it ranges between 700 and 750 °F.

250.0

Vespel  SP-21 unlubricated

200.0

150.0

100.0

50.0

0 32

212

390

570

750

930

Surface temperature, (˚F.)

Figure 5-41 Vespel® SP-21 wear factor vs. surface temperature (Courtesy: Du Pont)

Coefficient of friction, (ƒ)

5.15.4

Against an unlubricated carbon steel 0.40

Vespel  SP-211

0.30 0.20

Vespel  SP-21

0.10 0 32

212

390

570

750

930

Surface temperature, (˚F.)

Figure 5-42 Vespel coefﬁcient of friction vs. surface temperature (Courtesy: Du Pont)

370

5 Plastic Journal Bearing Design

Wear factor, "K" x 10-10 (in3 -min/ft-lb-hr)

100.0 90.0 80.0

The product designer must compensate for the higher frictional forces, caused during the start up and each restart of the bearing. First, the Vespel® polyimide bearing transfers a polymer layer to the mating surface only at start up. Second, the transition temperature must be compensated for the different types of Vespel® compounds, as shown in Figure 5-43. During the restart, breaking-in a new layer is not necessary under service conditions, but the transition temperature is reversible and will continue to operate at each restart.

Vespel  SP-21

Vespel  SP-211

70.0 60.0 50.0 40.0 30.0 20.0 10.0 0 32

212

390

570

750

930

Surface temperature, (˚F.)

Figure 5-43 Vespel wear factor vs. surface temperature against an unlubricated carbon steel (Courtesy: Du Pont)

DB

5.15.6

Vespel® Journal Bearings Length to Inside Diameter Ratio

For optimum performance of Vespel® journal bearings (using Figure 5-44 as a geometric model), the length per inside diameter (L / dB) ratio should be between 0.75 and 1.25. If a long journal bearing is required in an application, we recommend substituting the long bearing with two shorter journal bearings separated by an air gap in the middle. Small L / dB ratios offer the following advantages: • More efﬁcient debris removal (air gaps in the middle) • Less sensitivity to shaft deﬂection and misalignment • Better frictional heat dissipation

dB

• Cost advantages due to lower fabrication costs of short bearings

L

Figure 5-44 Journal bearing length to inside diameter ratio

d

5.15.7

D

For optimum performance of Vespel® thrust bearings (using Figure 5-45 as a geometric model), it is recommended that the thrust bearing outside diameter per inside diameter ratio (D / d) does not exceed a value of 2. Ratios greater than 2 can cause overheating at the outside edge; problems may arise from lack of ﬂatness and trapping wear debris within the thrust bearing mating surfaces.

5.15.8 Figure 5-45 Vespel thrust bearing ratio between diameters

Vespel® Thrust Bearing Ratio Between Diameters

Vespel® Journal Bearing Initial Clearance (cI)

Vespel® polyimide compounds have much lower coefﬁcients of linear thermal expansion than some alloy steels and most other thermoplastic materials. For optimum performance, the typical clearance between the Vespel® journal bearing inside diameter and the shaft outside diameter should be between 0.3 and 0.5% of the shaft diameter. Usually, heavier loads require larger clearances. Diametral clearances for Vespel® journal bearings do not have to be adjusted for moisture, because Vespel® polyimide absorbs very little moisture. The initial journal bearing clearances should be calculated by allowing for the circumferential thermal expansion of journal bearing, shaft, and housing. cI = DS × (α S × ∆TS + c) + 2 t × α B × ∆TB

(5-8)

371

5.15 Self-Lubricated Thermoplastic Journal Bearings

5.15.9

Vespel® Journal Bearing Inside Diameter (dB)

Vespel® journal bearing inside diameters dB can be calculated by using the following equation: dB = DS (1 + αS × ∆TS) + c + 2 t × αB × ∆TB – DH × αH × ∆TH

(5-9)

Where: DS = Shaft outside diameter (in) DH = Housing inside diameter (in) cI = Journal bearing initial clearance (in) c = Journal bearing typical clearance (in) αS = Shaft material coefﬁcient of thermal expansion (in/in/°F) αB = Bearing material coefﬁcient of thermal expansion (in/in/°F) αH = Housing material coefﬁcient of thermal expansion (in/in/°F) t = Journal bearing wall thickness (in) ∆TS = Shaft temperature variations (°F) ∆TB = Journal bearing temperature variations (°F) ∆TH = Housing temperature variations (°F) Table 5-7 Vespel® Coefﬁcient of Linear Thermal Expansion (Courtesy: Du Pont)

Material

Vespel® SP-1

Vespel® SP-21

Vespel® SP-22 Vespel® SP-211

Fabrication

Mach

Direct

Mach

Direct

Mach

Direct

Mach

Direct

30

28

27

23

21

15

30

23

α, 10 in/in/°F –6

Example 5-1: Vespel® Journal Bearing Design An automotive gear box device needs a Vespel® journal bearing design analysis to meet the following requirements: : Two journal bearings operating on a 1.50 in diameter shaft must support a total load of 4,000 lbs and operate at temperatures ranging between 73 and 525 °F. : The maximum journal bearing length is 1.50 in and the maximum allowable clearance between the bearing inside diameter and the shaft outside diameter is 0.017 in, in hot or cold conditions. : To prevent the contamination of sensitive surroundings, the bearings cannot be lubricated and they must operate without service 40 hours per week for three years. : The gear box shaft rotates at 20 rpm, the operating sequence is intermittent, 5% running and 95% in stop mode. Determine if Vespel® SP-21 journal bearings will meet these requirements. Solution 1) Determine the maximum journal bearing temperature Refer to Table 5-5: the maximum surface temperature of Vespel® SP-21 in air is 740 °F. The differential surface temperature between 740 °F and 525 °F should be 215 °F.

372

5 Plastic Journal Bearing Design

2) Calculate the PV limit 2.1) Calculate the bearing pressure: P=

2,000 (lb./bearing) W = = 889 psi L × DS 1.50 in × 1.50 in

2.2) Calculate the surface shaft velocity: V =

π × DS × N π × 1.50 × 20 = = 7.85 fpm 12 12

PV = 889 psi · 7.85 fpm = 6,978 psi-fpm At this low PV (6,978 psi-fpm), Vespel® SP-21 (300,000 psi-fpm), the PV value will not be critical, especially considering the intermittent operation. 3) Calculate the wear depth (χ) 3.1) Calculate the running time (ST): ST = 0.05 · 40 h/week × 52 week/yr × 3 yr = 312 h Wear depth (χ) = K × PV × ST Wear factor K (33 × 10–10 in-min/ft-lb-h) according to Figure 5-43 Wear depth (χ) = 33 × 10–10 × 6,978 × 312 = 0.0071 in This wear depth is less than the maximum allowable clearance of 0.017 in. The clearance and the wear depth difference are enough to adapt the thermal expansion, therefore, Vespel® SP-21 journal bearings meet the requirements for this application. 4) Calculate the journal bearing clearance Practical experience and good judgment lead to the assumption that the shaft diameter, journal bearing surface inside diameter, bearing outside diameter, and the housing will all reach different operating temperatures. The following assumptions are made: : The journal bearing contact surface reaches 100 °F higher than the gear box temperature (525 °F + 100 °F = 625 °F) : The journal bearing body average temperature is only 50 °F higher than the gear box (525 °F + 50 °F = 575 °F) : The housing remains at ambient temperature and restrains the journal bearing securely : The journal bearing surface inside diameter will expand inward when the temperature rises : The shaft surface outside diameter will expand outward With these assumptions, the initial room temperature clearance (cI) can be calculated by using Equation 5-8: cI = DS (αS × ∆TS + c) + 2 t × αB × ∆TB

5.16 Teﬂon® (TFE) Fabric Composite Bearings

Where: cI = Journal bearing initial clearance (in) DS = Shaft outside diameter (1.50 in) c = Journal bearing typical clearance (0.002 in) t = Journal bearing wall thickness (0.0625 in) αS = Shaft coefﬁcient of thermal expansion (6 × 10–6 in/in/°F) αB = Bearing coefﬁcient of thermal expansion (23 × 10–6 in/in/°F) ∆TS = Shaft temperature variation (625–73 = 552 °F) ∆TB = Journal bearing temperature variation (575–73 = 502 °F) We select the Vespel® journal bearing wall thickness (t) to be 0.0625 in. CI = 1.5 [(6 × 10−6 ) (552) + 0.002] + (2 × 0.0625 × 23 × 10−6 ) (502) = 1.5 (0.0053) + 0.0014 = 0.00796 + 0.0014 = 0.0093 in 5) Calculate the maximum journal bearing clearance The maximum journal bearing clearance is equal to the initial clearance (cI) plus the wear depth after three years. Maximum clearance = Initial clearance (cI) + Wear depth (χ) = 0.0093 + 0.0071 = 0.0164 in Therefore, the Vespel® journal bearings satisfy the maximum clearance speciﬁcations of 0.017 in to operate without lubrication at the given elevated temperature (525 °F) for the automotive gear box requirements.

5.16

Teflon® (TFE) Fabric Composite Bearings

Teﬂon® TFE woven fabrics have been made into a variety of bearing types, including spherical, ball joint, thrust, and journal bearings. In addition, nonwoven and knitted fabrics also have been used for special applications. There are two basic requirements for the successful use of the Teﬂon® TFE fabric as a bearing surface: • An adequate bond of the Teﬂon® TFE fabric to the substrate • Adequate support of the individual ﬁbers or ﬁber bundles in the fabric Teﬂon® TFE woven fabrics can be bonded very satisfactorily with conventional adhesives or bonding techniques by use of a double fabric in which one face is composed of cotton, an easily bondable ﬁber, and the other face composed in part or totally of Teﬂon® TFE fabric. Another method is to modify the surface tension of the face of a twill fabric of Teﬂon® TFE ﬁber. The raised ﬁbers in the fabric are trapped in a suitable adhesive, forming a mechanical bond. Many fabric structures have been used adequately to support the Teﬂon® TFE ﬁber in the bearing. The type of bonding agent used is one that has good load bearing properties and is applied so that good penetration of the voids of the fabric is obtained. Phenolformaldehyde resins are effective, because of their load carrying ability.

373

374

5 Plastic Journal Bearing Design

Figure 5-46 Teﬂon® ﬁlament weaving for composite bearing (Courtesy: Du Pont)

Metal backing

Teﬂon® ﬁlament weaving creates structures of exceptional strength to support the bearing surface. These are, by their nature, truly concentric and have no seam or overlap. They can handle high radial and longitudinal stresses while supporting bearing pressure loads of 60,000 psi. They are also resistant to shock loading and abrasion. Their high strength makes possible the use of thin walled (0.062–0.125 in) composite bearings with a coefﬁcient of linear thermal expansion similar to steel that also allows the dissipation of heat through the metal shafts. Low coefﬁcient of friction and high load carrying capacity come naturally with the Teﬂon® fabrics used for composite bearing surfaces. With tensile strengths 20 times greater than straight PTFE resins, they have excellent cold ﬂow resistance and have shown some remarkable wear characteristics. Figures 5-46 and 5-47 show Teﬂon® composite bearing structures.

5.16.1

Fabric Adhensive

Figure 5-47 Teﬂon® fabric composite bearing (Courtesy: Du Pont)

Bearing Physical Properties

Teﬂon® fabric bearings are made by the most advanced fabrication processes under strict quality control. This ensures uniformly high strength in the supporting structures and a strong bond to the Teﬂon® ﬁlament bearing surface. These structures will support a bearing load pressure of up to 60,000 psi, and the bearings meet operating requirements of high load pressures or high speeds in rotational and stroking movements and in oscillation applications. Teﬂon® fabric bearings have exceptional dimensional stability. A boiling water test produced no swelling or shrinkage. The linear thermal expansion is similar to steel and in fact less than some steels. The thin wall of the bearings helps the heat transfer away from the bearing, through the metal shaft surfaces. The bearings may be used in applications requiring operating temperatures from –300 °F to +350 °F. The speciﬁc gravity is lower than for bronze or carbon steel and is lower than other plastics.

5.16.2

Bearing PV Limit Rating

For many years, the plastics industry has used the pressure-velocity (PV) limit for evaluating performance of Teﬂon® fabric composite journal bearings. Knowing the PV limit of a bearing, a designer can determine the loads and surface running speeds under which a bearing can safely operate. However, there is another factor to consider. Heat generated by friction is a major cause of polymeric degradation. The rise in temperature is dependent on the running speed and is not a linear function of the PV limit. To evaluate the operating conditions of a composite journal bearing, a designer needs to know the approximate temperature generated at the wear surface. For Teﬂon® fabric bearings this temperature can be as high as 350 °F. As a guideline, a Teﬂon® fabric bearing has a maximum 25,000 PV limit. However, test results run at 15,000 PV gave only 0.002 in total wear after 10,000,000 cycles, with a 25° oscillation motion at 600 cycles per minute and a radial load of 343 pounds.

5.17 Thermoplastic Kevlar® Reinforced Bearings

5.16.3

Journal Bearing Clearances (c)

The effect of clearances on wear may not be as critical for Teﬂon® fabric composite journal bearings as for self-lubricated plastic or metal bearings. However, an important design consideration is allowance for linear thermal expansion of the Teﬂon® fabric composite journal bearing backing support, subjected to a wide range of service temperatures or continuous operation at surface speeds above 150 feet per minute. Table 5-8 provides clearance guidelines for sleeve bearings of Teﬂon® fabric composite journal bearings in plastic housings. Table 5-8 Journal Bearing Clearances (c) (Courtesy: Du Pont)

Shaft diameter (in)

Clearance (in)

0.50

0.001–0.003

1.00

0.002–0.005

2.00

0.004–0.007

Smaller clearances may be used when working with metal-backed bearing supports, as the coefﬁcient of linear thermal expansion of the thin layer of Teﬂon® can usually be ignored. Heavy Equipment Applications On crawler type tractors, heavy earth moving equipment, and farm machinery, Teﬂon® fabric composite journal bearings with their freedom from oil and grease are right at home in the dusty and sandy conditions in which this equipment operates. These bearings have no problems with rust, corrosion, and they do not need oil or grease seals that require more space and add cost.

5.17

Thermoplastic Kevlar® Reinforced Bearings

Thermoplastic resins that are reinforced with Kevlar® aramid chopped ﬁbers are injection moldable composites. Kevlar® possesses a combination of physical properties that cannot be found in any other commercially available ﬁber. Using Kevlar® composites, design engineers have a new family of superior wear and abrasion resistant thermoplastic materials from which to choose. The advantages offered by Kevlar® aramid chopped ﬁbers include: • Superior wear resistance • Non-abrasive bearing inside surface • Signiﬁcantly improved mechanical properties • Higher PV limit • Higher operational temperatures • Lighter bearing weight • Excellent process characteristics without wear on equipment and molds • Greatly improved machinability

375

376

5 Plastic Journal Bearing Design Several thermoplastic resins reinforced with Kevlar® aramid ﬁbers are commercially produced for the plastic industry, among them acetal homopolymer with 15% Kevlar®, nylon 6/6 with 17.5% Kevlar®, PET with 17.5% Kevlar®, polyphenylene sulﬁde (modiﬁed) with 25% Kevlar®, polypropylene with 17.5% Kevlar®, copolyester elastomer with 10% Kevlar®, and TPE polyurethane with 10% Kevlar®. All these composites are produced as injection moldable pellets ranging in size from 0.25 to 0.375 in. The addition of Kevlar® fortiﬁes the positive characteristics of the matrix resins, while reducing the abrasion wear problems. For superior wear resistance, Kevlar® aramid ﬁber composites are the nonabrasive answer. Many thermoplastic composites can achieve low wear rates by using a reinforcement that adds to their strength and/or stiffness. Most, however, also show a correspondingly high abrasiveness to the bearing surface. Kevlar® aramid ﬁber thermoplastic composites are the exception. The tough, strong reinforcing material makes the product extremely wear resistant without excessive galling to the mating wear surfaces. In addition, it improves mechanical properties and increases the operational temperature capabilities. In tests conducted using unreinforced nylon 6/6, erosion of the steel washer was more than 40 times greater than for nylon 6/6 reinforced with Kevlar® aramid ﬁbers. Nylon 6/6 parts wore 70% faster than nylon 6/6 parts reinforced with Kevlar®. Wear tests using unreinforced nylon 6/6 molded parts and nylon 6/6 molded parts reinforced with Kevlar® resulted in equally dramatic performance differences. Both parts were tested for 240 hours at a pressure of 250 psi and a velocity of 10 ft/min. The ﬁnal wear factor for unreinforced nylon 6/6 was 917. For nylon 6/6 parts reinforced with Kevlar®, the wear factor was 239, resulting in unreinforced nylon 6/6 wearing four times faster than nylon 6/6 reinforced with Kevlar® aramid ﬁbers. Figure 5-48 Components made from thermoplastic materials reinforced with kevlar aramid ﬁbers (Courtesy: Du Pont)

Figure 5-48 shows several applications made of thermoplastic materials reinforced with Kevlar® aramid ﬁbers.

377

6

Thermoplastic Molded Spring Design

6.1

Introduction

Thermoplastic resins, such as acetal homopolymer, nylon 6/6, polyethylene, polypropylene, TPE, and other materials have excellent ﬂexibility, hysteresis, and mechanical properties for injection molding ﬂexible products that function as springs. With these materials, thermoplastic springs can be incorporated into multifunctional product components, reducing the number of parts required for the application, lowering manufacturing, assembly, and inventory costs. Thermoplastic molded spring applications include cameras, water pressure systems, irrigation, printing mechanisms, door closures, electrical lock connectors, pen caps, tubing ﬂow metering, transportation equipment, a wide range of general industrial devices, and consumer products. The most critical material requirements for injection molded thermoplastic springs include rigidity, good fatigue resistance, high load carrying capacity, minimum creep, self-lubricity, and ease to process into thin, complex multifunctional components. Depending on the speciﬁc end use application, thermoplastic molded springs may be required to perform under a wide range of temperatures, in a corrosive atmosphere, or under other hostile environmental conditions. One of the most frequently selected thermoplastic resins for spring applications is acetal homopolymer. This polymer provides the best balance set of mechanical properties and molding process conditions. Other low creep resistance and high strength thermoplastic resins, such as polyethylene, polypropylene, TPE, polycarbonate, unreinforced, and glass ﬁber reinforced nylon 6/6 are also used for springs. To ensure a thermoplastic molded spring’s effective performance over the life of the product, they should be designed to operate at zero strain in the normal relaxed position. Springs should not be designed to ﬂex, storing mechanical energy on a continuous basis, because, even under low stress levels, the continuous load will cause a reduction in its spring characteristics over a long period. Thermoplastic molded springs should be designed to work for intermittent loading and without strain the rest of the time. Several types of thermoplastic molded springs have been used in many critical applications. Case histories of commercial applications and some general observations on these spring families are presented in this chapter. These spring families include helical compression and extension, cantilever, leaf, helical torsion, spiral torsion, belleville, lock, and wave washers.

378

6 Thermoplastic Molded Spring Design

6.2

Thermoplastic Molded Spring Design Considerations

Injection molded thermoplastic spring characteristics should meet the following considerations: • Compensate for temperature and chemical environment effects on the mechanical properties of the thermoplastic springs • Design thermoplastic springs for intermittent loads without exceeding the proportional stress limit and fatigue resistance under the end use operating conditions for the service life of the product • Sharp corners should be avoided, use generous ﬁllet radiuses Helical compression

Helical extension

Single leaf spring

• Spring designs based on constant strength cantilever beam equations operate at lower levels of stress than other spring families Not many injection molding thermoplastic materials satisfy all these spring requirements. For instance, reinforced materials that have appreciable ultimate yield strength and a high modulus of elasticity, also have some reduced elastic recoveries under signiﬁcant strains; therefore, these materials are not recommended for spring applications. The selection of the most suitable injection molding thermoplastic resin for spring applications requires a very careful examination of the properties, the ambient operating conditions (temperature, chemical resistance, etc.), types of loading, and additional requirements of the spring application, such as creep, wear resistance, dimensional tolerances, spring design complexity, molding process, and so forth.

Belleville washer

Wave washer

Acetal homopolymer, polycarbonate, and unreinforced nylon 6/6 resins possess a combination of properties that make them particularly suitable for the production of ﬂexible and elastic components with the spring characteristics.

6.3

Lock washer

Thermoplastic Helical Compression Springs

Free Load ed

Fixed

Free

Loaded

Helical torsion

Spiral torsion

Figure 6-1 Common types of metal spring designs

The use of injection molding thermoplastic materials for helical compression springs is restricted to very special applications. The most elementary and least demanding application is the polyethylene spring on the stopper for pharmaceutical containers. Polyethylene is not used in precision spring applications, because the force provided by a polyethylene spring is very low and requires a considerable deﬂection. Helical springs are ideal for the metal wire winding process; however, these geometries are very difﬁcult for injection molding. Figure 6-1 shows common types of metal springs available on the market. Figure 6-2 shows a molded helical compression spring for a pen actuator (difﬁcult), a reticulated notion, where the curved beams are converted into a bending spring (small strain) with higher spring constant, and a system using Belleville spring washers that provide the operating characteristics of a helical compression spring. Thermoplastic springs have a very low spring constant compared with metal helical springs of the same dimensions. It is very important that spring constant value four the round cross section of metal helical springs is used for metal springs

379

6.4 Thermoplastic Molded Cantilever Beam Springs only and not for compression springs made of thermoplastic material. This is because of the considerable value difference of G (material torsional modulus of elasticity) between the metallic and thermoplastic materials. Pen actuator spring

For example, the ratio between the torsional modulus of elasticity (G) of steel and acetal homopolymer is:

W

G (Steel) 8,500 = = 63 G (Acetal Homopolymer) 135 Calculating a thermoplastic helical spring’s compression is possible by using the engineering equations for metal springs. However, the design and construction of the mold are difﬁcult, complex, and expensive. Running and maintaining the mold in production creates too many manufacturing problems. Reticulated spring

6.4

Thermoplastic Molded Cantilever Beam Springs

Injection molded thermoplastic cantilever springs are the most common type of springs made and used in many applications. There are considerable advantages compared with the use of metals, even more so when the fundamental rules of multifunction designs are adopted. Figure 6-3 shows four types of cantilever spring designs and a stress versus weight comparison graph using the classical cantilever beam equations. The graph is a comparison between the different cantilever spring geometries that produce an equivalent spring rate. The cantilever spring (A) has a constant rectangular cross section and an initial spring rate calculated from the deﬂection equation for a cantilever beam (W / δ = E I / L3) where W is the load and δ is the deﬂection at the free end of the beam. The other cantilever springs were designed to provide an identical spring rate, using the equations for constant strength beams. This results in lower stress levels and, sometimes, in a reduction in spring weight. For example, in the cantilever spring (B) the stress is 75% of that developed in cantilever spring (A) and the weight is reduced by 25%. This weight reduction can be very important as a cost savings factor when a large injection molding production run is contemplated.

Stack of belleville springs

Figure 6-2 Thermoplastic molded compression spring illustrations

b W g "A"

h

Spring "B

Tensile stress, (x 103 psi)

Sprin

W

" h

L

L

b b

b

2

A

B C D

1

W

W Sprin

3

Sprin

g "C"

g "D

"

2.25 h L

h L

4.50

Spring weight, (grams)

Figure 6-3 Cantilever springs, tensile stress vs. weight comparison

6.75

380

6 Thermoplastic Molded Spring Design If sheet metal stamping or forming operations are considered for fabrication, the sheet metal manufacturing cost is more expensive than for the thermoplastic injection molding process. ε=

σ E

(6-1)

σ =

WL Z

(6-2)

I=

b h3 12

(6-3)

Z =

b h2 6

(6-4)

W =

Z×E×ε L

(6-5)

Where: δ = Deﬂection (in) σ = Stress (psi) Z = Section modulus (in3) ε = Strain of beam at root (in/in) E = Modulus of elasticity (psi) W = Force of deﬂection (lb) L = Length of beam (in) b = Width of beam (in) h = Beam thickness at root (in) I = Moment of inertia (in4) Figure 6-4 shows four cantilever beam spring geometries with their deﬂection equations used in the analysis of thermoplastic and metal materials. More detailed information about cantilever latch beams and the strain limits for common types of resins can be found in Chapter 8.

W h

b

W

h

b

L δ = 0.67 x

ε x L2 h

b 4

L 4 x W x L3 b x h3 x E

δ =

ε x L2 δ = 0.86 x h

δ =

W

W

b

5.15 x W x L3 b x h3 x E

h b h

h 2

L δ = 1.09 x

ε x L2 h

δ =

6.50 x W x L3 b x h3 x E

L δ = 1.33 x

ε x L2 h

Figure 6-4 Cantilever spring geometries and deﬂection equations

δ =

8.00 x W x L3 b x h3 x E

6.5 Cantilever Beam Spring Design Analysis

6.5

Cantilever Beam Spring Design Analysis

Several methods are used to analyze cantilever beam springs. Computer analysis, which requires the use of one of several available programs, is the latest method. One type of software program uses the classic cantilever beams equations. Another type employs graphical models of cantilever beams. A more sophisticated type of software uses a drawing program to create a scale model of the completed product, then a second stress analysis software program calculates the stress and deﬂection of the computer model using the product required end use service conditions. In addition, this software program also analyzes the dynamic characteristics of the thermoplastic material, such as the stress-strain curves at different temperatures, moisture absorption, and creep. The computer analysis methods are not discussed in this design handbook. Here, three spring analysis methods will be examined. Reference will be made to ﬂexible springs, in view of their practical importance for thermoplastic materials. Only springs loaded instantaneously and in normal environmental conditions will be considered.

6.5.1

Initial Modulus of Elasticity Cantilever Beam Analysis Method

This method is an approximation only for small strains, since it ignores certain properties of injection molding thermoplastic resins. The same equations as for metals are employed for calculating the initial modulus of elasticity. Because of the viscoelastic behavior of thermoplastic materials, the modulus of elasticity decreases with increased strain, an error in calculation is made which becomes greater as the strain increases. This method takes into account this error by ﬁxing a maximum limit (safety factor) or the “strain limit,” that has been established experimentally at 25%. For example, for acetal homopolymer cantilever springs the strain limit is 1.25% and the stress at 73 °F is 6,750 psi. The “strain limit” has a typical value for each material and is independent of the temperature. It has been proven experimentally that, within its “strain limit,” every thermoplastic material has practically complete elastic recovery. To calculate the deﬂection of a cantilever spring made of a thermoplastic material using this method of calculation, • Determine the deﬂection according to the elasticity equations applicable to metals, using the initial elastic modulus • Calculate the corresponding strain percentage. If the result is lower than the strain limit of the thermoplastic material in question, the calculation is within the permissible 25% error. Otherwise, a more accurate method of calculation must be used.

6.5.2

Stress-Strain Curve Cantilever Beam Analysis Method

This method of calculation also uses the classical equations developed for metals, but takes into consideration the secant modulus rather than the initial modulus. In this way, the typical shape of the stress-strain curve is taken into account, which was the main cause of the unreliability in this calculation, particularly for relatively high stresses.

381

382

6 Thermoplastic Molded Spring Design

6.5.3

Empirical Data Cantilever Spring Analysis Method

This method is based on experimental test data, obtained by stressing small samples injection molded of acetal homopolymer resins and measuring the actual deﬂection at the free end of the cantilever beam using various loads. With this method of calculation, the uncertainties caused by the stress distribution in the cross section areas are eliminated, which is very important in high stress conditions. Acetal Homopolymer Cantilever Spring Empirical Data The experimental test spring rates for acetal homopolymer cantilever springs are plotted in Figure 6-5. Deflection (δ) Length (L) at 73˚ F. and intermittent loading 2.000 Maximum stress 1.000 600

8 .0 0 0

4.0 00

800

L /h = 4 10 .00

9.0

2.000

0

400

00

200

6 .0 0 0

L /h = 6

100 80

L /h = 8

Spring rate, WM/δ, (psi)

60

8.88

L

40 L /h = 10 L /h = 12 20 L /h = 14 10

L /h = 16

8

L /h = 18

6

L /h = 20 L /h = 22

4

L /h = 24

W

b

L /h = 26 L /h = 28

2

δ h

Figure 6-5 Empirical data of acetal homopolymer cantilever spring (Courtesy: Du Pont)

1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

383

6.5 Cantilever Beam Spring Design Analysis

Example 6-1: Initial Modulus of Elasticity Analysis Method An acetal homopolymer cantilever spring is required to support a 2.0 lb load when deﬂected 0.600 in. The length of the cantilever spring is 1.50 in, the width is 0.375 in, and wall thickness is 0.0714 in. This example is used to illustrate the differences in the calculations among the three analysis methods for acetal homopolymer spring calculations. The initia modulus of elasticity for acetal homopolymer is: E = 410,000 psi : Calculate the moment of inertia (I): I = (b h3) / 12 = (0.375 × 0.07143) / 12 = 0.0000113 in4 : Calculate the spring load (W): W = (3 × δ × E × I) / L3 W = (3 × 0.60 × 410,000 × 0.0000113) / 1.503 = 2.487 lb : Calculate the section modulus (Z): Z = (b h2) / 6 = (0.375 × 0.07142) / 6 = 0.000318 in3 : Calculate the maximum stress (σ): σ = (W × L) / Z = (2.487 × 1.50) / 0.000318 = 11,731 psi : Calculate the spring strain (ε): ε = σ / E = 11,731 / 410,000 = 0.0286 = 2.86% The 2.86% value considerably exceeds the strain limit of 1.25% allowable for acetal homopolymer resins. Consequently, a more accurate calculation is necessary: 9 8

Calculate the load of the cantilever beam spring using the stress-strain curve analysis method and data from Example 6-1. From the stress-strain curve for acetal homopolymer shown in Figure 6-6 we see that with a strain of 1.25% the unit average tensile stress is 6,750 psi, which is a value below the yield point of the material (10,000 psi). Applying the cantilever beam Equation 6-2, the stress from Figure 6-6, L and Z from Example 6-1: W = (σ × Z) / L W = (6,750 × 0.000318) / 1.50 = 1.431 lb The force exerted by the spring is 1.431 lb rather than 2.487 lb.

122˚ F.

6.75

Tensile stress, (1.000 psi)

Example 6-2: Stress-Strain Curve Analysis Method

6

158˚ F.

5 212˚ F.

4 3 2 1 0 0

1.25

2

3

Strain, (%)

Figure 6-6 Stress-strain curve for acetal homopolymer (Courtesy: Du Pont)

Service temperature, (˚ F.)

73˚ F.

384

6 Thermoplastic Molded Spring Design

Example 6-3: Empirical Data Analysis Method An acetal homopolymer cantilever spring is required to support a 2.0 lb load when deﬂected 0.600 in. The length of the cantilever spring is 1.50 in and the width is 0.375 in. Calculate the cantilever spring wall thickness using the empirical data analysis method. L = 1.50 in W = 2.0 lb b = 0.375 in δ = 0.600 in : Calculate the modiﬁer load (WM): WM = W / b = 2.0 / 0.375 = 5.33 (lb/in)

: Calculate the spring rate (WM / δ): WM / δ = 5.33 / 0.600 = 8.88 (psi)

: Calculate the deﬂection/length ratio (δ / L): δ / L = 0.600 / 1.50 = 0.40

Find the intersection of the values for WM / δ = 8.88 (psi) and δ / L = 0.40 from the spring rate graph (Figure 6-5); these data provide the value for L / h. L / h = 20 : Calculate the cantilever spring wall thickness (h): h = L / 20 = 1.50 / 20 = 0.075 in Read off the maximum stress value from the empirical data graph from δ / L = 0.40 and the L / h = 20 (Figure 6-5): σMax = 7,250 psi A comparison of the three analysis methods used in Examples 6-1 through 6-3 shows that the initial modulus of elasticity analysis method contains an error about 62% too high in comparison with the empirical data analysis method in the estimating of the load exerted by the spring. The stress-strain curve analysis method error was 7% too low in comparison with the empirical data analysis method. The initial modulus of elasticity analysis method is more accurate, although not exact, largely because the data for the secant modulus was obtained from a tensile stress-strain curve, whereas the behavior of acetal homopolymer under bending stress is better. Note that within certain values of deﬂection the three methods give varying results, because of possible differences in the rate of application of the load and of crystallinity in the resin. It should also be noted that, within the normal design limits, calculation errors may even be lower than deviations of behavior due to injection molding factors (crystallinity) and the loading speed factor. Suitable testing is crucial, especially when carried out on a prototype injection molded spring under the same production conditions and tested under the actual conditions of use.

385

6.6 Thermoplastic Cantilever Spring Applications Deciding the actual load exerted by the cantilever spring is very important, in addition to the need to conduct tests to evaluate the useful life of the spring under actual working conditions (always being aware of strains lower than the limits for acetal homopolymer (1.25%) and for unreinforced nylon 6/6 at 50% RH (0.8%))

6.6

Thermoplastic Cantilever Spring Applications

With the introduction of engineering thermoplastic materials, many cantilever spring applications have been commercialized. Thermoplastic materials are different from metals; consequently, the cantilever spring designs for injection molding thermoplastic materials take advantage of the characteristics of these polymers, such as the ﬂexibility of the components that function as a spring and the multi-functional design capabilities of a complex component (gears, cams, etc.). This allows cost savings in the number of parts, fewer assembly operations, and less inventory. In addition, certain properties can be used to the advantage of thermoplastic materials, such as resistance to corrosion, chemicals and wear, low coefﬁcient of friction, low speciﬁc gravity, low noise, self-lubrication, electrical insulation, multi-functional design, lower manufacturing costs, etc.

Spring tongue Cam

Thermoplastic Spring Ratchet Applications The following illustrations show different injection molded thermoplastic spring ratchet and ﬂexing applications. Figure 6-7 shows a component from the mechanism of a cash register. The outer disc carries three ﬂexible tongues that act as springs allowing the central cam to rotate clockwise, but not in the counterclockwise direction. Figure 6-8 shows a garden spray; in the trigger a thermoplastic spring is arranged to set the rotation of an integral ratchet wheel with a cam. The trigger moves a lever that drives the upper piston that regulates how much water is discharged by the spray. Figure 6-9 shows a gear wheel with ﬂexible spring spokes molded in a single component. Figure 6-10 shows three spring ratchets and gear wheels with pointers for an instrument mechanism.

Wheel

Figure 6-7 Three spring ratchets for a cash register machine

Lever

Trigger

Spring

Figure 6-8 Garden sprayer with spring ratchet water regulator

Figure 6-9 Gear wheel with ﬂexible spring spokes

Figure 6-10 Three spring ratchets on gear wheel pointer

386

6 Thermoplastic Molded Spring Design Internal gear Spring

Rotor

Ratchet

Figure 6-11 Two spring ratchet gear systems

Figure 6-11 shows an electric meter assembly mechanism with an acetal homopolymer internal gear; two small ratchets made of metal are actuated by two springs made of acetal homopolymer. The springs are injection molded integrally with the central rotor. The springs are only 0.020 in thick for a working length of about 0.625 in. Figure 6-12 shows a portable tape labeling device. A spring ratchet mechanism is used for feeding the strip tape; the spring operating the ratchet was injection molded as an integral part of the trigger. Figure 6-13 shows an adding machine locating spring print gear. Figure 6-14 shows a gear and a spring ratchet with a tailpiece that exerts a counter force to eliminate backlash of the gear. The spring ratchet is preloaded on top of the gear teeth. The small and repeated deﬂections of the spring ratchet during the gear rotation cause negligible creep effects on material strength. Thermoplastic Locating Spring, Shock-Absorbing Applications

Gear Spring ratchet

Trigger

Figure 6-12 Labeling device spring ratchet tape feeder

Figure 6-15 shows a thermoplastic printing ring integrally injection molded with a locating spring. The rotating printing ring stops at the point mark position when the locating spring tongue enters or the spring backs inside a blank-marked slot of the printing ring. Figure 6-16 shows the mechanism for a locking cap of an automotive gasoline tank ﬁlling tube. When the cap is in the locked position the cap rotates idly, as the locating springs and the override cams avoid overtensioning the cap. Figure 6-17 shows a ﬁlm projector door handle locating spring that controls the rotation of the handle. Figure 6-18 shows a lock mechanism with shock absorbing spring, which functions in both horizontal directions. Locating spring tongue Locating spring

Locating spring

Central cam

Figure 6-13 Adding machine locating spring print-out gear

Figure 6-15 Printing ring locating spring

Figure 6-16 Locating springs, automobile gasoline tank cap

Figure 6-17 Film projector door handle locating springs

Figure 6-18 Lock mechanism with shock-absorbing spring

Spring ratchet Back lash spring

Ratchet gear

Figure 6-14 Gear backlash spring ratchet mechanism

Internal teeth

387

6.6 Thermoplastic Cantilever Spring Applications Figure 6-19 shows a cable stretcher system with a spring mechanism. Figure 6-20 shows a hose clamp adjustment mechanism for mounting various sizes of hoses. This type of hose clamp makes the installation and clamping of the hose simple, secure, fast, and safe. Figure 6-21 shows the body of an electrical switch, in which the pivoted on-off button is positioned by a locating spring. When the button is pushed down to either side, a direct indentation on the status of the switch is provided. Figure 6-22 shows an automotive roof light with snap springs to assemble the lens. Figure 6-23 shows an electrical slider switch with a leaf spring mechanism.

Figure 6-19 Cable stretcher system with spring mechanism

Thermoplastic Fixing Clip Spring Applications The following illustrations show injection molded thermoplastic ﬁxing clip spring applications. These ﬁxing clip spring applications are subjected to a permanent type of loading. They are designed for minimal stress loading of the ﬁxing clip springs, taking into account the eventual strength reduction over a long time (creep) and operating temperature conditions.

Locking spring Spring

Figure 6-24 shows a simple ﬁxing clip spring for a pen cap. Figure 6-25 shows a spring clip holder, where the ﬁxing clip spring is in the closed position. Figure 6-26 shows a ﬁlm projector locating compression spring, used to detect if the ﬁlm projector is loaded. Figure 6-27 shows a medical tubing ﬂow regulator for controlling the medication ﬂow rate.

Hose

Figure 6-20 Hose clamp system with ratchet locking springs

On-off bottom

Fixing clip spring Pivot

Figure 6-24 Common pen cap with retainer ﬁxing clip spring

Figure 6-25 Spring clip holder

Figure 6-21 Electrical switch locating position control spring

Clip spring flow position adjustment

Figure 6-22 Automotive roof light with snap springs

Slider switch

Leaf spring

Figure 6-26 Film projector locating compression spring

Figure 6-27 Medical tubing ﬂow regulator

Figure 6-23 Electrical slider switch with spring mechanism

388

6 Thermoplastic Molded Spring Design Flexible hinge

Pivot snap

Figure 6-28 Pivoted ﬂexible spring hinge clothespin

Flexible spring hinge

Thermoplastic Flexible Hinge Applications Combining the elasticity of some injection molding thermoplastics with their high resistance to repeated ﬂexing, it is possible to manufacture ingenious devices, as shown in the following illustrations. The function of a ﬂexible hinge is to remain open and strain-free when it is not in service; but when in operation, the ﬂexible hinge provides a differential spring force for clamping or returning to its original position. Figure 6-28 shows an integrally molded thermoplastic ﬂexible hinge clothespin. Figure 6-29 shows an encapsulated thermoplastic ﬂexible hinge staple remover. Figure 6-30 shows a thermoplastic ﬂexible hinge tap tool for recharging air conditioner refrigerant of cars. Figure 6-31 shows a ﬂexible hinge door catch device used to keep the door closed. Thermoplastic Torsional Springs

Metal teeth

Figure 6-29 Flexible spring hinge staple remover

Flexible hinge can refrigerant holder

Outlet

Locking guides

Figure 6-30 Flexible hinge tool for charging refrigerant of cars

Injection molded thermoplastic torsional springs are not commonly used in commercial applications. Thermoplastic torsional springs could provide new design opportunities in applications requiring intermittent load-bearing conditions. By the simple process of ﬁxing a thermoplastic ﬂat bar (Figure 6-32) at one end and applying torque at the other end, a spring action is produced by the torsional effect and the ﬂat bar’s tendency to return to its original free loading position. A thermoplastic torsional spring is subjected to the action of a bending moment, producing a normal stress and responds correctly to varying load conditions. The thickness of the ﬂat bar, its design, and the choice of a thermoplastic material with the ideal mechanical properties is critical. The properties required for some torsional spring applications are good fatigue resistance, spring like (hysteresis) characteristics, high rigidity, and good creep resistance. Figure 6-33 shows a heavy-duty torsional spring transmission shaft for an industrial grass mower tractor.

6.7

Thermoplastic Belleville Spring Washers

The Belleville spring washer takes its name from Julien Belleville, a French engineer who secured an English patent in the spring of 1866. His development was used largely in recoil mechanisms for artillery and buffer components used for railway applications.

Door catch

Flexible hinge door handle

Belleville spring washers, as shown in Figure 6-34, are convex round discs with a hole in the center having a saucer shape. Under load, they ﬂatten, resuming their original shape when the pressure is released. Mounted in series, back to back or nested on a rod, they provide a simple substitute for other types of springs.

Figure 6-31 Flexible hinge used to keep doors closed Free end torque

Fixed end

Figure 6-32 Flat bar torsional spring design

Figure 6-33 Grass mower tractor torsional spring transmission shaft (Courtesy: Du Pont)

Figure 6-34 Thermoplastic molded belleville spring washers

389

6.7 Thermoplastic Belleville Spring Washers Copolyester thermoplastic elastomers, acetal homopolymer, and nylon 6/6 are some materials employed in the fabrication of Belleville spring washers, used for applications requiring high spring loads and small deﬂections. Belleville spring washers injection molded from acetal homopolymer are especially desirable in demanding applications, requiring nonmagnetic and electrical insulation properties. Some potential end use applications are as a replacement for helical springs in switches and relays or return springs for push buttons on computer and control panels where nonmagnetic materials are essential. They also provide a toggle action for electrical and mechanical control levers. Belleville spring washers can be deﬂected from their original, convex shape to a concave shape, snapping inside out. They snap back to their original shape only when pressure is exerted in the opposite direction.

6.7.1

Acetal Homopolymer Belleville Spring Washer Analysis

Belleville spring washers injection molded from acetal homopolymer are used for testing the performance and comparing the experimental data to those obtained by using standard engineering equations developed for Belleville spring washers. Belleville spring washers of different sizes using the geometric model shown in Figure 6-35 were injection molded from acetal homopolymer. Each size of Belleville spring washer was tested using an Instron testing machine at a loading speed rate of 0.05 in per minute to develop the load and deﬂection empirical data. Standard equations were used to calculate loads and deﬂections for each spring size. Subsequently, the results were compared for both types of analysis methods. Belleville spring washer equation W =

E×δ δ    (H − δ)  H −  t + t 3   3   2  D   (1 − υ2 )   4 d ×  

(6-6)

Where: W E δ υ t H D d

= Compression load (lb) = Material modulus of elasticity (psi) = Belleville spring washer deﬂection (in) = Material Poisson’s ratio = Belleville spring washer wall thickness (in) = Belleville spring washer height (in) = Belleville spring washer outside diameter (in) = Belleville spring washer inside diameter (in)

Figures 6-36 through 6-44 are the plotted results for the compression load “W” (lb) vs. the deﬂection “δ” (in) obtained in both analysis methods. The solid lines in the spring graphs are the empirical data test results and the dashed lines are the calculated results generated by a computer analysis program.

W t H d D

Figure 6-35 Belleville spring washer geometric model

390

6 Thermoplastic Molded Spring Design

90

100

2.00 Dia. 1.25 Dia.

0.10

90

80 t = 0.10

Compression load "W," (lb.)

60 50 40 t = 0.07

30 20 10

80

80

70

70

60

t = 0.10

50 40 30 t = 0.07 20

0 0.05

0.10

0

0.05

50 40 30

Compression load "W," (lb.)

Compression load "W," (lb.)

60

0.20

180

80

t = 0.10

20

t = 0.03

Figure 6-39

0.10

2.00 Dia. 180

140

140

120

t = 0.10

100 80 t = 0.07

60 40

t = 0.03

0.20 0.5 Dia.

120

t = 0.10

100 80

t = 0.07

60 40 20

0

t = 0.03

0 0

0.05 0.10 0.15 0.20 0.25 0.30 Deflection "δ," (inch)

0.05 Deflection "δ," (inch)

200

2.00 Dia. 0.85 Dia.

160

20

10

0

0

0.10

160

t = 0.07

0

t = 0.07 20

Figure 6-38

200 2.00 Dia. 1.25 Dia.

70

30

Deflection "δ," (inch)

100 0.20

40

t = 0.03

Figure 6-37

90

t = 0.10

50

0

Deflection "δ," (inch)

Figure 6-36

0.5 Dia.

60

t = 0.03

0 0

0.10

10

10

t = 0.03

2.00 Dia.

90

Compression load "W," (lb.)

Compression load "W," (lb.)

70

100

2.00 Dia. 0.85 Dia.

0.10

Compression load "W," (lb.)

100

0.05 0.10 0.15 0.20 0.25 0.30

0

0.05 0.10 0.15 0.20 0.25 0.30

Deflection "δ," (inch)

Figure 6-40

Deflection "δ," (inch)

Figure 6-41 2.00 Dia.

2.00 Dia. 1.25 Dia.

360

280

240 200 t = 0.07

160 120

t = 0.03 80

Compression load "W," (lb.)

t = 0.10

0.30 320

0.30

320

280 Compression load "W," (lb.)

2.00 Dia. 0.85 Dia.

360

0.30

320

0.50 Dia.

400

280

t = 0.10

240

240 200 t = 0.07 160 120 t = 0.03

80

Compression load "W," (lb.)

400

200

t = 0.07 120 80

40

40

40

0

0

0

0

0.05 0.10 0.15 0.20 0.25 0.30 Deflection "δ," (inch)

Figure 6-42

0

0.05 0.10 0.15 0.20 0.25 0.30 Deflection "δ," (inch)

Figure 6-43

t = 0.09

160

t = 0.05 0

0.05 0.10 0.15 0.20 0.25 0.30 Deflection "δ," (inch)

Figure 6-44

391

6.7 Thermoplastic Belleville Spring Washers Reviewing the results found in Figures 6-36 to 6-44, using the empirical data and standard equation, it can be seen that the standard equation analysis method causes some type of error, especially for thick Belleville spring washers requiring large deﬂections. For example, for a 0.030 in wall thickness and a deﬂection of 0.100 in, the error is approx. 12%. For a 0.100 in wall thickness and a deﬂection of 0.100 in, the error is approx. 20%. Since the primary goal for using acetal homopolymer Belleville spring washers is to achieve high levels of spring rates and to be able to predict these values accurately, the standard equation, even when calculated by a computer program, does not produce accurate enough results. However, using the empirical data method and the graphs (Figures 6-36 to 6-44) will produce results that are accurate within a range of ±5%.

Example 6-4 W

Calculate the compression load (W) at a deﬂection (δ) of 0.060 in when the Belleville spring washer has the dimensions shown in Figure 6-45. To calculate the compression load (W), the Belleville spring washer requires a deﬂection of 0.060 in. Using the empirical data method, we select the spring graph in Figure 6-39 that has the same dimension. The compression load (W) value of 24.00 lb is found for a deﬂection (δ) of 0.060 in and the wall thickness curve (t) of 0.03 in.

0.030 inches 0.20 inches 1.25 inches 2.00 inches

Figure 6-45 Belleville spring washer

Example 6-5 Calculate the Belleville spring washer height (H) that will produce a compression load (W) of 26.00 lb at a deﬂection (δ) of 0.050 in, when the Belleville spring washer has an increasing spring force with deﬂection and the dimensions show in Figure 6-46. Figures 6-36 to 6-44 were developed using 0.10, 0.20, and 0.30 in heights; however, we need to calculate a new height different from those used in the graphs. Reviewing the Belleville spring graphs, Figures 6-37 and 6-43 can be used to calculate the height: Spring graph Figure 6-37 D d t H δ W

= 2.00 in outside diameter = 0.85 in inside diameter = 0.03 in wall thickness = 0.10 in spring height = 0.05 in deﬂection = 5.00 lb load

Spring graph Figure 6-43 D d t H δ W

= 2.00 in outside diameter = 0.85 in inside Diameter = 0.03 in wall thickness = 0.30 in spring height = 0.05 in deﬂection = 57.00 lb load

From Figure 6-37, we obtain a load of W = 5.00 lb and the height of H = 0.10 in and from Figure 6-43 we obtain a load of W = 57.00 lb and the height of H = 0.30 in. Interpolating both loads, we obtain the desired compression load (W) = (5.00 + 57.00) / 2 = 31.00 lb By interpolating both spring heights, we obtain the desired spring height (H) = (0.10 + 0.30) / 2 = 0.20 in.

W 0.030 inches H

0.85 inches 2.00 inches

Figure 6-46 Belleville spring washer

392

6 Thermoplastic Molded Spring Design

Load Ratio, WA/WC

1.20

1.10

1.00 0.05

0.50 5.00 Axial loading rate, (inch per minute)

50.00

Figure 6-47 Load ratio WA / WC vs. axial loading rate

6.7.2

Belleville Spring Washer Loading Rate

A characteristic of most thermoplastic materials is their sensitivity to the loading rates; this effect should be considered in any Belleville spring washer calculations. Figure 6-47 is a graph of the axial loading rate versus the ratio of the true load (WA) to the calculated load (WC). An axial loading rate of 0.05 in/min was used to develop the values for load and deﬂection shown in the graph. The appropriate load ratios can be found for speed rates between 0.005 to 50 in/min for converting the load to a realistic value. For example, if the axial loading rate is 5.0 in/min, the ratio WA / WC is 1.13, showing a realistic load 13% higher than the calculated load value (WC).

6.7.3

Belleville Spring Washer Long-Term Loading Characteristics

Acetal homopolymer and copolyester thermoplastic elastomer Belleville spring washers have an excellent recovery rate (hysteresis) from the applied compression loads. However, if a thermoplastic Belleville spring washer is constantly loaded or stressed for a long period at higher temperatures, the thermoplastic materials will be subjected to the effects of creep or loss of mechanical properties. Therefore, the load versus deﬂection characteristics of thermoplastic Belleville spring washers will be reduced, causing inferior performances of the product with lower than expected values. Acetal homopolymer and copolyester thermoplastic elastomer Belleville spring washers are not recommended for any application that requires the energy to be stored or continuous loading expecting a constant spring reaction force, and these materials will not perform under these requirements. The use of thermoplastic Belleville spring washers should be limited only to intermittent loads or deﬂection applications.

393

7

Thermoplastic Pressure Vessel Design

7.1

Thermoplastic Thin-Walled Pressure Vessels

The term thin-walled or thin-shelled pressure vessel describes a hollow cylinder in which the circumferential stress (frequently called hoop stress) in the wall is assumed to be constant throughout the thickness of the wall when the cylinder is subjected to an internal or external ﬂuid pressure. Thermoplastic materials have been used for the fabrication of many pressure vessel devices, such as toilet ﬂush valves, spray paint containers, butane lighters, irrigation sprayers and valves, brake master cylinders, radiator end cores, garden hoses, tubing, end connections, pumps, and so forth. Figures 7-1 and 7-2 illustrate some of these pressure vessel applications.

Water spray gun

Figure 7-1 Thermoplastics pressure vessel applications (Courtesy: Du Pont)

Figure 7-2 Thermoplastic pressure vessel applications

394

7 Thermoplastic Pressure Vessel Design

7.2

The circumferential stress (a) in a thin-walled cylinder subjected to an internal pressure (P) per unit area is found by applying an equation of equilibrium to the forces acting on the half cylinder shown in Figure 7-3. The length is uniform, wall thickness is (t), and inside radius is (r).

t 2r x P r P

da

Thin-Walled Cylinder Basic Principles

da

P × 2 r = 2 ∫ σ da

Figure 7-3 Thin-walled cylinder mathematical model

(7-1)

But under the assumed conditions, ∫ σ da = a σ = t σ . Therefore, Barlow’s Equation is:

t

σ =

P×r t

(7-2)

r

Spherical Closed End Thin-Walled Pressure Vessels P

r

To calculate the stress of a cylindrical pressure vessel with a spherical base, under uniform internal pressure, using Figure 7-4 as a model, the maximum stress equation should be applied:

Figure 7-4 Spherical closed end thin-walled pressure vessel

σ Max. =

P×r 2×t

(7-3)

Flat Closed End Thin-Walled Pressure Vessel

r

To ﬁnd the maximum stress of a cylindrical pressure vessel with a circular ﬂat bottom base, under uniform internal pressure and using Figure 7-5 as a model, the following equations should be applied:

t P

Center deﬂection:

 3 × P × r4  2 δ=  (1 − υ )  16 × E × t 3 

Maximum moment:

M Max. =

Maximum stress:

σ Max. =

P

Figure 7-5 Flat closed end thin-walled pressure vessel

P × r2 8

(7-4)

(7-5)

6 × M 0.75 × P × r 2 = t2 t2

(7-6)

Example 7-1 The shank/riser toilet ﬂush valve shown in Figure 7-6 needs to withstand a 2,000 psi burst pressure and 120 psi continuous internal pressure for 10 years. The burst pressure would be the controlling factor for the design, rather than the continuous pressure. The material selected for this application is acetal homopolymer with a tensile strength of 10,000 psi. Calculate the wall thickness of the shank/riser by using Barlow’s equation (Equation 7-2). Barlow’s Equation σ =

P×r t

or t =

P × r 2,000 × r = = 0.20 r σ 10,000

The stress for the shank/riser wall thickness at 120 psi pressure over ten years is calculated by using the isochronous creep stress/time long-term pipe test

7.2 Thin- Walled Cylinder Basic Principles

data shown in Figure 7-7. For 10 ycars the tensile stress is 1,750 psi. The wall thickness can be calculated as t, = (120 x r) 1 1,750 = 0.068 r. The burst pressure wall thickness t , = 0.20 r is the control bctor for dimensioning, because the wall: thickness for burst pressure requires 0.20 10.068 = 2.94 times the wall thickness calculated to retain the 110 psi internal pressure

over 10 ycars.

The shanklriser outside diameter at the base (left is theside) root diameter of the threads, and its wall thickness is t = 0.20 x 0.453 = 0.0906 in. Because its t = 0.20 the toy (right side) outside diameter is smaller, wall thickness x 0.375 = 0.075 in, without considering the reinforcement by the caused perpendicular wall of the toilet valve housing. twoThese shanklriser inside diameters form a tapered wall, which not only provides the efficientmost design, but aIso helps during the part ejection from the long core of the mold. To improve the molding process efficiency,the long core requires the surface to be hardened to 60 R,, the surface to be coated with a low coefficient of friction mold release and polished in the longitudinal direction (a RMS). The core requires alsoan independent water cooling system. The shanklriser is to he gated at the flange (thicker width wall sectinn equal 0.156 to i n ) using an insulated runnerless mold with three cavities.

Figure?-6 Acetal homopolymer shank/riser toilet flush valve

Internal pressure at 73" f.

10.000

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100

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1 Year

10 Years

Time, [hours)

Figure 7-7 Fsochronous acetal homopolymer pipe burst data (Courtesy:Du Pont)

396

7 Thermoplastic Pressure Vessel Design

7.3

Thick-Walled Pressure Vessels

P x 2r1

P

P da

da

Barlow’s Equation applies only to thin-walled cylinders; this fact will be evident from a consideration of the equilibrium of the forces acting on the thinwalled cylinder shown in Figure 7-8. In addition, a satisfactory solution for the thick-walled cylinder problem requires the determination not only of the circumferential stress at any point in the cylinder, but of both of the principal stresses whose vectors lie in the plane of the paper, namely, the circumferential or tangential stress σt and the radial stress σr as shown in Figure 7-9.

r1 r2

Figure 7-8 Thick-walled cylinder mathematical model

r

+d

r

7.3.1

t

d

Lame’s Equation for Thick-Walled Cylinders

Lame’s equation is used to calculate the maximum tangential and radial stresses of thick-walled cylinders subjected to internal and external pressures.

r

P2

For the pressure vessels with a relatively large wall thickness, the mathematical model shown in Figure 7-8 should be applied. If the variation in the stress from the inner surface to the outer surface is relatively large, the value of the stress found from Barlow’s Equation 7-2 is not a satisfactory measure of the signiﬁcant circumferential stress in thick-walled pressure vessels.

P1

Figure 7-9 represents a relatively long open-ended thick-walled cylinder subjected to internal and external ﬂuid pressures P1 and P2, respectively. We will use Figure 7-9 to ﬁnd the circumferential stress σt and the radial stress σr at a point at any distance from the central axis of the. From the conditions of symmetry, it is known that there is no shearing stress on the planes on which σt and σr act and therefore they are principal stresses.

r1 r2

Figure 7-9 Radial and tangential stresses caused by in/out pressures

d P1 P2 r

da

( r + d r ) da

Finding both σt and σr requires that sections be passed through the body so that the portion of the body isolated by the sections will be acted on by forces that involve the two stresses. Such a portion is obtained by ﬁrst passing two concentric sections through the body and thus isolating a thick-walled cylinder; this element with the forces acting on it is shown in Figure 7-10. A diametral plane is then passed through the element, isolating one half the element that may be expressed by the two stresses (one at each side of the half cylinder), as shown in Figure 7-11. By applying one of the equations of equilibrium to the forces in Figure 7-11 we ﬁnd that the algebraic sum of the vertical components of the forces is equal to zero and the following equation is obtained: σ t d ρ = 2 σ r d ρ − 2 ρ d σ r − 2 d ρ dσ r

Figure 7-10

(7-7)

The term 2 dρ dσr is negligibly small. The tensile stress becomes: (

r

r

) 2( + d ) r

+d

(7-8)

If the stresses σr and σt in Figure 7-11 are assumed to be positive, that is, if both stresses are assumed to be tensile stresses, the previous equation is:

2

d

d

Figure 7-11

σ t = σ r + ρ (dσ r /dρ)

(7-9)

P2

P1

t

σ t = − σ r − ρ (dσ r /dρ)

t

d

A rational assumption concerning the strains in a thick-walled cylinder is that longitudinal strains of the ﬁbers are equal. This means that transverse (parallel) sections that are plane before the ﬂuid pressures P1 and P2 are applied remain plane and parallel after the pressures are applied. This will be true at least for a

397

7.3 Thick-Walled Pressure Vessels cylinder with open ends, and it will also be nearly true for a closed cylinder at sections well removed from the ends of the cylinder. The relation between the longitudinal strain εl of any longitudinal ﬁber and the stresses acting on the ﬁber in an open ended thick-walled cylinder is: ε1 = υ (σ r / E ) − υ (σ t / E )

(7-10)

and according to the above assumption, εl is constant. In addition, Poisson’s ratio ν and the modulus of elasticity E are constants of the material, resulting in: σ t − σ r = Constant = 2 κ The constant is denoted by 2 κ for convenience. The previous two equations give two relations between σr and σt. From these two equations we obtain: 2 κ = −2 σ r − ρ (dσ r /dρ)

(7-11)

But the right hand side of this equation, when multiplied by ρ, becomes the derivative, with respect to ρ, of –(ρ2 σr), and therefore the equation may be written: d( ρ2 σ r )/dρ = −2 κ ρ

(7-12)

The integration of this equation gives: ρ2 σ r = − κ ρ2 + β

(7-13)

where β is a constant of integration. Therefore: σ r = ( β / ρ2 ) − κ

(7-14)

and from the equation σt – σr = Constant = 2 κ σ t = ( β / ρ2 ) + κ

(7-15)

The values of the constants β and κ are found by substituting the values σr and ρ that were obtained from the physical conditions or assumptions stated. For example, assuming the cylinder to be subjected to both internal and external pressures P1 and P2, respectively, we observe that σr = P1 where ρ = r1 and σt = P2 where ρ = r2 and from the Equation 7-14, σ r = ( β / ρ2 ) − κ we obtain: P1 = − κ + ( β / r12 ) and P2 = − κ + ( β / r22 ) from which κ and β are: κ=

P1 r12 − P2 r22

β=

r12 × r22 (P1 − P2 ) r22 − r12

r22 − r12

(7-16)

(7-17)

398

7 Thermoplastic Pressure Vessel Design The substitution of these values of κ and β in Equations 7-14 and 7-15: σt =

P1 r12 − P2 r22 + (r12 × r22 / ρ2 ) (P1 − P2 ) r22 − r12

(7-18)

σr =

P2 r22 − P1 r12 + (r12 × r22 / ρ2 ) (P1 − P2 ) r22 − r12

(7-19)

It is evident from Equation 7-18 that the maximum value of σt occurs at the inner surface where ρ has its minimum value r1. The maximum value of σr will always be the larger of the two pressures P1 and P2.

7.3.2

Maximum Stresses with Internal and External Pressures

By setting ρ = r1 in Equations 7-18 and 7-19, the maximum stresses (at the inner surface) are: Maximum tensile stress (Lame’s Equation) Maximum radial stress

7.3.3 P2 = 0

r

Max.

t

= P1

Figure 7-12 Maximum stress, internal pressure, mathematical model

σt =

P1 (r12 + r22 ) − 2 P2 r22 r22 − r12

(7-20)

σ r = P1 if P1 > P2

Maximum Stresses for Internal Pressure Only

The maximum radial stress value σr is equal to P1; these maximum stresses are shown in Figure 7-12. If the internal pressure is P1 and the external pressure is zero (P2 = 0), the maximum stress equations are reduced to: σt =

 P1 × r12  r22 + 1 2 2  2 r2 − r1  ρ 

(7-21)

σr =

 P1 × r12  r22 − 1 2 2  2 r2 − r1  ρ 

(7-22)

These equations show that the maximum values of σt occur at the inner surface, when ρ = r1. The new equation is: Maximum tensile stress

σt =

P1 (r12 + r22 ) r22 − r12

(7-23)

Table 7-1 provides more equations for calculating different pressure vessel types of loading.

399

7.3 Thick-Walled Pressure Vessels Table 7-1 Cylindrical Pressure Vessel Equations

Vessels geometry Thin-walled cylinder y P2

T

Type of loading

Cylindrical pressure vessel equations

Uniform internal or external pressure

σ1 =

σ2 =

P1 × R2 T

   σY T   P2 = 2 R2  σ Y  R2   1 + 4     E T  

2 1

P1 × R2 2×T

P1 R2

∆R2 =

R2 (σ 2 − υ × σ 1 ) E

P1 = 2 × σ1

R2 − R1 R2 + R1

P2

Thin-walled cylinder with spherical bottom base

Uniform internal pressure

Maximum stress at middle of spherical base Max. σ =

T

P1 × R1 2×T

P1 R1

Uniform internal pressure

Thin-walled cylinder with circular ﬂat bottom base

Maximum stress at middle of circular ﬂat base  3 × P1 × R14  2 δ=  (1 − υ )  16 × E × T 3 

T

M Max. =

P1 R1

Max. σ =

Thick-walled cylinder

Uniform internal pressure P2 = 0

P2

P1 × R12 8 6 × M 0.75 × P1 × R12 = T2 T2

R22 + R12 R22 − R12

σ1 = 0

Max. σ 2 = P1

∆R1 = P1

R1 E

σ1 = 0

Max. σ 2 = P2

∆R1 = P2

R1  2 × R12  E  R22 − R12 

Max. σ 3 = P1

R22 R22 − R12

P1 2 3 1

R2 P2

R1

Uniform external pressure P1 = 0

R2 = External radius, R1 = Internal radius, T = Thickness, υ = Poisson ratio, δ = Deﬂection, P2 = External pressure, P1 = Internal pressure, E = Flexural modulus, σ = Stress, σY = Yield stress, M = Moment of force

 R22 + R12  + υ  2 2  R2 − R1 

∆R2 = P1

2 × R22 R22 − R12

R2  2 × R12  E  R22 − R12 

Max. σ 3 =

∆R2 = P2

R2 E

Max. σ 2 2

 R22 + R12  − υ  2 2  R2 − R1 

400

7 Thermoplastic Pressure Vessel Design

7.4

Designing Cylinders for Cost Reduction

The typical design of spherical closed end pressure vessels uses smooth, thickwalled cylinders where the hoop stress is double the axial stress. The use of thinner walls and circumferential ribs to reinforce the vessel in the hoop direction improves the process efﬁciency, lowers part weight, cycle time, and manufacturing costs. A design comparison is shown in Figures 7-13 and 7-14. Figure 7-13 shows a typical pressure vessel design with 0.250 in wall thickness. The recommended ribbed design is shown in Figures 7-14 and 7-15. The wall thickness was reduced to 0.125 in and ribs were added to provide the same strength of the thick-walled cylinder, in both directions, hoop and axial. Fillet radius and draft angles on the ribs reduce stress concentration and simplify part ejection from the mold. The stresses caused by the spherical closed end cylinder wall were checked using standard pressure vessel equations. The stress level was negligible for both designs.

Figure 7-13 Thick wall cylinder (Courtesy: Du Pont)

The estimated cost savings of the ribbed design is 30%. Because of the reduction in the molding cycle from 85 s for the 0.250 in, thick-walled design to 45 s for the recommended 0.125 in walls and circumferential ribbed design.

7.5

Thermoplastic Pressure Vessels Design Guidelines

0.125

When good design, proper resin selection, mold design, quality injection molding, and testing are employed, thermoplastic pressure vessels will provide satisfactory and safe service performance to the end users. Because thermoplastics have nonlinear stress-strain relationships over a large range of strains, the usual analytical techniques applied to pressure vessels may not be accurate enough to predict failures. The design parameters for a thermoplastic pressure vessel should be very conservative and the injection molding process should be set up to comply with the “A-1 Quality Control” requirements.

Figure 7-15 Recommended design, ribbed cross section detail

7.5.1

Figure 7-14 Recommended design (Courtesy: Du Pont)

0.062 R.

0.250 Draft angle 0.375

Preliminary Pressure Vessel Design

A cylinder is considered thin-walled when the ratio between the wall thickness to inside radius is 0.50 or less; in this case, Barlow’s Equation (Eq. 7-2) should be used. Lame’s Equation (Eq. 7-20) should be used for thick-walled cylinders having a ratio greater than 0.50. Other sound engineering techniques are available to the designer (e.g., ﬁnite element analysis, more sophisticated equations, etc.) to determine the wall thicknesses and dimensions of the loadbearing members. When designing a pressure vessel cylinder with a snap-ﬁt end cap, the internal pressure will deform the wall of the vessel more than its cap. This reduces snap-ﬁt tightness and causes leakage. It is recommended to redesign the exterior snapﬁt cap and add an “O” ring to eliminate the leakage. Figure 7-16 shows a poor product design, operational problems, and design recommendations. When designing a pressurized cylinder with a bolted end cap and a top-seated “O” ring, a very high load is required to compress the “O” ring axially, consequently deformation and creep of the seal ﬂanges occur. This effect worsens when the distance between the “O” ring and the bolts is increased. To compensate for this

401

7.5 Thermoplastic Pressure Vessels Design Guidelines effect, redesign the end cap and the cylinder wall to move the “O” ring to seal by compressing radially. Additional beneﬁts are gained by reducing creep. Stiffening the ﬂange with ribs or a metal ring under the bolts can help. Figure 7-17 shows a poor design, operational problems, and design recommendations. When designing a pressurized cylinder with self-tapping screw end caps and a top axial seated “O” ring, the end cap requires snap-ﬁt ﬁngers with the “O” ring moved to seal radially. With the low internal pressure required for this valve application, the use of snap-ﬁts can make the assembly faster and more economical. Figure 7-18 shows a poor design while Figure 7-19 shows the design recommendations. Carefully study the thermoplastic material properties for the cylinder.

Valve assembly top view Axial "O" ring

Specify the sensitive areas of the pressure vessel, such as type, size, number, and location of the gate, sharp corners, weld lines, and ribs. The design pressure of the cylinder should be lower than 150 psi, or 15% of the maximum required bursting pressure. Creep stress of the resin must be used based on best available data.

Self-tap screw Valve cross section view

Pressure expands wall (leakage)

Pressure & "O" ring eliminates leakage

Figure 7-18 Poor valve design, “O” ring with self-tapping screw end cap (Courtesy: Du Pont)

"O" ring

Internal pressure

Poor design

Internal pressure

Operational problems

Internal pressure

Design recommendations

Figure 7-16 Pressure vessel with snap-ﬁt end cap design (Courtesy: Du Pont) Valve assembly top view Valve end cap with six lock springs Best seal, radial "O" ring in compression Axial "O" ring

Axial bolts compression forces apart joining walls

Radial "O" ring

Radial "O" ring

Valve base with six lock pockets Poor design

Operational problems

Figure 7-17 Pressure vessel with bolted end cap design

Valve cross section view Design recommendation

Figure 7-19 Valve with six snap-ﬁt end caps, design recommendations

402

7 Thermoplastic Pressure Vessel Design

7.6

Testing Prototype Thermoplastic Pressure Vessels

Build a cylinder prototype tool and mold several prototype samples. For each pressure vessel to be tested, use the same thermal and moisture environment of the product in its end use operation. All cycling and burst tests will be conducted in this environment. The designer must decide whether a cycling pressure test is needed. If the pressure vessel end use involves frequent pressurization and depressurization, a cycling fatigue test is recommended. The cylinder pressure should be tested starting from atmospheric to design pressure and back to atmospheric 100,000 times or less, if a lower service life is acceptable. If a cycling test was done on the prototype pressure vessel, a burst test must be done with the same specimen used in the cycling test. Otherwise, use any good sample conditioned to resemble the operating environment for the burst test. Burst pressure and the mode of failure should be carefully recorded.

7.6.1

Redesign and Retesting the Pressure Vessels

If needed, modify the pressure vessel design according to the outcome of the previous tests. Conduct a burst test with one of the modiﬁed pressure vessels to conﬁrm the improvements made by the redesign modiﬁcations. Build a production tool and mold commercial quality cylinders. Retest the pressure vessel periodically to conﬁrm if the quality of the injection molding process meets the quality control requirements established for the production pressure vessels.

7.7

Pressure Vessel Regulations

Pressure vessels are regulated by industry codes to establish design safety guidelines for dimensioning and testing vessels of various materials and end use applications. One of the most prominent groups that regulate pressure vessel design is the American Society of Mechanical Engineers (ASME). The ASME Boiler and Pressure Vessel Codes have become an American National Standard (accepted by the American National Standards Institute) and are mandatory by law in several states. Injection molded thermoplastic pressure vessels do not always fall under the jurisdiction of the ASME code because of the small size of the pressure vessels and the type of end use applications. However, the National Sanitation Foundation is very active in regulating the design, testing, and manufacture of injection molded thermoplastic pressure vessels.

7.7 Pressure Vessel Regulations

7.7.1

ASME Pressure Vessel Code

This regulation is applicable for materials of flexural modulus as low as 1.0 × 106 psi. Some thermoplastic materials meet this criterion. Creep of thermoplastic materials is not considered by the code, but must be taken into account in designing thermoplastic pressure vessels. Tensile Strength The materials considered by this code are the materials that have tensile strengths from 12,000 psi to 25,000 psi. Design Pressure Sets the maximum design pressure lower than 150 psi, or 15% of the bursting pressure. Design Temperature Temperature is set at 150 °F and the code requires the burst pressure test to be done at this temperature. Operating Pressure Should be less than, or equal to the design pressure. Bursting Pressure Bursting pressure is the hydrostatic pressure at which a prototype pressure vessel bursts. Loadings Several types of pressure loadings are considered; the most important being internal and external pressure. The other types are mechanical impact loads, reactions of supporting lugs, rings, and so forth. Stress Caused by Combined Loads Stresses are analyzed using the membrane stresses produced by bending and shearing loadings. The maximum membrane stress during normal operation of the pressure vessel must not exceed 15% of the maximum membrane stress at the bursting pressure. Proof of Design Adequacy A prototype pressure vessel must be subjected to 100,000 pressure cycles, ranging from atmospheric to the design pressure. After this test, the same pressure vessel should burst at a pressure no less than six times the speciﬁed maximum design pressure. The test ﬂuid should have a minimum temperature of 150 °F. Pressure Relief Devices It is require that all pressure vessels be provided with protection against over pressures.

403

404

7 Thermoplastic Pressure Vessel Design Three types of pressure relief devices are acceptable: • Direct spring loaded safety relief valves • Pilot operated valves • Rupture disks Set Pressures A single pressure relief device will be set at the design pressure. Set Pressure Tolerances For safety and safety relief valves: • ±2 psi when the operating pressure is lower than 70 psi • ±3 psi when the operating pressure is higher than 70 psi • Rupture discs ±5 psi for all pressures Permissible Over Pressures Single relief devices should withstand 110% of the design pressure. Testing Requirements Apply cyclic pressure and adhere to burst test conditions: • Test ﬂuids should be water or other liquids • Temperature of test ﬂuid should be 150 °F • Cycle pressure from atmospheric to design pressure and back, 100,000 times • Following the cyclic test, the pressure vessel should burst. The minimum burst pressures to be six times the design pressure.

405

8

Thermoplastic Assembly Methods

8.1

Introduction

Injection molded thermoplastic components lend themselves to a number of assembly and ﬁnishing methods. The best suited method for a particular design depends on performance, cost, and number of units to be produced. In most cases, design considerations should be examined early in the product development. In thermoplastic product engineering, designing a one-piece item for molding is the ideal situation, because it excludes any assembly operation. However, mechanical limitations and other considerations often make it necessary to join thermoplastic components either to each other or to other thermoplastic or metal components to complete the assembly. In such instances, the joining process can be an efﬁcient production method, if a few precautions are taken and established procedures are followed. Various methods of joining can be used successfully with most injection molding thermoplastic materials. These include hot plate welding, cold heading, press-ﬁtting, snap-ﬁtting, mechanical fastening, solvent bonding, vibration welding, spin welding, electromagnetic welding, and ultrasonic welding. Choosing the best method requires a fundamental knowledge of good joint design and a thorough understanding of the purpose of the joint, the type of thermoplastic material, the geometry and nature of the components, the type of load involved, and the properties of the particular thermoplastic material to be used. Because of these engineering complexities, careful product planning to eliminate needless joints is essential. The best system of joining two components is never as good as a well designed component molded in one piece.

8.2

Cold Heading Method

Cold heading is an assembly method useful in forming strong mechanical joints between either similar or dissimilar materials of various geometrical shapes at low cost. This fastening method was one of the ﬁrst assembly processes developed for the plastics industry. It is similar to the riveting operations used for fastening metal sheets onto a metal structure; the metal rivet head is formed at the end of the stud by applying a compressive load. Staking or cold heading is the process of deforming a thermoplastic stud that has been molded in a component to join mechanically with a plate-like item, forming a strong and permanent joint. The process is performed with a contoured tool that transfers energy to the thermoplastic, producing a melt phase and exerting pressure to reform the thermoplastic. Tight assemblies require that pressure be maintained until the thermoplastic becomes solid in the new holding contour. Selection of the optimum staking method for a given application is based on considerations such as joint strength, reliability, resin grade, assembly time, tool wear, assembly, and equipment costs.

406

8 Thermoplastic Assembly Methods Most common types of thermoplastic unreinforced resins can be cold-headed, because the physical properties of these resins are compatible with this assembly method. However, this technique is being replaced by the ultrasonic staking process method, which provides several advantages, such as the contact tip of the tool staying cool during processing, forming a clean head without over-stressing the joint, and producing a stronger joint with less elastic recovery and a tighter head. The ultrasonic staking method allows the use of thermoplastic reinforced resins, thus expanding the design scope of thermoplastic applications. Figures 8-1 and 8-2 show a pump impeller made of acetal homopolymer that was assembled by cold-heading all studs simultaneously in a punch press. Figure 8-1 Pump propeller (Courtesy: Du Pont)

8.2.1

Cold Heading Procedure and Equipment

Cold heading may be accomplished by applying a uniform load at the end of a thermoplastic stud, while holding and containing the thermoplastic unit, forming a rivet-type head at the end of the shaft. The magnitude of the compressive load is such that it exceeds the yield strength of the thermoplastic material causing it to permanently deform into a rivet-like head. Equipment requirements for cold heading may range from a simple arbor press and hand vice to a punch press and automatic clamping ﬁxture for multiple heading operations. Figures 8-3 and 8-4 show the basic principles for a cold heading tool with the thermoplastic stud to be joined in position under the tool. As the tool is brought into contact with the thermoplastic stud, a spring loaded sleeve preloads the area around the protruding stud to assure a tight ﬁt between the ﬂat metal plate and the thermoplastic rivets. The heading portion of the tool then cold-heads the end of the stud, forming a strong, permanent, mechanical joint.

Figure 8-2 Cold heading details (Courtesy: Du Pont)

There are a number of staking proﬁles available to provide a variety of appearances and strengths. The standard proﬁle produces a head form having twice the diameter of the original stud. Other proﬁles, which provide appearance enhancements, generally provide less holding strength than the standard proﬁle. The staking horn tip should be contoured to contain the molten thermoplastic and assure a tight assembly. The staking ultrasonic horn tips used for glass or mineral reinforced thermoplastic materials, hardened steel, or carbide tips are suggested to minimize surface erosion and extend tool life.

1.1 t Preloaded spring Pilot sleeve 0.1 t

The basic principles of cold heading can be adapted to many speciﬁc applications. The following design considerations are presented as a guide to the applicability of cold heading as an assembly method.

0.7 t 0.2 t

1.4 t 2.5 t

Tool

1.5 t Metal plate

t

0.7 t 1.5 t

t

Plastic rivet

Metal plate

r Thermoplastic

Position of tool Figure 8-3 Cold heading equipment prior to assembly

Stroke the tool

Finished head

Figure 8-4 Cold heading expanded rivet method, processing steps

407

8.2 Cold Heading Method Spring

Metal plates

Pilot tool

Plastic rivet

Sleeve

Position of tool

Stroke the tool

Finished head

Figure 8-5 Cold heading flanged rivet method, processing steps

Heading tool Metal plates

Plastic stud

Knurl or flat surface

Position of tool

Recovery

Stroke the tool

Finished head

Figure 8-6 Cold heading flat rivet method, processing steps

Shaft Geometry When a compressive load is applied to a stud that exceeds the yield strength of the thermoplastic material, permanent deformation takes place at the minimum cross section of the thermoplastic stud that is not contained. It is therefore possible to accurately locate the formation of a cold head either by variation of the heading tool proﬁle or by clamp location. The heading tool shaft can be tapered, bored out, or shaped like tubing to reduce the cross section and concentrate the load at a point where heading is desired. If the heading tool shaft is contained in a jig or by another part, it is not necessary to reduce the cross section to form a head. However, a tighter joint is formed when the cross section is reduced. The joint is tighter because less force is needed to form the head. With less force necessary, the heading tool support is subjected to less elastic deformation that is regained after the pressure is released and less elastic recovery provides a tighter head. The length of the thermoplastic stud also affects the degree of recovery from the deformation in cold heading. A short heading tool shaft is deformed less in the axial direction, which results in a tighter joint. The joint can be made tighter by incorporating a spring effect in the design. The stud length necessary to form a satisfactory head varies with the diameter and also with the material. To determine the length, the ﬁrst step is to estimate the ﬁnished head dimensions. Knowing these dimensions and the diameter of the thermoplastic stud, the length can be easily calculated by using equivalent volumes. The length to diameter ratio of 2 : 1 is suggested for the stud geometry wherever applicable.

408

8 Thermoplastic Assembly Methods Rate and Type of Loading The rate of loading for cold heading is not critical but should not be an impacttype load. The impeller shown in Figures 8-1 and 8-2 was headed at room temperature in a punch press at 20 in/min using a spring loaded ﬁxture. Relaxation of Stud and Head The tendency of a material to recover to its original shape after deformation is dependent on its elongation properties. The deformation under the head is recovered when the load that formed the cold head is released, thus forming a gap (recovery) in the joint. With thermoplastic materials, stress relaxation is encouraged by exposure to elevated temperatures. Materials with high elongation tend to regain their original shape with an increase in temperature. The head itself tends to regain its original shape and does not affect the tightness of the joint to the same degree, because the shaft length below the head does not change appreciably. Force Required to Form Head The compressive stress necessary to reach the yield point at room temperature and 50% relative humidity using unreinforced nylon 6/6 is 12,000 psi. The compressive yield stress for acetal homopolymer is 18,000 psi. The total force increases as the cold head is formed, because the shaft cross section increases. The load required to head both materials at 200 °F is approx. 50% of the load at 73 °F. Heading at higher temperatures is advantageous because less force is needed and tighter ﬁts can be obtained. In some applications it may be possible to perform the heading operation at the molding machine before the injection molded component cools. Strength of Cold Headed Joints Tensile pull strengths of heads formed by cold heading are approximately equal to 50% of the shear strength of the material under consideration.

8.3

Electro Fusion Fitting System

The world wide use of polybutylene for thermic-mechanic-chemical high stressed pipe systems is increasing. It’s most important applications are for ﬂoor heating, hot water conduits, district heating, piping for industrial applications and thermal water conduits. The welding system was developed to create an easy to install connection system for polybutylene pipes, similar to the electro fusion systems for HDPE and MDPE. The Salen Electro Fusion Fitting System (SEF) was developed to weld thermoplastic materials by means of electro fusion. A deﬁnite welding pressure is absolutely necessary to obtain a homogeneous reliable connection of the welded surfaces. Two procedures were developed for this system: • One procedure is based on exact tolerances of the gap clearance between the inner diameter of the electro fusion ﬁtting and the outer pipe diameter. Due

409

8.3 Electro Fusion Fitting System to the tight tolerances for manufacturing this procedure did not succeed in the market. • The second procedure is based on the “memory-effect” of stretched polyethylene, which is the solution for the electro fusion system. The electro fusion socket is widened at the end of the manufacturing process, so that the ﬁtting can be inserted easily over the pipe. During the welding procedure the socket shrinks, thus ensuring the required welding pressure. This system is advantageous because it results in a reliable welding connection regardless of the wide tolerances of the PB pipes. Figure 8-7 Electro fusion socket

8.3.1

The SEF-System

Electro fusion socket, pipes and ﬁtting of the SEF-System have welding surfaces tapered to a deﬁnite shape. When joining these tapered surfaces a close tolerance ﬁt is obtained, which ensures the required welding pressure. The welding wire is injection molded into the ﬁtting so it cannot be damaged during the assembly procedure. Compensation can be made, within reasonable limits for inaccurate tolerances of sockets and ﬁttings, and oval pipe sections; the assembly can be easily performed by means of the specially developed tools; the welding procedure requires no special personnel skills. In all cases the result is a 100% reliable welding connection.

Figure 8-8 Fix clamping device

The Components of SEF-System • Electro fusion socket in dimensions from 18 through 110 millimeters according to ISO • Fittings • Socket saddle in dimensions 63, 90 and 110 millimeters • Welding machine for dimensions 18 and 20 millimeters • Welding machine for dimensions 20 through 110 millimeters • Tools: Clamping device Fixing clip Pipe milling tool

Figure 8-9 Connect wires to the power supply

Advantages of SEF-System • Easy preparation of welding procedure • Safe and easy handling, granting a speedy assembly on site

Figure 8-10 Melt formation bottoms show good weld joints

• Easily adapted in case of on site restrictions

Welded joint

• No welding seams at the welded joint • High safety factor

Pipe

Sequence of Operations The simplicity of the electro fusion welding process is as easy as shown in Figures 8-7, 8-8, 8-9, 8-10 and 8-11.

Pipe

Inner flash

Electro fusion socket

Welded joint

Figure 8-11 Cross section of electro fusion welded joint

410

8 Thermoplastic Assembly Methods

8.4

Hot Plate Welding Method

Hot plate welding has been used successfully as a rapid assembly method for producing strong, permanent, leak free joints using amorphous thermoplastic and some crystalline materials with slow crystallization rates like is the case of some high viscosity polymers. Hot plate welding consists of heating two injection molded products to be joined with a hot plate. The hot plate or heated platen, at the correct temperature, is placed at the interface of the injection molded products to be joined. The products are then brought into contact, or in some cases very close to, the hot plate surface. Stops on the insert and the ﬁxtures that hold the components determine the depth of the initial melt. The hold time determines the depth of the secondary softening. This initial hot plate contact produces a good mating surface as both components are made level by the hot plate. At this point the insert is removed and the components are pressed together by the holding jigs. The stops on the holding jigs (ﬁxtures) now determine the amount of joint material displaced during mating. Figure 8-12 illustrates the basic principles of the hot plate welding process. After the components’ joint surfaces have been melted, they are quickly brought together and held under a slight pressure for a short holding time until the molten surface layer cools off, producing a weld between the two joint surfaces. The recommended hot plate welding equipment is an aluminum plate with a thin coating of Teﬂon® ﬂuorocarbon resin and a thermostatic switch. This will maintain a controlled and uniform temperature throughout the plate and the Teﬂon® will prevent the molten material from sticking to the hot plate.

Moving fixture

Top cap Joint

Hot plate Body Fixed jig

Loaded cap and body in holding fixtures, while hot plate slides-in

Moving fixture lowers down to contact the hot plate surface

Moving fixture raises away the hot plate

Molten joints

Cap and body molten joints, moving jig up, hot plate retracts

Figure 8-12 Hot plate welding process method sequence

Moving fixture lowers down pressing both molten joints

Moving fixture raises, welding done, ejection

411

8.4 Hot Plate Welding Method Figure 8-13 shows some typical thermoplastic applications welded by the hot plate welding process method. Figure 8-14 shows two types of typical hot plate welding equipment. Figure 8-15 shows the hot plate welding process for thermoplastic tubing. Injection molded or extruded thermoplastic products can be hot plate welded and sealed following procedures that are similar, but may differ slightly with the type of material. For example, products molded of polycarbonate resin that are to be joined by hot plate welding or sealing must be predried at 250 °F if maximum bond strengths are to be obtained between the joint surfaces. Drying times at this temperature depend on part wall thickness. This range varies from 30 minutes for 0.031 inches wall thickness to 14 hours for 0.187 inches wall thickness. Advantages • Joint can be curved in all planes (irregular shapes) • Achieves consistent hermetic seal • Eliminates glues, solvents and adhesives • Replaces the use of ultrasonic horns • Eliminates screws, clips, gaskets and preforms

Figure 8-13 Typical hot plate welding applications

• Can weld internal walls • Allows internal products to be installed before welding Disadvantages • Long cycle times • Equipment cleaning • Material degradation possible • High temperature requirements • Only the same type of materials can be joined by this method. • Polyamide or nylon based resins are unsuitable for hot plate welding since they oxidize when the melted resin is exposed to air during the hot plate welding cycle. The injection molded nylon product will oxidize and will not weld properly. • Some sticking problems between the polymer and the hot plate are possible. Teﬂon® PTFE coating of the plate tends to reduce this considerably.

Welded joint of both tubes

Fixture Tubing Hot plate

Tubing alignment

Hot plate melting both joints

Tubing weld finished

Figure 8-15 Thermoplastic tubing hot plate welding process sequence

Figure 8-14 Two types of hot plate welding equipment

412

8 Thermoplastic Assembly Methods

8.4.1

2.5 T T

1.5 T - 0.06 in.

0.03 in. weld

0.75 T

Figure 8-16 Butt joint with external flash

2.5 T T

0.38 T x 45˚ Chamfer (typ)

0.03 in. weld

0.75 T

Hot Plate Welding Joint Design

Products must be designed correctly to avoid rejects and failures. The ﬂatness of the joint surface area is essential and therefore the design principles for injection molding thermoplastic materials should be strictly applied. In particular, uniform and symmetrical wall sections, suitably designed with radiused corners are vital. Joint design is also important in hot plate welding for cosmetic reasons. The material displaced during the joining process will form a bead at the weld area. It may be necessary to design the joint to hide such a bead, or the bead can be removed after the welding process is completed. Because the hot plate welding method applies heat to the joint surface areas, these joints must be ﬂat and the use of only butt joint and variations on butt joint designs is recommended. If the proper joint design, as illustrated in Figures 8-16, 8-17, 8-18, 8-19, 8-20, and 8-21 (detailed butt joint designs) is chosen during initial design stages of a product, the hot plate welding method can produce a welded joint that is equal to or stronger than any other area of the injection molded thermoplastic product. Designing the proper joint for hot plate welding depends heavily on an understanding of the hot plate welding process itself.

1.5 T - 0.06 in.

Figure 8-17 Butt joint with flash pocket

T 0.5 T

3T T 2.5 T

0.75 T 0.5 T

0.03 in. weld

3.0 T - 0.06 in.

0.03 in. weld 1.5 T

0.75 T

T

2.0 T - 0.06 in.

Figure 8-18 Butt joint with single flash trap

Figure 8-19 Butt joint with double flash trap

0.03 in. weld

T

T

1.5 T

0.5 T

0.5 T T 0.5 T

0.75 T

T T

0.25 T 2.25 T

Figure 8-20 Bi-level butt joint with internal flash trap

413

8.5 Solvent and Adhesive Bonding Methods During operation, a heated platen moves into position between the two separate products, which are held in place with specially designed holding ﬁxtures. The holding ﬁxtures move inward and press the two separate products against the heated platen and typically a melt displacement of 0.030 in occurs on each side of the two product parts. This displacement may vary slightly depending on part design, material used, and the strength of the weld required.

0.5 T

0.03 in.

After the two separate part edges are molten, the holding ﬁxtures open and the heated platen is withdrawn. The ﬁxtures then close, forcing the two separate parts together at their melted joint, creating a molecular seal as the thermoplastic cools. During this cooling phase a “seal” displacement, typically 0.015 in, occurs on each side of the welded components.

1.5 T + 0.03 in. 0.03 in. weld

8.4.2

Flash or Weld Bead

During the hot plate welding process method, approx. 0.060 in of total molten polymer is displaced from each welded assembly. This includes 0.03 in material displacement per side: 0.015 in during melt and 0.015 in during seal. This displaced material creates a weld bead that is often referred to as “ﬂash”.

T

T

T

The molten polymer displacement (flash) after hot plate welding can be critical: • Tight tolerances: In intricate assemblies, where dimensional tolerance is critical, material needs to be added at the mating surfaces of the joint to allow for material displacement during the weld cycle. • Aesthetics: A small uniform bead of ﬂash around the joints indicates that a good weld has been formed. When the appearance of ﬂash is undesirable, either from a decorative point of view or from a functional standpoint, an extra processing step is required. In parts where external appearance is important, the weld bead must be hidden or removed, for example by designing a ﬂash trap into the part (see Figures 8-17, 8-18, 8-19, 8-20, and 8-21). If there is insufﬁcient design ﬂexibility to hide the bead, it can be removed in a ﬂash pinch-off operation after the welding process.

8.5

Solvent and Adhesive Bonding Methods

Solvent bonding, or solvent welding, as it is sometimes called, is a common technique used for joining injection molded components of amorphous thermoplastic resins. When the components are bonded in this manner, the solvent dissolves the surface of the two mating components and allows the material to ﬂow together. After the solvent evaporates, it leaves a pure material-to-material bond. When solvent bonding, the components should be dampened with solvent and then assembled using moderate pressures to hold them together. This wetting process can be accomplished by wetting a piece of felt or a mating preform made of wood with solvent and lightly pressing the parts to be bonded against them. The components can be safely set aside to dry after 40 to 60 s of hold time, but full bond strength will not be achieved for 24 to 48 hours. In some cases, it may be desirable to heat the assembly in an oven to drive off the excess solvent.

T 2T

Figure 8-21 Recessed butt joint with external flash

414

8 Thermoplastic Assembly Methods It is very important for the solvent bonding method to have well matched component joint surfaces. Often, unsuccessful attempts are made to resolve bonding problems by using excessive amounts of solvent or pressure on the mating components. These “remedies” will only result in a sloppy or overly stressed joint. Locator pins and/or tongue and groove assemblies can help to provide good part matching as well as easy part location after the application of solvent. Both joint surfaces should be shallow to avoid entrapment of excessive amounts of solvent. If mating the molded components continues to be a problem, slurries made of solvents and up to 25% base resin may be used. Advantages • Distributes stress over the bonded surface areas • Provides aesthetic bond • Can provide bond instead of bosses • Can provide hermetic seal Disadvantages • Bonded products cannot be disassembled • Sufﬁcient surface area must be available for proper joining • Solvent vapors released may be hazardous • Only practical with amorphous and some semicrystalline materials • Tight molding tolerances are required to produce properly bonded mated products Adequate ventilation should always be provided to avoid possible health hazards posed by the use of solvents. OSHA, EPA, and local regulations should always be researched and obeyed. Make sure to avoid direct contact with the solvents.

8.5.1

Solvents Used to Bond Thermoplastic Polymers

For many amorphous and some semicrystalline resins there is more than one solvent that can be used for bonding. Several polymers and various types of solvents that are appropriate for bonding are listed here. ABS • Acetone • Methylene chloride • Methyl ethyl ketone • Methyl isobutyl ketone • Tetrahydrofuran Acrylic (PMMA) • Methylene chloride • Ethylene dichloride • Trichloroethlene

8.5 Solvent and Adhesive Bonding Methods Cellulosic • Acetone • Methyl ethyl ketone Nylon (PA) • Aqueous phenol • Resorcinol in alcohol • Calcium chloride in alcohol Polycarbonate (PC) • Ethylene dichloride • Methylene chloride Polystyrene (PS) • Ethylene dichloride • Methylene chloride • Ethylene ketone • Trichloroethylene • Toluene • Xylene Polysulfone (PSU) • Methylene chloride PPO/PPE • Trichloroethylene • Ethylene dichloride • Methylene chloride • Chloroform PVC • Methyl ethyl ketone • Cyclohexane • Tetrahydrofuran • Dichlorobenzene

415

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8 Thermoplastic Assembly Methods

8.6

Adhesive Bonding Method

Adhesive bonding is one of the most convenient methods of assembling a thermoplastic product to either a similar or dissimilar (metals) material. Adhesives distribute stresses over the entire bonded surface area and can provide a hermetic seal if needed. Flexible adhesives allow some movement between the mating joint surfaces and thus can compensate for differences in the coefﬁcients of linear thermal expansion of the materials. Adhesives are relatively inexpensive and often require little or no special equipment for application. Advantages • Ease of application using automatic equipment • Can bond dissimilar materials • Uniform stress distribution along the bonded surfaces • Bonded surface sealing • Elasticity • Low cost • Reduce hardware • Insulation • Ease of repair • Aesthetics

8.6.1

Adhesive Families

There are ﬁve major families of adhesives, each with a somewhat unique set of characteristics. When choosing an adhesive for a speciﬁc application, these characteristics can provide an initial direction for the search. For example, if the end use environment will exceed 212 °F, an epoxy or an anaerobic adhesive would be worth considering. Or, if set time and cure time are critical during assembly, a cyanoacrylate or anaerobic adhesive might be appropriate. Epoxy (EP) Epoxies are known for their versatility. Their bond strength, electrical conductivity, and temperature resistance can be modiﬁed to ﬁt almost any speciﬁc application needs. Epoxies are made in either one or two part formulations. The two-part systems consist of a resin and a hardener that must be mixed together in strict proportions for maximum bond strength. They can be cured at room or at elevated temperatures. One-part epoxies require no mixing; however, they must be heat-cured at approx. 300 °F for one hour or more. Heat cured epoxies tend to exhibit greater strength than their mixed counterparts. However, two-component systems are more widely used because they may be stored for long periods of time and will not activate until mixed. Unlike other adhesives, epoxies are not solvent-based, but cure as the result of a chemical reaction. Liquid paste form Shear strength: from 5,000 to 10,000 psi Operating temperature: –70 to 450 °F

8.6 Adhesive Bonding Method Advantages • Good adhesion • High tensile and shear strength • Creep resistance • Good rigidity • High heat tolerance • Easy to cure Disadvantages • Poor peel strength • Brittle • Low impact strength • High cost Urethane (TPU) This adhesive family, also called polyurethanes, can provide strong bonds for a variety of substrates. Urethane is primarily found in applications that require high strength as well as ﬂexibility. Urethane is available in both one- and twopart systems. One-part formulations require heat curing, while two-part systems may be cured at room temperature. Liquid paste solvent-based form Shear strength: up to 8,000 psi Operating temperature: –300 to 300 °F Advantages • Toughness • Flexibility • Impact strength • Abrasion resistance • High peel strength Disadvantages • Volatile and excessive creep • Poor strength at high temperature • Chemical sensitivity • Lacks long term durability • Usually needs primers • Moisture sensitive in uncured state

417

418

8 Thermoplastic Assembly Methods Acrylic (PMMA) The acrylics used today are second generation or modiﬁed acrylic systems. These “improved” acrylics provide many of the same attributes as the epoxies and urethanes; in addition, they also offer the advantage of rarely needing primers. There are one- or two-part systems, the latter consisting of a catalyst primer and the adhesive. Usually, the two-part systems do not need mixing or weighing, which simpliﬁes their application immensely. Acrylics boast rapid cure at room temperature with a setting time of approximately 60 to 90 s and full cure within 30 min or less. The application of heat may be used to reduce cure times. Liquid paste form Shear strength: up to 6,000 psi Operating temperature: between 240 and 350 °F Advantages • Bonds to contaminated surfaces • High strength

• Superior toughness • Fast curing

Disadvantages • Strong odor

• Problems with ﬂammability • Minimal gap ﬁlling Anaerobics Anaerobics are a one-part thermosetting adhesive family whose curing mechanism is triggered by the absence of oxygen. This eliminates the problem of premature curing. Curing occurs at room temperature and the addition of heat or ultraviolet radiation will increase the speed of the curing process. The cure cycle may be as short as 15 s set time and 2 to 24 h for full cure. Anaerobics also exhibit the useful property of being easily cleaned from unbonded surfaces after the bond line has set up. Anaerobics are excellent for critical sealing and bonding applications where strength is not needed. Their use is also expanding into the sealing of welds and soldered joints. Liquid form Shear strength: up to 5,000 psi Operating temperature: between 65 and 400 °F Advantages • Good solvent resistance • Bond ﬂexibility

• High peel strength

• Good impact strength

8.6 Adhesive Bonding Method Disadvantages • Sensitive to surface cleanliness • Poor gap ﬁlling properties Cyanoacrylates Cyanoacrylates are single-part, fast curing “convenience adhesives”. With a normal setting time of 2 to 3 s and a full cure time of 24 h at room temperature, these systems are popular in tacking and quick contact assembly operations. Curing is initiated by the presence of surface moisture, even in limited quantities, such as humidity in the air. These adhesives are highly application speciﬁc in their use. Liquid form Shear strength: up to 5,000 psi Operating temperature: between 65 and 180 °F Advantages • High tensile strength • No shelf life limitations Disadvantages • Brittle • Not usually suggested for dissimilar materials • Poor gap ﬁlling • Not suggested for constant water exposure • Limited impact and peel strength

8.6.2

Adhesive Concerns

Other properties to be considered in the choice of an adhesive include resistance to the end use chemical environment, bond strength in the speciﬁc joint design, and long-term temperature stability. Assembly facility, shelf life, handling requirements, and potential toxicity issues must also be considered. The following adhesive characteristics should be evaluated before a ﬁnal choice is made. • Chemical resistance • Bond strength • Temperature stability • Handling • Toxicity • Low cost • Compatibility • Surface preparation

419

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8 Thermoplastic Assembly Methods

8.6.3

Adhesives Bonding Selection

The best adhesive for a job may not be the one that produces the strongest bond to the material in question. The most important factor in selecting an adhesive is the intended use of the product. Will it be used underwater or exposed to sunlight? Is the adherent transparent, requiring a colorless adhesive? Some guidelines for selecting the proper adhesive are listed below: Flexibility Flexible materials must be bonded with ﬂexible adhesives. A bead of ﬂexible adhesive can act as strain relief between the molded components. When dissimilar materials are bonded, the adhesive must be ﬂexible enough to accommodate their different coefﬁcients of linear thermal expansion. Other Assembly Considerations • Will the adhesive be expected to cure quickly in preparation for the next fabrication step, or can it rest in a ﬁxture for a while? • Does the manufacturing facility have access to equipment for UV curing or visible light curing? • Will the part be assembled by human hands (subject to burns and poisoning) or by robotic equipment? • Do you expect the joints to perform better in a laboratory set-up environment or when assembled on the production line? • An adhesive that performs marginally in testing probably will fail when applied in haste during production. Solvent degreasing stations recycle the solvent; therefore, the quality of bonds may suffer from gradual or sudden contamination of the solvent. Many adhesives require a human touch for assembly because the joining surfaces must be slid together to exclude air. An adhesive may produce fumes from which workers need to be protected. Temperature If a heat curing adhesive will be used, the molded products must be able to withstand the curing temperatures without deformations and expansion. Aging Strength retention values with aging (creep) are often determined in a laboratory, with the thermoplastic sample bar loaded under controlled temperature and humidity conditions. The joints aged under load, especially cyclic load under the end use environment, may lose their strength more rapidly than in the laboratory.

8.6 Adhesive Bonding Method

8.6.4

Ultra Violet Curable Adhesives

In the mid 1960’s, ultra violet curing technology was ﬁrst used for inks in the packaging industry. Since that time, the technology has advanced and spread into other applications. Besides inks, adhesives, potting compounds, encapsulates, and sealants are now available with ultra violet curing systems. Ultra violet curable adhesives are one of the fastest growing families of adhesives in use today. Ultra violet curing uses light energy to initiate polymerization. In industrial applications, ultra violet light is provided by low pressure mercury vapor lamps. In the early stages of development, ultra violet curable adhesives could be used only with transparent substrates or other applications that would allow the entire bond line to be exposed to ultra violet light. Since that time, technology has overcome this problem. Today, most ultra violet curable adhesives include a secondary cure mechanism. Anaerobic, oxygen, and moisture cures are often used to cure those areas not directly exposed to light. The main drawback of ultra violet adhesives is cost. Ultra violet curing equipment as well as the adhesives are relatively expensive. Therefore, careful consideration must be given to whether these adhesives are economical for a given application. Trigger cure is an alternative for bonding opaque substrates. This method consists of applying the adhesive to one half of an assembly and exposing it to ultra violet light. The molded components to be bonded are then ﬁxed together trapping the previously activated adhesive between them. The initiated cure then continues to completion and forms a secure bond. This method is slower than the typical ultra violet cure cycle and may provide bond strengths somewhat lower than direct light curing. Ultra violet adhesives offer many advantages. They are 100% reactive liquids that avoid the hazardous solvents used in other systems. This aspect also improves dimensional stability in these adhesives, because there is no loss of volume caused by solvent evaporation. Ultra violet curing is also extremely fast, typically from 3 to 10 s. This characteristic allows much faster production times and eliminates the need for dedicated ﬁxtures in the assembly lines. Cyanoacrylate adhesives can provide similar production times, but they cannot compete with the physical properties of the ultra violet adhesives. Tables 8-1 to 8-5 list the various adhesives that have been tested and found to be compatible with thermoplastic resins. All joints were tested as solvent wipe and machined surface at room temperature. The tables are only guidelines; it is necessary for the end user and the adhesive supplier to work together in resolving speciﬁc adhesive problems.

421

422

8 Thermoplastic Assembly Methods Table 8-1 Acetal, Polyacetal (POM)

Supplier

Adhesive

Adhesive type

Dexter

EA 9412, EA-6

Epoxy

Ren Plastics

RP 5540/5542

Polyurethane

3M

Scotchgrip EC-1711

Rubber

M&T

Uralane 5738

Polyurethane

Sears

Hot Melt

Du Pont

46950, 2% RC 803

Polyester

Du Pont

4684

Rubber

Lord

Chemlock 220, 203

Rubber

Harman, Inc.

Phenoweld 7

Phenolic

Crest Products Corp.

Narmco 3135A, 3135B

Modiﬁed Epoxy

Fiber Resin Corp.

Resiweld 7004

Epoxy

Fiber Resin Corp.

Cycleweld C-14

Epoxy

Applied Plastics Co.

Apco 5363, 5313

Epoxy

Chemical Development

CD 200

Vinyl

Permabond International

Permabond

Alpha Cyanoacrylate

Oneida Electronics Mfg.

Instant Weld

Alpha Cyanoacrylate

Eastman Chemical

910 MHT, THT

Alpha Cyanoacrylate

Table 8-2 Polyamide (PA)

Supplier

Adhesive

Adhesive type

American Cyanamid

Cybond 1101, 1110

Epoxy

Dexter

Hysol EA 9412

Epoxy

Crest Products Corp.

Crest 3147/7112

Epoxy

Amicon

XT 1242

Epoxy

M&T

Uralane 5738

Polyurethane

Ren Plastics

RP 5540/5542

Polyurethane

Goodyear

6000/6012G,

Polyurethane

Oneida Electronics Mfg.

Instant Weld 104, 101

Cyanoacrylate

Loctite

414

Cyanoacrylate

Permabond International

Permabond 105, 102

Cyanoacrylate

Lord-Hughson

Cylok G, M

Cyanoacrylate

Armstrong

A-12

Epoxy

3M

Scotchweld 2214

Epoxy

3M

Scotchcast 4407

Dexter

Hysol EA-6

Epoxy

Crest Products Corp.

3135

Epoxy

H. B. Fuller

Resiweld 7004

Epoxy Phenolic

Hardman

Phenoweld #7

Applied Plastics

Apco 5363

Eastman Chemical

910, 910 ET

Cyanoacrylate

Loctite

430, 414

Cyanoacrylate

Oneida Electronics Mfg.

747, 240

Cyanoacrylate

423

8.6 Adhesive Bonding Method Table 8-3 Polyethylene Terephthalate (PET)

Supplier

Adhesive

Adhesive type

Goodyear

6000, 6040, AX37J922

Polyurethane

Ciba-Geigy

RP 5540/5542

Polyurethane

M&T

Uralane 5738

Polyurethane

H. B. Fuller

UR 2139

Polyurethane

3M

Scotchweld 2214

Epoxy

3M

Scotchweld CA-40

Cyanoacrylate

Armstrong

A-12

Epoxy

Dexter

Hysol EA 9412

Epoxy

American Cyanamid

Cybond 1101

Modf. Epoxy

Loctite

430, 496, 414

Cyanoacrylate

Eastman Chemical

910, 910ET

Cyanoacrylate

Oneida Electronics Mfg.

Instant Weld 130, 107, 747

Cyanoacrylate

Permabond International

Permabond 102, 105, 747

Cyanoacrylate

Lord-Hughson

Cylok R, G, M

Cyanoacrylate

Supplier

Adhesive

Adhesive type

Lord

Tyrite 7411

Urethane

Table 8-4 Copolyester (TPE)

Lord

Tyrite 7500 A/B or A/C

Urethane

Lord

Chemlok 250

Primer

Locktite

Prism 406

Cyanoacrylate

Table 8-5 Other Thermoplastic Polymers

Supplier

Adhesive

Adhesive type

Family of plastics

3M Company

Scotch-Weld 1838L

Epoxy-2-Part

PC, PBT, PPO, PC/ PBT, PEI, ASA

3M Company

Scotch-Weld 2214

Epoxy-1-Part

PBT, PPO, PC/PBT

3M Company

Scotch-Weld 2216 4-7 Epoxy-2-Part

PC, PBT, PPO, PC/PBT, PEI, ASA, ABS, PC/ABS

3M Company

Scotch-Weld 2216 4-7 Epoxy-2-Part

PC, PBT, PPO, PC/PBT, ASA, ABS, PC/ABS

3M Company

Scotch-Weld 3532

Urethane-2-Part PC, PBT, PPO, PC/ PBT, ASA, ABS

Synthetic Surfaces

80 J

Epoxy-2-Part

CIBA-GEIGY

Arathane 5540/41

Urethane-2-Part PBT, PPO, PC/PBT, PEI, ASA

PEI, PC/PBT

Essex Chemical

Betamate 73553/72034 Urethane-1-Part PBT, PPO, PC/PBT, Primer PEI

Lord Corp

Tyrite 7500 A/C

Urethane-2-Part PC, PBT, PPO, PC/PBT, PEI, ASA, ABS, PC/ABS

Emerson & Cuming LA2337-8

Epoxy-1-Part

PBT, PC/PBT, PEI

Emerson & Cuming Uniset A 401-37

Epoxy-1-Part

PBT, PC/PBT, PEI

Miller Stephenson

Epoxy-2-Part

PBT, PC/PBT, PEI

EPON 828 V-40

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8 Thermoplastic Assembly Methods

8.6.5

Adhesive Surface Preparation

For optimal bonding results, the surface of the substrate must be clean and dry. Surface preparation includes any method, mechanical, chemical, or electrical, used to remove contaminants that might interfere with the bonding process. The amount and type of surface preparation needed depends on the adhesive used, the bond strength required, the end use environment, and economic practicalities. To achieve the most effective adhesion for the molded products, bonding should take place soon after surface preparation has been completed. Degreasing Cleaning This is the most common method of surface preparation and is usually used before any type of adhesive bonding. The surface of the molded product to be bonded should be clean and relatively free of contaminants. Among the cleaning methods used are vapor degreasing, immersion baths, solvent spraying or wiping, and acid etching. Care should be taken in ensuring that the chemicals used for degreasing are compatible with the molded products being cleaned as well as EPA and OSHA approved. Abrasion This method of surface preparation includes sandblasting, sandpaper, and “Scotch-Brite”. An increase of up to 15% in adhesion strength after abrasion can be achieved for some materials. Notch sensitive materials require careful application of this method. It should also be noted that some type of solvent cleaning is usually needed both before and after abrasion to remove any surface contamination. Flame Treatment With this method, the area of the molded product surface to be bonded is exposed to the oxidizing area (i.e., blue, not yellow) of a ﬂame. This is done with a controlled ﬂame, usually natural gas, until the surface is shiny. Care must be taken to avoid overheating or melting the product. This method is used for oleﬁn materials. Corona Discharge The injection molded thermoplastic component to be bonded is placed between two electrodes where the bombardment of electrons and ions causes the surface to be oxidized. Corona discharge equipment was mainly developed for changing the surface tension on ﬁlms, but new developments allow for more complex parts to be treated. The corona discharge treatment is helpful in improving the bonding of adhesives as well as a primer for printing inks and paints. Plasma Treatment The plasma treatment is a variation of the corona discharge. The injection molded thermoplastic components are placed in a vacuum chamber that is ﬁlled with an inert gas. A high electrical charge is applied to the gas and the thermoplastic component surfaces are bombarded with electrons causing a similar surface oxidation. Chemical Etching This is a wet process in which the injection molded thermoplastic components to be bonded are dipped into an acid solution. The etching solution forms either

8.6 Adhesive Bonding Method highly oxidized or complex inorganic surface layers. Chemical etching requires careful control.

8.6.6

Adhesive Application and Curing Methods

How an adhesive is applied is as important to successful bonding as the type of adhesive applied. Methods of application fall into one of three categories: manual, semi-automatic, and automatic processes. Manual processes, although slow, offer the combined advantages of simplicity and low equipment cost. Semi-automatic processes represent a compromise between the speed and expense of automatic equipment and the comparative slowness and economy of manual equipment. Automated processes, on the other hand, provide greater speed and the potential for improved consistency, but usually require expensive machinery. Occasionally, there are also unique applications when conventional methods are not adequate and a special process must be devised to apply the adhesive. Following are some common guidelines to be followed for the application of adhesives on injection molded thermoplastic components: • The assembly should be kept clean of dust and contaminants. Control of humidity and temperature in the assembly area is dependent on the demands of the adhesive system. • Separation of the adhesives may occur during storage. Care should be taken to thoroughly mix the adhesive before use. In the case of two component systems, proportional blending affects the bond strength and mixing should be done carefully. • For both automated and manual application methods, bond line thickness may affect the strength of the bond. Generally, maximum strength and rigidity will be achieved using as thin an adhesive layer as possible without starving the joint. Thin bond lines require greater forces to deform, are less apt to creep, and have a lower probability of containing air bubbles. Thin ﬁlm layers also have a greater resistance to cracking. • The curing of adhesives should take place under controlled humidity, temperature, and timing conditions. The selection of cure time, rate of heat rise, and rate of cool down are dependent on the adhesive formulation, type of joint, and expected service condition of the bond. • Uniform pressure should be maintained over the entire joint surface area during the curing cycle. Some pressure is required to prevent movement of the assembled molded components to obtain a uniform bond layer, as well as to overcome the internal pressures created by out gassing of the adhesive.

8.6.7

Joint Design for Adhesive Bonding

Adhesive bonding is a complex phenomenon involving chemical reaction, electrical attraction at the molecular level, and various mechanical factors. The adhesive introduced at the mating surface of two components must be compatible chemically, electrically, mechanically, and with the end use environment. In addition, the adhesive medium should have a similar coefﬁcient of linear thermal expansion to the injection molded thermoplastic components, or, if they are dissimilar, the adhesive system must be ﬂexible.

425

426

8 Thermoplastic Assembly Methods Load

Joint design is critical to the optimum performance of a bond. Factors to be considered in choosing a joint design include:

Load

• Joint orientation will deﬁne the type of loading applied to the bond shear, peel, and tensile characteristics. Poor

Good

Load

Load

Load

Load

• Optimize the joint surface area of the bond to match the adhesive strength and the expected loadings. • Aesthetics in the bond area may restrict the choice of joint design. • Mold design and injection molding process efﬁciency can be compromised by some joint designs.

Poor

Load

Good

Load

Poor

Load

Load

Good

Load

Poor Load

Good Load

Good

• Molding processing tolerances on the mating joint surface areas are critical to some joint designs, such as the tongue and groove. • Product design and mold design must anticipate the tolerances required at the mating joint surface areas of the components. • Gap ﬁlling ability of the adhesive must be considered in the dimensions of the joint. • Set time for curing the adhesive and the time for handling the thermoplastic products after assembly should be controlled to obtain the best bonding strength and consistency of the joints. The type of joint design selected for an application depends on the characteristics of the adhesive system. The adhesive method does not operate just in a single point like rivet fasteners, instead, the adhesive structural bond is dispersed over the entire joint surface area,. An adhesive joint should be designed without any type of internal sharp corners to minimize stress concentrations. Figure 8-22 shows four typical adhesive joint designs. Comparisons of poor designs and recommended joint designs are shown. Figure 8-23 shows poor adhesive butt joint designs of two mating ﬂat plates and solid round bars. Figure 8-24 shows three fair to good adhesive joint designs with larger bonding surface areas. Figure 8-25 shows nine recommended adhesive joint conﬁgurations.

Figure 8-22 Typical poor and recommended adhesive joint designs

Single lap joint

Strap joint Flat plates adhesive butt joint

Joggle lap Solid round bars adhesive butt joint

Figure 8-23 Poor adhesive joint designs

Figure 8-24 Fair to good adhesive joint designs

427

8.7 Metal Fasteners Method

Taper lap joint

Double lap joint

Scarf joint

Stepped lap joint

Taper double strap joint

Double strap joint Collar improves seal

Joggle lap/strap

Hollow cylinder

Tube “T” junction

Figure 8-25 Recommended adhesive joint designs

8.7

Metal Fasteners Method

Injection molded thermoplastic components are frequently assembled with metal self-tapping screws, bolts, nuts, washers, and locking springs that are reliable and economic. Some basic design principles should be applied to the injection molded thermoplastic components that are going to be assembled with fasteners. The thermoplastic components under the screw head should always be under compressed load and never under tension. The top illustration in Figure 8-26 shows a poor design, where the screw bends the supported beam, stressing this member under tension. The bottom illustration is the recommended design, where the screw load is distributed by the washer, compressing the free support beam against the boss. The left illustration in Figure 8-27 shows a poor design, where the bolt and nut stress puts tension on the beam. The right illustration shows the recommended design, where the compression load is supported by the spacer and distributed by the washer. Figure 8-28 shows different types of screw heads; the countersunk head is not recommended for thermoplastic assembly designs. Countersunk screws, as shown in Figure 8-29 left, produce undesirable hoop stresses around the tapered hole and should not be used in thermoplastic applications. The right illustration shows a hex bolt head with a ﬂat underside. This type of head is preferred because the stress produced is more compressive. Flat washers are helpful in distributing the assembly force and should be used under both the nut and the fastener head. Hex nuts and hex ﬂange head nuts and bolts have ratchet-like circular groove teeth on the underside of the fastener heads. When tightened, the circular ratchet-like teeth of the fastener bite into the thermoplastic surface to provide locking action to the fastener; the ratchet-like circular grooves increase spring action (similar to a spring lock washer), resist vibration, and maintain clamping load. Figure 8-30 shows the circular ratchet-like teeth locking mechanism.

Concentrated screw load

Bending (tension)

Screw Boss End support Poor design Washer

Compression

Beam

Free support Recommended design

Figure 8-26 Type of loading caused by selftapping screw

428

8 Thermoplastic Assembly Methods Concentrated screw load

Bending (tension)

Washer

Compression

Spacer

Bolt

Free support

End support

Figure 8-27 Type of loading caused by bolt and nut

Figure 8-28 Different types of screw heads

Nut

Nut

Recommended design

Poor design

Truss head with washer

Pan head with washer

Stress concentration by countersunk screw

Hex head with washer

Avoid countersunk head for plastics

Hex bolt & washer

Compression

Nut

Figure 8-29 Compression load effects caused by countersunk screw

Figure 8-30 Circular ratchet-like serrated teeth fasteners

Figure 8-31 Arched spring lock nut

Compression

Recommended design

Poor design

Hex nut

Washer

Nut

Hex flange head nut

Figure 8-32 Blind spring expansion fastener

Hex flange head bolt

Figure 8-33 Self-retainer U-type snap fastener

429

8.7 Metal Fasteners Method

Post

Plastic expand rivet

Compression

Expansion Plastic arrow clip rivet

Poly drive 2 piece rivet

Metal screw

Snap lock fingers

Taper fins Round taper fins rivet

Plastic fastener

Plastic snap push-in rivet

Spread by screwing Screw spread grommet nut

Metal screw Metal screw

Deformed by screwing Cold-pressed anchor metal fastener

Drywall plastic punch & screw anchor fastener Locking drive pins

4 spreaders to lock in plastic hole wall Screw spread metal knurl fastener

In/out threads lock drive pins metal fastener Threaded product

Heli-coil thread insert

Bolt

Heli-coil Heli-coil steel alloy diamond wire internal and external thread insert

Figure 8-35 Typical commercial fasteners for assembly applications

Joining methods for thermoplastic components by means of nut-type metal springs are used in many applications that require a reliable and economic fastener. Figure 8-31, 8-32, 8-33, and 8-34 show the different types of metal spring fasteners. Figure 8-35 shows more commercial types of fasteners made either by injection molding thermoplastics or by a sheet metal forming process.

8.7.1

Thermoplastic Bosses and Self-Tapping Screws

Thermoplastic boss design depends on the boss wall thickness, the boss inside diameter depth, the thermoplastic boss material modulus of elasticity, the type of self-tapping screws, the pull-out force, and the stripping torque required for the application. Figure 8-36 shows basic thermoplastic boss and self-tapping screw parameters.

Figure 8-34 Single thread crown bite locking nut

430

8 Thermoplastic Assembly Methods DS = Screw major diameter dS = Screw minor diameter DB = Boss Outside Diameter dB = Boss inside diameter L = Full thread engagement

D B dB

L

Boss Inside Diameter (dB) For the highest pull-out force and stripping torque, it is recommended that the boss inside diameter equals the pitch diameter of the self-tapping screw. Boss Outside Diameter (DB)

D S dS

It is recommended for the boss outside diameter to be equal to 2.5 times the self-tapping screw major diameter, or the diameter recommended for special self-tapping screws. If a boss wall is too thin, the boss may crack and if a boss wall is too thick, the boss will produce unacceptable stripping torque. Screw Length Thread of Engagement (L)

Figure 8-36 Thermoplastic bosses and selftapping screws

The pull-out force and stripping torque values increase proportionally to the screw thread of engagement until the length reaches 2.5 times the screw outside diameter. This length should be the minimum distance used for screw applications.

8.7.2

Thread Forming and Thread Cutting Screws

In many self-tapping screw applications, failures occur because of inadequate design of the thermoplastic bosses. When analyzing cracked bosses caused by thread forming or cutting screws, the failure is often the result of high stresses from the screw or improper boss geometry. Self-tapping, and in particular, thread forming screws generate high stress in the thermoplastic boss section. It is of the highest importance to distribute stresses equally around the screw, especially when the boss has a uniform circular section. Even if a boss has a perfect circular shape, weak areas can still be present. Depending on the overall design, the gate location and the molding conditions, there may be a weld line over the length of the boss. If the gate is located close to a screw hole, an additional possible weak spot may exist. If improper molding conditions are used, molded-in stress may be present at the gate and a poor weld line may be formed. Therefore, the designer should be familiar with possible molding problems and take care of them before they cause part failure. It is advisable to provide the screw hole with a slight taper or a recess. This not only facilitates screw engagement, but it signiﬁcantly reduces the stresses in the most critical area. On an injection molded thermoplastic thin walled boss, an additional problem may arise if the boss is ﬁlled in the longitudinal direction. Flow orientation reduces tangential elongation, therefore the use of self-tapping screws may not be possible. The stress level in the screw boss is highest immediately after assembly. Over time the stress decreases with relaxation. Molded bosses may still crack after

431

8.7 Metal Fasteners Method assembly, because the post-molding shrinkage forces have not been removed from the injection molded components. Self-tapping screws are one of the methods used to firmly join injection molded thermoplastic parts. With a hole in the thermoplastic boss, an initial thread tapping process can be eliminated, because either the thread forming or thread cutting screw itself taps the thread, moves cut material downwards, and simultaneously accomplishes a tight joint. Figure 8-37 shows the difference between thread forming and thread cutting screws.

Thread forming screw

Thread Forming Taper Lead Screws Thread forming screws deform the material into which they are driven and force it to conform to the screw thread. To reduce the risk of over-stressing, thread forming screws should have a taper lead as shown in Figure 8-38. Thread Cutting Screw Leads Thread cutting screws function as a tap to cut threads in a thermoplastic boss. Chips generated are pushed ahead of the screw and deposited at the bottom of the boss pilot hole. Figure 8-39 shows some types of thread cutting screw leads. As a general rule, thread cutting screws are capable of supporting higher pull-out forces and stripping torque than thread forming screws; however, they are more expensive. In applications that may cause screws to be loosened by motion and vibrations, special types of thread cutting screws are used. Thread cutting screws and leads are selected based on the application and the thermoplastic modulus of elasticity. The boss diameters and thread length of engagement are dependent on the size and type of thread cutting screw used.

Thread cutting screw

Figure 8-37 Thread forming and thread cutting screw assemblies

Taper lead "A"

Table 8-6 illustrates the relationship between these parameters for both types of self-tapping screws. For a same-size self-tapping screw with a different design geometry, the thermoplastic boss inside and outside diameter requirements are different. This is because the self-tapping screws of different types use speciﬁc geometries and technologies. Therefore, with various self-tapping screws the boss inside and outside diameters, and depth thread engagement are dependent on stiffness of the thermoplastic, screw geometry, screw size, molding conditions, and product design.

Taper lead "B"

Taper lead "AB"

Figure 8-38 Typical thread forming taper lead screws

Lead "BT"

Lead "F"

Lead "D"

Lead "BF"

Lead "T"

Lead "G"

Figure 8-39 Typical thread cutting screw leads

432

8 Thermoplastic Assembly Methods Table 8-6 Self-Tapping Screw Selection

Modulus of elasticity, K-psi

Self-tapping screw type

Thread length of engagement

Reassembly limits

Less than 200

Thread forming

80%

5 times

200–400

Thread forming Thread cutting

60% 90%

5 times

400–1,000

Thread cutting

90%

5 times

Over 1,000

Thread cutting

35%

None without insert

The pull-out force of a self-tapping screw assembled in a thermoplastic boss may be approximated by the shear strength of a cylinder having the same diameter as the screw pitch diameter and a length equal to the screw thread of engagement. Tests show that oil present on a self-tapping screw, washer, or thermoplastic boss to be assembled can cause a decrease in stripping torque of 50%. Pull-out Force Equation σS = 0.577 × σ, A = π × DP × L, F = σS × A F = 0.577 × σ × π × DP × L

(8-1)

Stripping Torque Equation T =

F × DP  p + π × μvDP  2  π × DP − μ × p 

(8-2)

Where: F = Pull-out force T = Stripping torque σS = Shear stress σ = Tensile stress A = Shear area DP = Pitch diameter L = Full thread length engagement μ = Coefﬁcient of friction p = Reciprocal of threads per unit length A length thread engagement of at least 2.5 times the self-tapping screw outside diameter will provide adequate pull-out strength. The depth of the boss inside diameter must be increased beyond the engagement value to include the partial threads in the lead of the self-tapping screw. For thread cutting screws, the depth of the boss inside diameter must also be compensated for length tolerances of both the screw and the hole, the hoop stress relief at the cavity, plus 0.2 the pitch diameters for chip clearance. Also, the draft angle for mold core holes should be a maximum of 1° to avoid excess hoop stress toward the bottom of the hole. Table 8-7 shows the most common self-tapping screw sizes used in both types of threads, including the different screw leads. This table provides the boss dimensions for thermoplastic materials of different modulus of elasticity.

Selftapping screw size

Screw major diameter [in]

Length thread engagement [in]

Depth of screw penetration [in]

Boss hole depth [in]

Boss hole diameter [in]

Thread forming

Thread cutting

Thread forming

Thread cutting

Thread forming

Thread cutting

AB

BT, BF

AB

BT, BF

A, B, AB

BT, BF, T, D, F, G

T, D, F, G

T, D, F, G

200 kpsi

400 kpsi

200 kpsi

400 kpsi

#1

0.073

0.182

0.278

0.236

–

0.295

0.262

–

0.064

0.066

0.061

0.065

#2

0.080

0.200

0.344

0.287

0.292

0.380

0.316

0.321

0.076

0.079

0.074

0.078

#3

0.099

0.247

0.398

0.330

0.338

0.414

0.361

0.368

0.089

0.091

0.086

0.090

#4

0.112

0.280

0.454

0.377

0.389

0.470

0.410

0.422

0.101

0.104

0.098

0.102

#5

0.125

0.312

0.510

0.431

0.418

0.526

0.467

0.454

0.114

0.117

0.110

0.115

#6

0.138

0.345

0.553

0.460

0.482

0.559

0.498

0.520

0.123

0.126

0.119

0.125

#7

0.151

0.377

0.604

0.498

–

0.620

0.540

–

0.136

0.140

0.132

0.138

#8

0.164

0.410

0.649

0.536

0.541

0.665

0.581

0.586

0.146

0.150

0.142

0.148

# 10

0.190

0.475

0.745

0.614

0.657

0.785

0.663

0.705

0.167

0.172

0.162

0.169

# 12

0.216

0.540

0.852

0.699

0.715

0.872

0.754

0.770

0.191

0.196

0.186

0.194

1/4 in.

0.250

0.625

0.988

0.776

0.837

0.988

0.837

0.898

0.221

0.216

0.224

0.233

5/16 in.

0.312

0.780

1.203

0.952

1.008

1.223

1.024

1.080

0.283

0.299

0.276

0.287

3/8 in.

0.345

0.862

–

1.094

1.187

–

1.178

1.271

0.347

0.384

0.340

0.351

8.7 Metal Fasteners Method

Table 8-7 Thermoplastic Boss Dimensions for Self-Tapping Screws

433

434

Table 8-8 Thermoplastic Boss Hole Diameter for Shakeproof Hi-Lo Screws

Hi-lo screw size

High thread diameter [in]

Low thread diameter [in]

Pitch diameter [in]

Boss pilot hole diameter [in] Thread forming modulus of elasticity

Thread cutting modulus of elasticity

Up to 200 Kpsi

200 to 400 Kpsi

Up to 200 Kpsi

400 and Up Kpsi

0.084–0.090

0.069

0.050–0.058

0.0670

0.0700

0.0700

0.0760

# 3–28

0.095–0.105

0.078

0.057–0.065

0.0730

0.0781

0.0760

0.0820

# 4–24

0.105–0.115

0.086

0.061–0.070

0.0810

0.0860

0.0820

0.0890

# 5–20

0.119–0.125

0.100

0.073–0.082

0.0935

0.0995

0.0960

0.1040

# 6–19

0.135–0.145

0.108

0.080–0.090

0.1015

0.1100

0.1060

0.1160

# 7–19

0.148–0.158

0.130

0.089–0.100

0.1200

0.1250

0.1250

0.1285

# 8–18

0.160–0.170

0.130

0.097–0.105

0.1200

0.1285

0.1250

0.1360

# 9–18

0.175–0.185

0.145

0.102–0.110

0.1360

0.1440

0.1405

0.1495

# 10–16

0.185–0.195

0.145

0.108–0.118

0.1360

0.1440

0.1405

0.1520

# 11–16

0.198–0.210

0.150

0.113–0.125

0.1495

0.1562

0.1520

0.1695

# 12–16

0.210–0.220

0.167

0.125–0.137

0.1570

0.1660

0.1610

0.1770

# 13–16

0.220–0.230

0.180

0.127–0.139

0.1695

0.1800

0.1730

0.1875

1/4–15

0.250–0.260

0.200

0.161–0.175

0.1890

0.2010

0.1960

0.2187

8 Thermoplastic Assembly Methods

# 2–32

435

8.7 Metal Fasteners Method

High thread

Low thread

30˚ Included angle

60˚ included angle

Figure 8-40 Shakeproof® Hi-Lo screw and details of both threads

Shakeproof® Hi-Lo Thread Forming and Thread Cutting Screws The Shakeproof® Hi-Lo screw has a unique thread form. It was designed to improve the screw and the thermoplastic material performance. The screw has a double lead consisting of a high and a low thread. The high thread is ﬂatter and sharper than a conventional screw thread, having a 30° included angle as compared to the conventional 60° included angle. The low thread has the 60° included angle and has a height ranging from 40–50% of the high thread height. The 30° included angle of the high thread form reduces the hoop stress by 50% of the value generated by the conventional 60° screw. Boss cracking in thermoplastic is dramatically reduced and smaller diameter bosses can often be speciﬁed. The Hi-Lo thread design has a smaller minor diameter than a conventional screw. The high threads make a deeper cut into the boss wall, leaving a greater volume of material between the threads. There is also a greater amount of material in contact with the high, ﬂat thread; the boss axial shear area is increased, providing a higher pull-out strength. The 30° included angle of the high thread displaces less material when it is driven into a thermoplastic boss. thus requiring lower driving torques. A greater amount of material remains between the high threads, increasing the stripping torque. Low driving torque and high stripping torque provide maximum protection against stripping problems. Figure 8-40 shows a Shakeproof® Hi-Lo screw and a close view of both types of threads. Table 8-8 shows the most common Shakeproof® Hi-Lo thread forming and thread cutting screw sizes. This table provides the boss dimensions for thermoplastic materials with a modulus of elasticity lower than 200,000 psi, between 200,000 to 400,000 psi and higher than 400,000 psi. Plastite® Tri-Lobular Twin Lead Thread Forming Screw The Plastite® self-tapping screw is a unique tri-lobe shaped thread forming screw, designed speciﬁcally for thermoplastic materials. It combines the trilobular cross sectional form with a deep spaced thread. The tri-lobular principle provides three sharp swaging lobes with full radial relief of the thread form. The lobular section eliminates wedging, thus reducing the hoop stress, and provides effective resistance to vibrational loosening by allowing the material to recover and ﬁll in behind the lobes.

436

Next

Table 8-9 Thermoplastic Boss Hole Size for Plastite® Tri-Lobular Screw

Plastite® screw major diameter [in]

“C”, diameter of circumscribing circle [in]

“D”, diameter measured across the center [in]

Minimum screw out of round tolerance [in]

Recommended boss pilot jole diameter [in]

Length thread engagement [in]

# 2–28

0.089

0.086–0.092

0.083–0.089

0.002

0.076

0.500

# 3–24

0.106

0.104–0.110

0.100–0.106

0.002

0.089

0.500

# 4–20

0.123

0.121–0.127

0.117–0.123

0.002

0.102

0.750

# 6–19

0.143

0.141–0.147

0.137–0.143

0.003

0.120

0.750

# 7–18

0.160

0.160–0.166

0.154–0.160

0.004

0.136

1.250

# 8–16

0.179

0.179–0.185

0.173–0.179

0.004

0.154

1.250

# 9–15

0.193

0.193–0.199

0.187–0.193

0.004

0.166

1.250

# 10–14

0.208

0.206–0.212

0.202–0.208

0.004

0.180

1.250

# 12–11

0.230

0.229–0.235

0.224–0.230

0.005

0.199

1.250

# 12–14

0.226

0.226–0.232

0.220–0.226

0.005

0.203

1.250

1/4–10

0.268

0.270–0.276

0.262–0.268

0.006

0.238

1.250

5/16–9

0.335

0.335–0.345

0.325–0.335

0.006

0.281

1.250

8 Thermoplastic Assembly Methods

Plastite® tri-lobular twin lead screw size

437

The threads are considerably coarse, up to 30% fewer threads than comparable sizes of self-tapping type “B” screws and similar coarse-threaded self-tapping screws. This condition offers a heavier shear stress in the thermoplastic boss, less material displacement, and improved holding power. The increased thread depth provides deeper thread engagement, greater tolerance for the pilot hole, and freedom from the dangerous wedging action and friction of root interference, which commonly causes bursting of thin-walled bosses.

Semi cup point

8.8 Press Fitting Method

1.2 Taper thread 48˚ Twin lead

The Plastite® screw has a single lead thread for starting and alignment. The point taper is limited to 1.2 threads to gain additional engagement. The Twin Lead Plastite® thread forming screws are designed to meet the most critical requirements for self-tapping screws in thermoplastics, which is resistance to stripping torque and pull-out force. The screw size designation combines the commonly used numerical screw size with the major diameter “D” as measured with ordinary micrometers, followed by the number of threads per inch.

D C

Figure 8-41 shows a general drawing of a Plastite® tri-lobular twin lead thread forming screw. Table 8-9 shows the Plastite® tri-lobular twin lead screw sizes. This table recommends the thermoplastic boss pilot hole diameters and the length thread engagement. The boss sizes are subject to variation, depending on depth of engagement, the stripping torque, pull-out force, ease of driving, and the thermoplastic material modulus of elasticity required for the application.

8.8

Figure 8-41 Plastite® tri-lobular twin lead screw

Press Fitting Method

Press ﬁtting assembly methods are universally applicable for,joining different components made from the same material, other thermoplastic resins, or metals. The press ﬁtting method does not require welding equipment or any foreign element such as adhesive, sealer, cement, mechanical expander, or metal insert. Properly applied, this assembly method produces serviceable assembled joints with good strength at minimum cost. Figure 8-42 and 8-43 show two press-ﬁt light-duty assembly applications. The press ﬁtting assembly method for injection molded thermoplastic parts is similar to the press ﬁtting process used by the metal industry to join different components. The clearance or interference dimensions (including tolerances) used for press ﬁtting thermoplastic components are generally greater than the dimensions required to press ﬁt metals. Higher interference dimensions are necessary to compensate for the lower tensile strength, modulus of elasticity, and creep effects of the thermoplastic materials compared with metals. However, the elongation property of thermoplastics is higher than that of metals; therefore, injection molded thermoplastic components can be designed using higher strain ratios than the values recommended for metals. The strain ratios used for designing injection molded thermoplastic components are dependent on the product design, the modulus of elasticity, the ambient temperature, and moisture conditions. For maximum joint strength in thermoplastics, the interferences should reﬂect product design, mechanical properties of the resin, molding process conditions, mold ejection system used to remove the molded parts from the mold (without molded-in stresses), and the assembly process conditions (join force and velocity). All these parameters are required to obtain a good press ﬁtted joint without stressing the thermoplastic component beyond

Plastic handle

Metal shaft Figure 8-42 Press-fit handle

Press fit gauge

Molded plastic parts

Figure 8-43 Tire pressure gauge

438

8 Thermoplastic Assembly Methods its yield point. The theoretical relationship between interference and stress level is based on the geometry of the part and the mechanical properties of the resin. Interference can be calculated by using standard stress analysis procedures. Residual joint strength in press ﬁtted injection molded thermoplastic parts is affected by complex variables, such as modulus of elasticity and coefﬁcient of friction. For most thermoplastic resins, the variations in modulus of elasticity become negligible after one year and the joint strength is constant. Because the coefﬁcient of friction is affected by lubrication, moisture, temperature, and hoop stress, the coefﬁcient of friction under each of these conditions must be known in order to calculate accurately the strength of the press ﬁt joints. When the torsional strength is critical, ribs should be provided around the boss to distribute the load around the shaft. When the hoop stress is critical, circular ribs should be used. When both torsional and hoop stress are critical, a coarse diamond knurled shaft or combination of grooves for the shaft and ribs around the hub provide a good and well balanced structure. Plastic sleeve (Hoop stress) Load

Internal metal Ttube

When injection molded thermoplastic parts are to be press ﬁtted for maximum holding power immediately after molding, they should be free from molded-in stresses. Molded-in stresses can be reduced by annealing the molded part under controlled conditions, recommended by the resin supplier. Also in designing such joints, environmental conditions should be carefully considered. Expansion by heat and moisture can be compensated for by designing for expected growth in the worst conditions. These dimensional changes should be adjusted in addition to the press ﬁt interference selected for the desired joint strength. Molded-in stresses in a press ﬁtted part may tend to promote crazing in some amorphous thermoplastics, such as acrylics, polycarbonate, and polystyrene, reducing the impact strength of these resins. When a molded part is subjected to repeated impact loads, an impact test method similar to the end use operating conditions should be run to determine the performance of the press ﬁt joint.

Poor design Plastic sleeve (Failure) Load

Figure 8-44 shows how the thermoplastic and metal press ﬁtting joint designs affect the performance and structure of the molded thermoplastics parts, causing failure of the thermoplastic products under extreme high loads,. Press ﬁtting the hubs into the shafts is not limited to only metal parts. Combinations of plastic-to-plastic and plastic-to-metal parts offer dependable, pressure tight joints at moderate costs.

Operational problems Plastic tube (Compression)

Press seal External metal tube

Recommended design Figure 8-44 Thermoplastic and metal press fit joint design

Shrink ﬁtting is one of the assembly methods for joining molded parts without interference dimension problems. Interferences for shrink ﬁtting are determined by adding the thermoplastic mold shrinkage rate to the hub, shaft diameter, and shaft thermal expansion. The shrink ﬁtting assembly method takes advantage of volume reduction (mold shrinkage) properties of the molded part immediately after molding (cooling stage), while the molded part is still hot (almost the same temperature of the mold cavity). At this time, the hot molded thermoplastic part continues its cooling process, the polymer geometry structure springs back (polymer volume expansion during plastiﬁcation) to its original structure (organized, oriented, reduced, and small). This rate of expansion and contraction of the polymer is known as the mold shrinkage rate. The hot molded part immediately after molding has already recovered more than 50% of its shrinkage rate during its ﬁrst cooling stage inside the mold cavity. The remaining mold shrinkage (ﬁnal size reduction) of the hot thermoplastic part continues when the part is ejected from the mold, until the part cools off at room temperature.

439

8.8 Press Fitting Method The shrink ﬁtting process begins when the hot part (immediately after molding) is pressed into the shaft (pre-heating the shaft to the same temperature of the part immediately after molding is recommended). The hot hubs and shafts are then placed in a well ventilated area for cooling and aligned so that both sides of the hub are being cooled at the same rate. Do not apply any mechanical stress or allow cooling of the hub (joint assembly side mounted, lay down on one side). For shrink ﬁtting, the shaft may be cooled or the hub may be heated to reduce the interference. The required temperature change may be calculated by the usual method, using the coefﬁcient of linear thermal expansion for the resins involved.

8.8.1

Press Fitting Interference

Because of the lower modulus of elasticity of thermoplastic materials (as compared to metals), interferences required for press ﬁtting thermoplastic components are much looser. The tolerances for thermoplastic parts are usually wider than those for similar metal parts. The amount that the thermoplastic hub inside diameter may be expanded to accommodate an interference ﬁt shaft is generally limited by the tensile strength of the thermoplastic hub in the hoop direction. Shaft and Hub Made of Different Materials The relationship of the design stress in a hub to the diametral interference of a press ﬁt shaft is: Int. =

σ d DS  GF + νH 1 − νS  + GF  EH ES 

(8-3)

where GF is a geometry factor obtained from: GF =

1 + (DS / DH )2 1 − (DS / DH )2

(8-4)

and where the design stress, σd, equals the yield stress σt divided by a safety factor. Small safety factors can be used (1.0 to 3.0), because creep will reduce the stress level from maximum stress, which occurs only at assembly. Shaft and Hub Made of the Same Material When both parts are of the same material (EH = ES), Eq. 8-3 simpliﬁes to: Int. =

σ d DS  GF + 1  E  GF 

(8-5)

Metal Shaft and Plastic Hub When the shaft is of metal or other high modulus of elasticity material, the last term in Eq. 8-3 becomes negligible and the equation simpliﬁes to: Int. =

σ d DS  GF + νH  EH  GF 

(8-6)

440

8 Thermoplastic Assembly Methods 180

110

150

MDPE shaft

90

LDPE shaft

120

70

Steel shaft 90

0

0.2

Steel shaft 0.4

0.6

0.8

50

1.0

65

HDPE shaft 55 45

Steel shaft 35

0

0.2

0.4

0.6

0.8

1.0

High-density polyethylene hub inside diameter 55

Acrylic shaft

45 35

Steel shaft 25

0

0.2

0.4

0.6

0.8

1.0

Acrylic hub inside diameter 100

0.2

0.4

0.6

0.8

1.0

60

PP shaft 50 40

Steel shaft 30

0

0.2

0.4

0.6

0.8

1.0

Polypropylene dub inside diameter 50

Acetal shaft 40 30

Steel shaft 20

0

0.2

0.4

0.6

0.8

1.0

Acetal homopolymer hub inside diameter 50

80

TFE shaft

40

Nylon shaft

60 40

0

Medium-density polyethylene hub inside diameter Diametrical interference limit, (0.001 per inch of shaft diameter)

Diametrical interference limit, (0.001 per inch of shaft diameter)

Low-density polyethylene hub inside diameter

30

Steel shaft

Steel shaft 0

0.2

0.4

0.6

0.8

Unreinforced nylon 6/6 hub inside diameter

1.0

20

0

0.2

0.4

0.6

0.8

1.0

TFE (fluorocarbon) hub inside diameter

Figure 8-45 Press fit interference

The curves of Figure 8-45 can be used as guides for determining interference limits for press ﬁt applications for a number of material combinations. Values shown represent theoretical interference limits for press ﬁts based on yield point (σt) and modulus of elasticity data at room temperature. These limits should be reduced by a safety factor required for the application. Metal Hub and Plastic Shaft When the hub of an assembly is made of a high tensile strength, high modulus of elasticity material such as steel, and the shaft is made of a thermoplastic material, the strength of the shaft may be limiting. Here, failure will occur in compression, and Eq. 8-3 becomes less reliable. However, with a small safety factor (between 1 and 2) for the design stress, the equation: Int. =

σ d DS  GF + νS  ES  GF 

provides a reasonable approximation of allowable interference.

(8-7)

441

8.8 Press Fitting Method

8.8.2

Circular Press Fitting Assembly Method

Press "OFF"

Round thermoplastic parts can be assembled by press ﬁtting. To calculate the force required to press ﬁt a shaft into a hub, the use of the following equation is needed for the analysis. F = π μ P DS L

(8-8)

Where: F = Assembly force (lb) P = Joint pressure = σd / GF (psi) DS = Shaft diameter (in) L = Joint length (in) μ = Coefﬁcient of friction σd = Design stress (psi) GF = Geometry factor (Eq. 8-4)

Housing shaft mount on fixture Shaft holding drive sleeve Press "ON"

The coefﬁcient of friction depends on many factors and varies considerably from application to application. Table 8-10 may be used as an approximation. When greater accuracy is important, end use testing is recommended. Figure 8-46 shows some illustrations of press ﬁt assembly equipment.

Heading tool

Table 8-10 Hub/Shaft Coefficient of Friction (μ)

Plastic to plastic (molded surface)

0.25–0.35

Plastic to steel (ﬁne surface ﬁnish)

0.15–0.25

Plastic to steel (rough surface ﬁnish)

0.35–0.65

Plastic to steel (polished, hard, lubrication)

0.05–0.15

Press fit joint completed

Figure 8-46 Press fitting assembly equipment

Press Fitting Joint Strength Axial strength resulting from an interference press ﬁt of a joint after a period of time may be approximated using Eq. 8-8. On a smooth joint surface shear area, creep and/or stress relaxation reduce the effective interference that, in turn, reduces joint pressure and holding power.

To deter the effect of creep on an injection molded thermoplastic hub press ﬁtted into a metal shaft to form a joint, the metal shaft joint surface can be ﬁne diamond knurled or roughened. The force required in the press ﬁt joint process generates a frictional energy inside the thermoplastic hub shear area, melting the polymer joint surface, causing the molten thermoplastic to move or producing a cold ﬂow effect into the knurled grooves and valleys of the metal shaft. The result is an effective increase in coefﬁcient of friction that can compensate for loss of joint pressure caused by creep effects.

450

Force to separate the joint, (lb)

Effects of stress relaxation on an interference ﬁt between a hub and shaft made of acetal homopolymer resin were examined by Du Pont to demonstrate the creep effects caused on the force to separate the press ﬁt joints for extended time periods. Figure 8-47 shows the test result data on the joint strength versus time for an interference press ﬁt. These curves show a reduction of separation force with time for a 0.011 in initial interference press ﬁt of a 0.50 in outside diameter, 0.25 in inside diameter hub on a 0.25 in diameter shaft, and a joint length of 0.50 in. The tests were performed at 73 °F and 50% relative humidity conditions.

Machined parts, "

0.25

350

250 Molded parts, "

150 0

102

0.15 10

3

10

4

Time, (hours)

Figure 8-47 Joint strength vs. time for an interference press fit

10

5

442

8 Thermoplastic Assembly Methods Dimensional Changes of the Press Fit Joint When a shaft is press ﬁtted into a thermoplastic hub, both outside and inside diameters of the hub are expanded. Similarly, the inside diameter of a hub shrinks when the hub is pressed into a housing. The relative dimensional changes for each of these cases are important, because these changes affect the relative joint strength of the components, subsequently causing assembly quality control problems. The inside diameter reduction of a thermoplastic hub due to an external interference with an essentially rigid housing can be found from: dH (Reduction) =

DS2

2 × DS × DH (1 + νH ) + DH2 (1 − νH )

(8-9)

where νH is Poisson’s ratio for the hub material. The increase in outside diameter of a thermoplastic hub due to an internal interference with a rigid shaft is found from: DH (Increase) =

2 × DS × DH DS2 (1 − νH ) + DH2 (1 + νH )

(8-10)

Changes in temperature and moisture content also affect dimensions of a thermoplastic part. These dimensional changes are relatively unimportant for plastic-to-plastic joints; the differential expansion is not present to any great degree. However, they must be considered in the design of metal-to-thermoplastic joints that operate over a temperature range. For example, as a thermoplastic hub, ﬁtted over a metal shaft, is cooled from room temperature; the plastic hub contracts more than the metal shaft and the effective interference increases. If the temperature is lowered far enough, effective interference increases to the point where the stress level may exceed the yield point of the material and permanent deformation or failure may occur. In raising the temperature of this joint, the thermoplastic hub expands more than the metal shaft, resulting in a lower effective interference and causing a decrease in joint strength. Examples 8-1 to 8-4 use DS DH dH E F μ Int. L P σd σt T t GF α ν

= Shaft diameter (in) = Hub outside diameter (in) = Hub inside diameter (in) = Modulus of elasticity (psi) = Assembly force (lb) = Coefﬁcient of friction = Diametral interference in the joint (in) = Joint length (in) = Joint pressure (psi) = Design stress (psi) = Tensile stress (psi) = Torque (lb/in) = Temperature (°F) = Geometry factor = Coefﬁcient of linear thermal expansion (in/in/°F) = Poisson’s ratio

443

8.8 Press Fitting Method

Example 8-1 Find the allowable interference of a nominal 1.00 in outside diameter hub of acetal homopolymer resin that is to be ﬁtted on a 0.50 in shaft of the same material. Assume an average service environment of 73 °F and 50% RH. For these conditions, the yield strength, σt, of acetal homopolymer resin is 10,000 psi and the modulus of elasticity is E = 410,000 psi. Using a safety factor of 2, σd = 10,000/2 = 5,000 psi. From Eq. 8-4, GF =

1 + (DS / DH )2 1 − (DS / DH )2

=

1 + (0.50/1.0)2 1 − (0.50/1.0)2

= 1.67

Substituting these values in Eq. 8-5, Int. =

5,000 × 0.50  1.67 + 1    = 0.0097 in 410,000  1.67 

Therefore, the shaft diameter can be 0.0097 in larger than the hub inside diameter without danger of hub failure during assembly.

Example 8-2 Determine the allowable interference for an acetal homopolymer hub with a outside diameter (DH) of 1.00 in and inside diameter (dH) of 0.50 in, if a steel shaft is used (for the hub material, νH = 0.35). Substituting into Eq. 8-6, Int. =

5,000 × 0.50  1.67 + 0.35    = 0.0073 in 410,000  1.67

Example 8-3 Estimate the holding power of the assembly in Example 8-1 initially and after one year, assume constant environment. At the time of assembly, P =

σd 5,000 = = 3,000 psi GF 1.67

and from Eq. 8-8 (μ = 0.2), F / L = π × 0.2 × 3,000 × 0.5 = 950 lb/in of joint holding power. After one year, the modulus of elasticity of acetal homopolymer resin drops from 410,000 to 190,000 psi. Therefore, σ d = 5,000 ×

190,000 = 2,317 psi 410,000

and P = 2,317 / 1.67 = 1,387 psi. Holding power after one year is then: F / L = π × 0.2 × 1,387 × 0.5 = 435.88 lb/in of joint holding power.

444

8 Thermoplastic Assembly Methods

Example 8-4 Calculate the axial press ﬁt interference for a 0.25 in diameter brass shaft in a hub made of acetal homopolymer. The design requires a tight seal during temperature cycles from –30 to +200 °F. The seal must remain effective for 10 years, EB (brass) = 13 × 106; EA (acetal) = 4.10 × 105 at assembly, and 1.75 × 105 after 10 years; αB = 10.4 × 10–6; αA = 5.5 × 10–5; σA = 10,000 psi; DS = 0.25 in; DH = 0.50 in; νB = 0.333; νA = 0.35. Assembly is done at room temperature. From Eq. 8-4, GF = 1.67. Maximum interference that can be tolerated at room temperature (letting σd = σA) is: Int.max =

10,000 × 0.25  1.67 + 0.35 1 − 0.333  +   = 0.0074 in 1.67 410,000 13 × 106 

To form a tight seal throughout the service temperature range, a large axial interference is necessary. Large interferences with small safety factors are needed, because the maximum stress occurs at installation and diminishes with time. Effective axial interference after 10 years is: Int. = (0.0074)

175,000 = 0.0031 in 410,000

Because the acetal hub expands more than the brass shaft at 200 °F, interference is reduced. ∆Int. = DS (αA − αB ) (t 2 − t1 ) = 0.25 (5.5 × 10−5 − 10.4 × 10−6 ) × (200 − 70) = 0.0015 in and from –30 to +70°°F the increase in interference is 0.0011 in. The effective interferences at these extreme temperatures after 10 years are: : At 70 °F after molding = 0.0074 in : At 70 °F after 10 years = 0.0031 in

: At 200 °F = 0.0031–0.0015 = 0.0016 in

: At –30 °F = 0.0011 + 0.0031 = 0.0042 in

8.9

Snap Fitting Methods

Snap ﬁtting provides a simple, economical, and rapid means of attaching thermoplastic parts to other plastic materials or metals. A snap ﬁt is strong but is not usually pressure tight, unless other features such as a compression seal are incorporated in the joint design. Designers can dramatically reduce production and component costs in fastening and assembly of both plastic and nonplastic parts by taking advantage of the design possibilities of thermoplastics. The flexibility and “springiness” of thermoplastics permit simple, efﬁcient designs. In addition, the ability to mold fasteners or joint members integrally with a thermoplastic part saves both hardware and extra assembly operations.

445

8.9 Snap Fitting Methods Snap ﬁts can be divided into two distinct categories. The ﬁrst is the complete circular snap internal or external undercut that is ejected by snapping the part from a mold spring loaded core, thus requiring the correct strain rate of the material to elongate the molded part for removal from the mold. The second is formed of individual ﬂat or annular cantilevered latching beams.

Bevel gear

Snap Retainer

Bearing

8.9.1

Circular Undercut Snap Fitting Joints

Injection molded thermoplastic products can have undercuts for functional reasons or for decorative effects. When undercuts in a chamfered molded product are less than the recommended wall thickness and strain for the undercut, it may be possible to strip or eject the product out of the mold cavity. However, when this is contemplated, the mold must be designed to ensure that ejection takes place only when the thin walled molded product is free to expand or compress. In such cases, it may be necessary to provide the mold with an ejection ring or plate, rather than ejector pins. The ejection of an undercut from the mold usually presents no problem, as long as the part and the mold are correctly designed and dimensioned. Because the resin is still at a relatively high temperature, the modulus of elasticity is low and the elongation high. Often, an undercut that has been ejected correctly from the mold, will crack during the assembly operation. This may be caused by weak spots, such as weld lines, gates, or voids. Therefore, a gate should never be placed close to an undercut. This type of breakage can occur even a long time after the assembly has been made due to excessive post mold shrinkage.

Plastic guard

Shaft

Figure 8-48 Gear snap fitting Snap Spur gear

Shaft guard

Shaft Gear guard Snap

Figure 8-49 Shaft gear circular

There are nevertheless a great number of applications where undercut snap ﬁts have proven to be simple, inexpensive, and reliable. Figures 8-48 and 8-49 show typical applications where these circular undercut snap ﬁttings are used. Figure 8-50 shows a poor snap ﬁt ball joint design; the snap logs are too rigid to strip from the mold or to install the ball joint. We recommend using four uniform wall thickness snap logs and 30° taper entrance without sharp corners. Figure 8-51 shows the vertical wall restricting the joint expansion, making it difﬁcult to snap ﬁt the shaft. Relocating the wall to the back side is recommended.

Poor design Failure

Operational problems

Failure Wall restricts snap elongation

Poor design

Operational problems

Figure 8-51 Circular snap fitting joint designs

Recommended design

Recommended design

Figure 8-50 Circular snap fitting ball joint designs

446

8 Thermoplastic Assembly Methods

8.9.2

Suggestions for Stripping Circular Undercut Snap Fitting

• Undercut thermoplastic product must be free to stretch or compress. • The undercut should be rounded and chamfered to permit easy slippage of the thermoplastic product over the core of the mold and to minimize stress concentration during the ejection stripping action. • Adequate mold ejection contact area should be provided to prevent penetrating or collapsing of the thin-walled thermoplastic product during the stripping action. • Some permanent deformation may occur when the undercut is stripped. The deformation depends on the thermoplastic product, mold design, and injection molding process variables. • Acetal homopolymer resins: It is possible to strip injection molded acetal homopolymer products from the cavities, see Figures 8-52 and 8-53, if the undercuts are less than 5% of the diameter and have a 30° taper. • Unreinforced nylon 6/6 resins: Injection molded nylon 6/6 products having between 6% to 10% undercut of the diameter and a 30° taper can be stripped from a mold. To calculate the allowable undercuts (that vary with thickness and diameter) see Figures 8-52 and 8-53. • Unreinforced resins: Thermoplastic resins having high elongation properties could produce circular internal undercut snap ﬁt components by using the general design guidelines shown in Figure 8-54. To calculate the allowable undercut for a speciﬁc material, the dimensions given by Table 8-11 should be checked by using the permissible strain and the crystalline structure of the material in question.

Allowable internal undercut equation 30˚

45˚

30˚

(A - B ) x 100

D

% U n d er c u t =

B

B

B

A

A

Figure 8-52 Circular internal undercut snap fitting joints

4t

Allowable external undercut equation (F - D ) x 100 30

% U n d er c u t =

˚

B t

Figure 8-54 Circular internal undercut snap fitting joint

C

30˚ Chamfer

C 30˚ Chamfer

C

D

D

F

F

Figure 8-53 Circular external undercut snap fitting joints

447

8.9 Snap Fitting Methods Table 8-11 Circular Internal Undercut Depth

D (in) Outside diameter

B (in) Inside diameter

t (in) Undercut depth

0.197 0.315 0.394 0.433 0.669 0.866 1.102 1.299

0.078 0.118 0.157 0.197 0.394 0.590 0.787 0.984

0.0019 0.0019 0.0039 0.0039 0.0078 0.0135 0.0196 0.0255

8.9.3

Figure 8-55 Cantilever conical latches snap fit permanent

Cantilevered Latch Snap Fitting Joint

The second category into which snap ﬁts can be classiﬁed is based on cantilevered latches, the retaining force of which is essentially a function of bending stiffness. They are actually special spring applications, which are subjected to high bending stress during assembly. Under working conditions, the lugs are either completely unloaded for moving parts or permanently distorted in order to achieve a tight assembly. Snap ﬁts provide a strong and reliable fastening. Cantilevered latch snap ﬁts are usually molded from polypropylene, polyethylene, nylon, acetal, polycarbonate, or thermoplastic elastomers. These materials are particularly useful because, being somewhat elastic, they permit tighter, more secure snap ﬁt joints that are dependent on friction. Changes in temperature and moisture content affect dimensions and the holding friction.

Figure 8-56 Cantilever round logs snap fit ball removable

Cantilevered latches should always be designed in a way so as not to exceed allowable stresses during assembly operation. Too short a bending length may cause breakage, which can be corrected by using considerably longer ﬂexible lugs. Cantilevered snap ﬁt latches should always be dimensioned to obtain a constant stress distribution over the whole length. This can be achieved by providing a slightly tapered beam. Special care must be taken to avoid sharp corners and other possible stress concentrations. Injection molded cantilevered latches are the most common type of devices used for assembling thermoplastic components. The latching mechanism is simple: a lip or ball at the end of a springy lever engages a lip or socket built into the mating surface. The design requires a balance of stiffness and ﬂexibility, with enough resilience to withstand the deﬂection during latching and unlatching without damage. Poor design can result in permanent set from over stress or cracking at the base of the lever.

Figure 8-57 Cantilevered annular snap fit latches

Figures 8-55, 8-56, 8-57, 8-58, 8-59, and 8-60 show typical injection molded thermoplastic cantilevered latch snap ﬁt designs.

Figure 8-60 Device with two cantilever snap fit latches

Figure 8-59 Mounted switch by cantilever snap fit latches

Figure 8-58 Pressure valve, snap fitted cantilever latch, and “O” ring

448

8 Thermoplastic Assembly Methods Cantilever thin rectangular latches Cantilever taper latches

The several design possibilities have been reduced to a few basic shapes and calculation principles have been developed for the following basic designs: cantilevered snap ﬁt latches, torsion snap joints, and cantilevered annular snap ﬁt latches. With cantilevered snap ﬁt latches, the load is mainly ﬂexural. In torsion snap joints, shear stresses carry the load. Cantilevered annular snap ﬁt latches are rotationally symmetrical and involve multi-axial stresses. Figure 8-61 shows an injection molded thermoplastic device with four cantilevered snap ﬁt latches that hold the mating component ﬁrmly in place, yet allowing the device to be removed when necessary. An economical and reliable snap ﬁt joint can also be made by using rigid lugs on one side, in combination with cantilevered snap ﬁt latches on the other side. This design is particularly effective for joining two similar halves of a housing that need to be easily separated. Figure 8-62 shows how these four latches work.

Figure 8-61 Plastic device with four cantilevered snap fit latches Flexible latches

Figure 8-63 shows a poorly designed cantilevered snap ﬁt latch with sharp corners at the base. This causes breakage of the snap lever base during loading the component at assembly. The recommended design shows the cantilevered snap ﬁt latch with a spring like ﬂexing beam without stress concentration. Figure 8-64 shows a poor design of dual cantilevered snap ﬁt latches that are connected, making the structure rigid and unable to bend. The recommended design has two separated cantilevered snap ﬁt latches connected at the base with a generous ﬁllet radius.

Rigid lugs

Figure 8-62 Combination of two pairs of snap fit latches

Figure 8-65 shows a circular device with cantilevered annular snap ﬁt latches for a permanent assembly; these latches have a 90° stop to lock the round cover in place and a 30° chamfer at the base to help at assembly. The cantilevered annular latches are ﬂexible, round, uniform width, and proportionally distributed, which is required for assembly.

Sharp corner

Latch support prevent bending

Poor design Failure

Recommended design

Poor design

Figure 8-64 Dual cantilevered snap fit latch designs Operational problem

Frame Round cover Recommended design

Figure 8-63 Cantilevered snap fit latch designs

90˚

Cantilever latch

Figure 8-65 Cantilever annular snap fit latches, permanent

449

8.9 Snap Fitting Methods Figure 8-66 shows a round plug with cantilevered annular snap ﬁt latches used in applications requiring the plug to be removable; these latches have a 45° slope to hold the device in place and a 30° chamfer at the base to help at assembly. Figure 8-67 shows six cantilevered annular snap ﬁt latches that are very ﬂexible, rounded with uniform width, proportionally distributed, with a generous ﬁllet radius between each latch lower section; all these details are required for assembly. Figure 8-68 shows a heater valve cover secured in place by cantilevered annular snap ﬁt latches.

Round plug

Frame

Cantilever latch

45˚

Figure 8-66 Cantilever annular snap fit latches, removable

Figure 8-69 shows an injection molded acetal homopolymer split worm gear. The two worm gear halves are designed in such a way that each component is injection molded using the same mold cavity. The two worm gear halves are post-mold assembled by rotating 180°, facing each other, aligning the two pins with the holes, and pressing both components until they are snap ﬁtted together in four locations. The four cantilevered annular snap ﬁt latches provide a strong coupling force with good annular alignment in the assembly of both worm gear halves.

8.9.4

Cantilever Snap Fit Latch Design Guidelines

The most economical method of assembling thermoplastic components is with cantilever snap ﬁt latches. While simple in appearance, snap ﬁts require complete design analysis to achieve functional performance and structural stability. Cantilever snap ﬁts have been successfully applied in various industries from toys to automotive, appliances to TV, and electrical engineering construction, and business machines. Equally important to the design engineer is the selection and evaluation of the proper thermoplastic resin.

Figure 8-67 Plastic device with six cantilever annular snap fit latches

Valve cover

In all the cantilever snap ﬁt latch designs, it is assumed that one of the mating parts remains rigid. If the two components are of approximately equal stiffness, half the deﬂection can be assigned to each part. If one component is more rigid than the other, the strength available can be utilized to the fullest. The design goal is to develop a cantilever latch geometry for the speciﬁc resin being used, that can withstand and recover from the strain produced on the cantilever latch during deﬂection. That strain is usually calculated from the standard cantilever beam equation for a rectangular or taper cross section beam. Finite element structural analysis studies show that the stresses caused by the deﬂection tend to concentrate at the root section of the cantilever beam (the point where it joins the structural support wall) and can produce cracks at this section from repeated loading operation of the beam. Generous root ﬁllet radius will help with the stress concentration, additionally a tapered wall thickness will help to uniformly distribute the peak stress level on the beam by 25% or more. An important objective is to prevent the cantilever beam from taking a permanent set, or retaining residual deﬂection after bending. How much, if any, residual deﬂection is produced and affected by the ratio of beam length to its wall thickness? The smaller the ratio, the greater the strain the beam can withstand without taking a permanent set. The amount of stress that can be tolerated up to the yield point depends not only on the basic polymer, but also on the additives used for compounding the thermoplastic resins.

Heater pipe

Figure 8-68 Heater valve cap with annular snap fit latches

Figure 8-69 Split worm gear assembled with four annular snaps

450

8 Thermoplastic Assembly Methods Sharp corner causes high stress concentration

Figure 8-70 shows how a tapered cross section cantilever beam moves the stresses away from the root area compared with a rectangular uniform beam.

Force

8.9.5

Uniform cantilever beam Taper beam and radius distribute stress uniformly

e Forc

Radius = 60% H

H Taper cantilever beam Figure 8-70 Cantilevered snap latch beams, design differences

Cantilever Latch Snap Fit Mathematical Model

A wide range of design possibilities exists for cantilever snap ﬁt latch joints. In view of their high level of ﬂexibility, injection molding thermoplastic resins and a good product design are ideal for this assembly method. Figure 8-71 shows a mathematical model for a cantilever latch beam with tapered width and length used in the developing of equations. The designs of cantilever snap ﬁt joints take into account the fact that, for brief periods, they are subjected to very high mechanical loads. This means the stress/strain behavior of the material is already outside the linear elastic range and the modulus of elasticity must be replaced by the strain-dependent secant modulus. The snap ﬁt joints are basically simple cantilever snap latch beams (see Table 8-12), that have rectangular, trapezoidal, or annular cross sections. It is recommended that the cantilever beams be designed so that either the wall thickness or the width is tapered from the root to the snap latch with a value at the end of 50% of the value at the root. This design technique allows the load bearing cross section (at any point) to distribute the stress uniformly. The maximum strain on the injection molded thermoplastic beam is reduced and less thermoplastic material is required for the application. Notes: 1. If the tensile stress occurs in the convex surface, use C2 in Figure 8-79 to determine C1and use C2 if it occurs in the concave surface; use C1 accordingly. 2. (c) is the distance between the outer ﬁber and the center of gravity (a neutral axis) in the surface subjected to tensile stress. 3. The section modulus should be determined for the surface subject to tensile stress. Section modulus for an annular cross section is given in Figure 8-80; use the nomogram to determine the annular modulus (Z2). The section modulus for other basic geometrical shapes can be found in mechanical engineering manuals.

B 1 (Width)

F (Force)

H 2 (Wall)

H 1 (Wall)

(Deflection) L (Length)

Figure 8-71 Cantilever snap fit latch mathematical model

B 2 (Width)

451

8.9 Snap Fitting Methods Table 8-12 Equations for Cantilever Beams

Cross section

Rectangle c

Trapezoid c2

H

Annular H

B

c2

H B

Design type

F

δ = 0.67 × δ=

H

4×F×L B × H3 × E

δ = 1.09 ×

F

δ=

H H/2

Tapered beam width L B B/4

ε × L2 H

6.5 × F × L3 B × H3 × E

δ = 0.86 × δ=

ε × L2 H 3

Taper cross section L

r2

c1

r1

Permissible deﬂection

Equal cross section L

A

c1

ε × L2 H

5.15 × F × L3 B × H3 × E

δ=

B+A ε × L2 × 2B + A H

δ = C2 ×

ε × L2 r2

δ = 1.64

B+A ε × L2 × 2B + A H

δ = 1.64 × C2

δ = 1.28

B+A ε × L2 × 2B + A H

δ = 1.28 × C2 ×

ε × L2 r2

ε × L2 r2

Deﬂection force B × H2 6 E ×ε F =Z× S L

Z =

δ ε L H B A c I Z Es F W Z2 C1 C2 μ φ ϕ

Z =

H 2 B 2 + 4 B A + A2 × 12 2B + A

E ×ε F =Z× S L

= Bending deﬂection permissible = Permissible strain percentage at the root of the beam = Length of the cantilever latch beam = Wall thickness at the root = Upper width at the root = Lower width at the root = Distance between outer ﬁber and neutral axis (center of gravity) = Axial moment of inertia = Section modulus, Z = I / c = Secant modulus of the material = Permissible deﬂection force = Assembly force = Annular modulus cross section = Annular concave cross section = Annular convex cross section = Coefﬁcient of friction = Angle of inclination = Annular angle of concave width

F = Z2 ×

ES × ε L

452

8 Thermoplastic Assembly Methods

8.9.6

Cantilever Snap Latch Beam Permissible Deflection (δ)

The deﬂection (δ) occurring during the assembly operation is equal to the depth of the latch. The permissible deﬂection (δ) depends not only on the size and geometry of the cantilever beam snap latch but also on the permissible strain ratio of the injection molding thermoplastic resin used for the application. A good cantilever beam snap latch design, molded without stresses and using an unreinforced thermoplastic with good spring rate characteristics, performs without problems during the brief snap ﬁtting operation. A partially crystalline melt (molding problems) may be stressed almost to the yield point (probably failure), during the snap ﬁtting operation. When designing with unreinforced semi-crystalline thermoplastic materials, it is recommended to use up to 70% of the yield strain, while for amorphous resins only 50% of the yield strain will do, because stress/strain curves of amorphous resins do not have well deﬁned yield points. Figure 8-72 shows the permissible strain curves for these types of polymers. The test values for permissible strains for injection molding thermoplastic polymers provided by the resin supplier are based on the type of polymer (semi-crystalline or amorphous). The permissible short-term strain values for several materials are:

εYield point

70% εYield point

Stress, (σ)

Unreinforced semi-crystalline polymers with yield point

• Acetal homopolymer

5%

• Nylon 6/6 unreinforced

6 to 10%

• Polycarbonate unreinforced

4%

• PC/ABS alloy

2.5%

• PET glass reinforced

1.7%

• PPO

4%

• PPO glass reinforced

1%

For cantilever snap latch beam designs where frequent separations and rejoinings are required, only 60% of these values should be used. Strain, (ε )

Using the cantilever beam equations shown in Table 8-12, the permissible deﬂection δ can be determined for cross sections of complex shapes. The procedure is illustrated in the cantilever beam snap ﬁtting sample calculations.

Unreinforced amorphous polymers without yield point

Strain, (ε )

Figure 8-72 Permissible strain for unreinforced polymers

εBreak

50% εBreak

Stress, (σ)

The most efﬁcient geometry for a cantilever beam snap ﬁt latch is the taper design, with the thickness of the cantilever beam decreasing linearly to half its initial root value. This version increases the permissible deﬂection by more than 60% compared to a snap ﬁtting cantilever beam of constant cross section. The deﬂection force (F) required to bend the cantilever beam can be calculated by using the equations in the bottom row of Table 8-12, cross sections of various shapes. ES is the strain-dependent modulus of elasticity, or secant modulus.Values for the secant modulus for various thermoplastics can be obtained from resin suppliers. The strain value used should be the one on which the dimensioning of the latch was based.

453

8.9 Snap Fitting Methods

8.9.7

Cantilever Latch Beam Assembly Force (W)

F

During the assembly operation, the deﬂection force (F) of the cantilever latch beam and friction forces have to be derived from the geometrical relationship shown in Figure 8-73.

W

The assembly force is given by:  μ + tan θ  W = F × tan (θ + ρ) = F ×   1 − μ tan θ   μ + tan θ  The values for  can be taken from Figure 8-74.  1 − μ tan θ 

(8-11)

F Friction cone

+ W

The coefﬁcient of friction for various materials are given in Table 8-13. Friction coefficient, µ = tan ρ

Table 8-13 Coefficient of Friction

Plastic

Against metal

Against itself

ABS

0.50–0.65

(×1.2)

Acetal

0.15–0.25

(×1.5) (×1.2)

0.12–0.21

(×1.5)

PE, ﬂexible

0.55–0.60

(×1.2)

PE, rigid

0.20–0.25

(×2.0)

PET

0.30–0.40

–

PBT

0.35–0.40

–

PC

0.45–0.55

(×1.2)

PP

0.25–0.30

(×1.5)

PTFE

0.12–0.22

–

PS

0.40–0.50

(×1.2)

PVC

0.55–0.60

(×1.2)

SAN

0.45–0.55

–

µ = 0.2

µ = 0.4

µ = 0.6

µ = 0.8

0.50–0.60

µ=1

Acrylic PA

10

µ=0

µ + tan 1 - ( µ) tan

8

6

4

2

0 0

10

20

30

40

50

60

Angle of inclination,  μ + tan θ  Figure 8-74 Diagram for determining   1 − μ tan θ 

70

80

Figure 8-73 Relationship between deflection and assembly force

454

8 Thermoplastic Assembly Methods Notes: The values depend on the speed of the mating parts, the pressure applied, and the surface. Friction between two different thermoplastic materials gives values equal to or slightly below those shown above. With two components of the same thermoplastic material, the friction coefﬁcient is generally higher. Where the factor is known, it has been indicated in parentheses. In the case of cantilever latch beams of separable joints, the separation force can be determined in the same way as the assembly force, using Eq. 8-11. The angle of inclination that should be used here is the angle φ.

8.9.8

Design and Material Considerations

A cantilever snap ﬁt latch beam that undergoes dynamic ﬂexing in a snap ﬁt assembly must be designed to approach a “stress-free”, or predetermined,“ allowable” stress condition, when locked in position. This ensures minimum creep distortion of the ﬂexing member for a positive snap ﬁt engagement during repeat assembly operations. Care should be taken to design the ﬂexing member to the elastic limit of the thermoplastic material used (normally below a yield point). If improperly designed, the strain developed can cause permanent deformation and stress cracking. Because thermoplastic materials vary in performance, this stress strain characteristic coupled with the thermoplastic’s physical properties makes the selection of material a critical decision. When choosing the proper resin, a variety of factors must be considered, including molding tolerances, exposure of the assembled unit to high temperatures, and other environmental conditions.

8.9.9

Uniform Cross Section Cantilever Beam

When a uniform cross section snap beam is used, the snap ﬁt design requires a minimum amount of space. It is used in internal and less demanding external snap ﬁt assemblies. The overall design advantage is that the uniform cross section beam makes it easier to analyze the structure and stress concentrations. The tensile strength of a cantilever beam can be increased by extending the beam’s effective length rather than by decreasing the beam wall thickness. Example 8-5

F

0.15 0.06

Determine the maximum strain and load developed for an allowable snap deﬂection of 0.06 in on a uniform cross section cantilever beam, with an effective length of 0.40 in, a thickness of 0.07 in, and a width of 0.15 in as shown in Figure 8-75. The design eliminates the stress concentration (factor of 1), which assumes adequate radii at the root, and the structure is stable at the root and adjacent side wall of the ﬂexing cantilever snap latch beam. From Table 8-12 for a rectangle cross section uniform cantilever beam: δ = 0.67 ×

0.40 0.07

Figure 8-75 Cantilever snap latch beam

F =

ε L2 H

or ε =

δ×H 0.06 × 0.07 = = 0.0391 = 3.91% 0.67 × L2 0.67 × 0.42

B H 2 ES ε 0.15 × 0.072 × 175,000 × 0.0391 = = 2.09 lb. 6L 6 × 0.4

455

8.9 Snap Fitting Methods

In this particular situation, the amount of strain is within the design limits of unreinforced nylon 6/6 at 73 °F and at 50% relative humidity, which possesses a consistent stress-strain behavior from 6 to 10% strain at the yield point. In order to build a safety factor of two, it is recommended that a 4% strain be used as a limit. It will require approximately 2.09 lb of force perpendicular to the cantilever beam applied at the snap latch to create 0.06 in deﬂection for the secant modulus of 175,000 psi for unreinforced nylon 6/6 at 73 °F and at 50% relative humidity.

8.9.10

Tapered Cross Section Cantilever Beam

In the nonuniform cross section cantilever beam, the ﬂexing member can be a combination of varying wall thicknesses and widths. This design approach offers two distinct advantages. First, there is an increase in allowable snap ﬁt deﬂection without increasing the induced strain and second, there is an increase in stability by increasing the spring constant and side load bearing capabilities without increasing the induced strain.

Example 8-6 Consider a cantilever snap ﬁt latch beam of rectangular cross section and with a constant decrease in wall thickness from (H) at the root to (H / 2) at the end of the latch, as shown in Figure 8-76, made of unreinforced polycarbonate. Determine the wall thickness (H) at which full deﬂection δ will cause a strain of 50% of the permissible strain, the deﬂection force F, and the assembly force W. Determination of wall thickness (H):

0.37

F

0.094

30˚ H 2

0.75 H

Figure 8-76 Cantilever snap latch beam

Permissible strain for polycarbonate is: εPC = 4%.

Strain required is: εW = 0.50 × εPC = 0.50 × 4 = 2%.

EO

δ = 1.09 ×

ε × L2 H

1 Secant modulus σ Es = 1 ε1

Stress, σ

Deﬂection equation from Table 8-12 for tapered beam and rectangular cross section is:

1 Strain, ε

Transposing in terms of wall thickness yields: 1.09 × ε × L 1.09 × 0.02 × 0.75 = = 0.130 in δ 0.094

Determination of deﬂection force (F): Deﬂection force from Table 8-12: Z =

E ×ε E ×ε B×H B×H , F =Z× S , or F = × S 6 L 6 L 2

600

2

2

Secant modulus, K-psi

H =

2

ES

400 264 200 0 0

1 2 Strain, ε (%)

3

Figure 8-77 Unreinforced polycarbonate stress, modulus vs. strain

4

456 µ = 0.6

µ = 0.4

µ = 0.2

µ=1

10

µ = 0.8

8 Thermoplastic Assembly Methods

50

60

70

µ=0

µ + tan 1 - ( µ) tan

8

6

4

1.75 2 0 0

10

20

30

40

Angle of inclination, Figure 8-78 Diagram for determining the constant

80

μ + tan θ 1 − μ tan θ Annular angle ϕ˚

10

15˚

8

30˚

6

C1- concave side

45˚ 4 60˚ 75˚ 2 90˚ 105˚ 120˚ 135˚ 150˚ 165˚ 180˚

1.0 0.8 0.6 0.5

0.6

0.7 Radius ratio, (r1/r2)

0.8

0.9

1.0

Annular angle ϕ˚ 10

30˚

15˚

8

45˚

6 60˚ C2- concave side

4

75˚ 90˚ 105˚

2

120˚ 135˚ 150˚ 165˚ 180˚

1.0 0.8 0.6 0.5

0.6

0.7

0.8

0.9

1.0

Radius ratio, (r1/r2)

Figure 8-79 Annular Curves to Determine C1 and C2, constants from Table 8-12

457

8.9 Snap Fitting Methods 0.039 0.000061 0.078 0.2 0.157

0.787

0.0061

1.574

0.0122

2.362 3.149 3.937

0.0242 0.0366

180˚ 165˚ 150˚ 135˚ 120˚ 105˚ 90˚

0.02

Z/r23

0.00122 0.00244 0.00366

0.1 0.06 0.04

0.236 0.314 0.400

0.061

0.01 0.006 0.004 0.002

75˚

0.001 0.0006 0.0004

60˚

0.0002

30˚

45˚

0.0001 0.50

0.60

0.70

0.80

0.875

Radius ratio, (r1/r2)

Figure 8-80 Nomogram to determine modulus Z2 for annular cross section from Table 8-12

From Figure 8-77 at ε = 2.0%, we ﬁnd, ES = 264,000 psi F =

0.31 × 0.132 × 264,000 × 0.02 = 1.33 lb 6 × 0.75

Determination of assembly force: W =F×

μ + tan θ 1 − ( μ) tan θ

From Table 8-12, PC/PC is: μ = 0.50 × 1.2 = 0.6. From Figure 8-78, given μ = 0.6 and θ = 30°, we ﬁnd the constant: μ + tan θ = 1.75 1 − ( μ) tan θ Assembly force: W = 7.33 × 1.75 = 12.82 lb

Example 8-7 ϕ° = 75°, r1 = 0.350 in, r2 = 0.400 in Results: Z2 = 0.000242 in3

15˚ 0.95

Annular angle ϕ˚

0.000242 0.000488 0.000610

r2 (in)

Convex side under tensile stress - Z2 (in3)

0.000122

458

8 Thermoplastic Assembly Methods

8.10

Electromagnetic Welding Method

Electromagnetic welding of injection molded thermoplastic components provides a simple, rapid, and reliable assembly method to produce structural, hermetic, or high pressure seals on most thermoplastic materials. It employs the basic principles of induction heating by developing fusion temperature at the welding interface of the joint components to be welded by using a thermoplastic electromagnetic inter-layer compound.

Speaker housing

The process is versatile enough to weld certain dissimilar thermoplastics plus paper and aluminum to thermoplastics, whether they are reinforced or unﬁlled. Most of the engineering high performance injection molded thermoplastic resins can be used, including such difﬁcult to weld materials as polyoleﬁns. Understanding the design criteria for electromagnetic welding is essential to the new product designer. It involves an understanding of the thermoplastic materials, induction coil designs, and electromagnetic joint designs. Electromagnetic Welding Applications Figure 8-81 shows the diverse welding capabilities of the electromagnetic assembly method.

Gas filter

Advantages of Electromagnetic Welding • All thermoplastic materials can be welded, whether they are crystalline or amorphous • Fusion is developed from within at the bond joint. Variation in the thickness is not a problem • Minimum contact pressure required Attaché case

• Rapid welding cycles • Automatic ﬁxturing for large volume production is possible • Structural, hermetic, and high pressure seals are achieved • Physical and chemical properties of the welded area are similar to the joined materials • Welded joints can be hidden from the surface

Station wagon loadfloor

• Size of the welded area can vary from a small spot to a continuous length up to 10 feet • Storage life of electromagnetic bonding material is indeﬁnite • No pretreatment required for the joint areas • Clean and quiet assembly operation Disadvantages of Electromagnetic Welding • Prototype welding is required to assemble new products, because the electromagnetic weld depends on the welding coil and the part’s joint surfaces ability to be designed for complicated geometries

Portable toilet

Figure 8-81 Electromagnetic welding applications

• The electromagnetic welding process is not recommended for welding electrical devices encapsulated in thermoplastic, such as winding coils, magnets, print circuit boards, sensors, etc.

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8.10 Electromagnetic Welding Method • Electromagnetic welding equipment is expensive • Additional cost for Emaweld bonding system • Welding several different components is difﬁcult to balance using a single electromagnetic welding induction generator • Electromagnetic welding coils have the tendency for overloading, overheating, and arcing

8.10.1

Electromagnetic Welding Process

Welding one thermoplastic component to another is achieved by inductively heating a thermoplastic electromagnetic compound (layer) at the joint interface to the fusion temperature of the abutting thermoplastic components. A chemical bond is obtained within this composite. The electromagnetic layer consists of a dispersion of ﬁne micron-sized metal powders, such as iron, stainless steel graphite, or ferrite. A thermoplastic matrix develops heat losses from the ﬁller concentration in the form of eddy currents and hysteresis losses. When the entire composite is subjected to a high frequency alternating current source (see Figure 8-82), the heat developed in the ﬁller melts the thermoplastic matrix that, in turn, melts the thermoplastic to be welded. Four basic components (see Figure 8-83) comprise the process of an induction generator that converts 60 Hz electrical supply from 3 to 40 MHz output frequency. Most common ranges are from 4 to 10 MHz with output power from 1 kW to 5 kW. The coils are water-cooled copper inductors that develop the requested magnetic ﬁeld intensity through the weld gasket. Fixturing is used to hold the parts and the weld gasket in place in the magnetic ﬁeld. Before welding A weld gasket (e.g., Emaweld®) is deposited in the joint, the mating components are brought together and placed within a holding ﬁxture containing a coil and a nest. Emaweld material Coil

During welding The coil is activated and the weld gasket heats and starts to melt the surrounding surfaces of the thermoplastic joint.

Figure 8-82 Electromagnetic welding principles

Timer

Induction generator

Pressure

Press holder Active coil

Emaweld material

After welding The weld gasket has ﬁlled the joint and has fused with the mating parts, initiating polymer-to-polymer linkage (electromagnetic weld).

Cylinder

Heat exchanger Coil

Coil Final weld

Coil

Emaweld® material

Figure 8-83 Electromagnetic welding process

460

8 Thermoplastic Assembly Methods General Design Considerations When considering the electromagnetic welding process for an application, the following issues should be considered. • Thermoplastic material to be welded: Select the appropriate matrix polymer for the weld gasket • Environmental end use: Determine the best ﬁller type • Physical properties required (structural, hermetic seal, etc.) Determine the ﬁller particle size and joint design • Geometry of the welded area: Determine the coil design and ﬁxturing requirements to hold the parts • Coupling distance of coil/joint. Determine the welding gasket’s ﬁller particle size and frequency • Production volume: Determine the welding requirements for the application, such as batch, continuous, or automatic ﬁxturing • Welding cycle speed: The cycle is determined by ﬁller type, ﬁller particle size, coil type, and cross sectional dimension of the welding gasket‘s electromagnetic material, power output, frequency output, part size, and joint conﬁguration. The two most important parameters in determining if a thermoplastic part can be welded using the electromagnetic welding process are the coil and joint design. All other variables, such as power frequency, ﬁller type and size, and polymer types are easily controlled and are tailored for each speciﬁc application.

8.10.2

Electromagnetic Welding Coil Design

Primary consideration for a successful electromagnetic welding application is the proper design of the coil. If fundamental principles of good coil design are maintained, costly failures are eliminated and efﬁcient assembly welding operation is achieved. Coils used in electromagnetic welding are similar to those used in standard induction heating in annealing or heat treating of metals. In electromagnetic welding applications, the preferred frequency ranges from 4–10 MHz. Because of the higher frequencies, coil design becomes important for reducing any tendencies for overloading and arcing. The coils can be made in a large variety of types and styles, depending on the shape of the joint to be welded and the location of the weld area. Their design must follow certain principles for maximum efﬁciency of the high frequency generator to be obtained. The induction coil quickly raises the temperature of the welding gasket located at the bond interface by passing high frequency current through the coil and inductively into the electromagnetic bonding material. This, in turn, develops heat from hysteresis and eddy current loss. The coil becomes the primary of a transformer and the work becomes the secondary. The electromagnetic bonding material to be heated is in no way a part of the closed electrical circuit; therefore the generation of heat is solely by induction.

461

8.10 Electromagnetic Welding Method Heating the weld gasket is the result of energy losses within the material that cause the temperature to rise. In ferrous metal powders such as iron or stainless steel, which have magnetic properties, these losses are caused by hysteresis and eddy currents up to the Curie point. At this point, magnetic properties cease to exist and from here on, heating is attained only through eddy current losses. In nonmagnetic ﬁller materials, the only losses present are from eddy currents; therefore, the induction of heat requires more time. Since magnetic ﬁelds occur in the area surrounding the coil and are stronger next to it than at any distance away from it, there is an advantage in locating the bond line close to the coil so that a maximum of heat energy may be transferred to it. The strength of the ﬁeld varies inversely with the square of the distance between the work and the coil. This means that the coupling distance will have a direct relation to the amount of heat generated in the welding gasket at a given time. Materials Used for the Coils Coils are made from copper tubing, copper sheet stock or machined copper blocks with brazed water cooling. Coils made from tubing can be round, square, or rectangular and range in size from 0.125, 0.187, 0.250 to 0.375 in. The 0.125 in size should be used very sparingly, because its small area may cause overheating due to insufﬁcient ﬂow of water for cooling. For short heat cycles and small parts, where slight heating of the coil may have no effect, this size can be considered. Round copper tubing can be used for many types of coils; however, it is usually preferable to use square tubing because coupling distance to the work piece is very important. Copper sheeting stock of 0.062 in thickness has also been used for applications running up to 5.00 in wide by 20.00 in long. This allows for large surface area sealing for irregular ﬂat shapes. Solid copper blocks are made from rectangular stock generally 0.375 in by 0.750 in and are mainly for electrical contacts on split coils. Types of Electromagnetic Welding Coils The following types of coils are used for electromagnetic welding: • Single turn coils

• Multi-turn helical coils • Pancake coils • Hairpin coils • Split coils

• Combination coils

• Multiple coils in series or parallel Single Turn Electromagnetic Welding Coils A single turn coil is shown in Figure 8-84. Due to the direction of the current and resultant magnetic ﬁeld there is a tendency to develop a relatively weak ﬁeld intensity at the end of the coil leads. This is sometimes overcome by using a crossover lap and keeping the leads at that point within 0.187 in. If the injection molded thermoplastic component design permits the use of a two turn coil, this will eliminate this deﬁciency.

Figure 8-84 Single turn electromagnetic welding coil

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8 Thermoplastic Assembly Methods Multi-Turn Electromagnetic Welding Coils Figure 8-85 shows the most common cylindrical or helical multi-turn coils of copper tubing, which is suited for welding round containers. Rectangular or square coil and the irregular shaped coils are used for welding irregular shaped items. Figure 8-85 Multi-turn electromagnetic welding coil

Figure 8-86 Pancake electromagnetic welding coil

Pancake Electromagnetic Welding Coils The pancake coil shown in Figure 8-86 is used for heating ﬂat surfaces, while the spiral, helical coil is used for heating conical surfaces. For maximum efﬁciency, a reﬂector coil is placed in the internal coil diameter to concentrate the ﬁeld within the joint area. This is used for part sizes from 6–25 in. Modiﬁcations on these coils are possible by connecting them as single coils in series or parallel. As a general rule, the length of a helical coil should not be more than three to four times its diameter. Pancake coils are made from tubing by winding a ﬂat horizontal plane with multiple turns until the desired working diameter is reached. Typical applications are for sheeting, ﬂat or contoured surfaces, or large surface areas. Hairpin Electromagnetic Welding Coils Hairpin coils are actually single turn coils squeezed together resulting in a coupling distance between the turns equivalent to the part thickness. The magnetic ﬁeld is highly concentrated as the coupling distance is reduced. Such coils can be wound into irregular shapes and are very often used as dies, as in a press action with movable and ﬁxed platens. They are very efﬁcient for bonding long ﬂat sheets or perimeter seals of two sheets for tray or mat composites. By far, the hairpin type coil and helical coil are the most commonly used in many applications. Because of the magnetic ﬂux path around a helical coil, the greatest ﬁeld strength is within the coil itself, rather than on the outside turns. For this reason, it is best to locate the weld joint within the internal diameter of the helical coil. Split Electromagnetic Welding Coils Split coils are used quite often for easy removal of the part. For example, instead of making a single turn coil and moving the part in and out, the coil can be split in half and electrical contacts made by actuating with an air cylinder. Naturally, all coils must be water-cooled by providing a separate water circuit. Impolene hose is generally used to make the interconnections. Split coils are very useful for large parts, such as pipe or conduit or for bonding two halves of a tray. Combination Electromagnetic Welding Coils Combination coils are used for welding multiple units with the same induction generator. They can be a series or parallel connection of two or more coils. Each coil can be either of the same or of a different type. There is a limit to the number of coils that can be connected. By using too many coils the heating efﬁciency can be drastically reduced. In some cases, a 2-station or 3-station generator is used to connect multiple coils. Each coil has its own independent cycle and they cannot operate both at once. Electrical circuits are designed to sequentially activate each cycle.

8.10 Electromagnetic Welding Method Multi-Electromagnetic Welding Coils in Series or Parallel Solid type coils are made of rolled copper plate and can be arranged for single or multiple operation. The coil can be made to weld multiple round units either by moving the coil with the part or by introducing the part in the coil. They are made of a thick copper plate, from 0.250 in to 0.375 in as required, bored to suit the diameter of the part, with sufﬁcient coupling distance provided for easy entry and removal. Two connecting blocks are brazed to this plate, then the plate is cut out to allow the high frequency current to follow the path, coming in at one block and going out at the other.

8.10.3

Electromagnetic Welding Joint Design

There are many approaches to acieve the proper joint design, depending on the application, whether the part is injection molded, blow molded, extruded proﬁle, or thermoformed. Ideally, the design should contain the electromagnetic bonding layer by having it exert an internal pressure on the abutting surfaces to be welded. The analogy is somewhat similar to ﬁlling a cavity in an injection mold. It should not be under-ﬁlled. The electromagnetic bonding layer is inductively heated and must transfer, by conduction, sufﬁcient fusion temperature to the bonding surface. Because the bonding layer will melt when heated, it will ﬂow in the path of least resistance and ﬁll any void areas. The joint designs and uses are shown in Figure 8-87. Flat-to-ﬂat joint This joint is suited for parts with long weld lines, or for continuous welding such as solar panels. Flat-to-groove joint This joint is used when the location of the weld is critical, such as in automotive panel applications. Tongue-and-groove joint This joint provides the ultimate weld strength; it is well suited for hermetic weld seal applications. Step joint Excellent pressure-proof weld; this joint is used for welding plastic parts with wide tolerance variations.

Shear joint Superior weld strength, suitable for high pressure seals, used for pressure vessel applications.

Figure 8-87 Electromagnetic welding joint designs

463

464

8 Thermoplastic Assembly Methods

10°

0.04

0.158

0.036 0.06 0.14

0.045 10°

0.06 0.02

0.18

0.095 Dia.

0.245

R.

0.125

0.086

0.04 0.194

0.02

0.03

Figure 8-88 Tongue and groove joint

0.185 0.035

A proper joint design is essential to the ultimate success of the weld. Because the welding gasket located at the joint interface becomes molten when activated, it ﬂows under pressure into grooves and irregular surfaces to produce reliable welds. Ideally, the molten ﬂow should be contained and subjected to an internal pressure against the abutting weld surface. Equation 8-12 is used to determine the amount of bonding material required to ﬁll the joint: AE = AG k

0.043

(8-12)

Where: AE = Cross sectional area of the bonding material AG = Cross sectional area of the groove in the joint k = Constant ranging from 1.02 to 1.05 depending on the amount of joint interface pressure desired

0.157

Figures 8-88, 8-89, 8-90, and 8-91 show examples of actual joint detail conﬁgurations. A quick calculation will show that a factor of k = 1.042 was applied in Figure 8-89. Figure 8-91 shows how two different components are joined in a double groove of the third component; all three injection molded thermoplastic components are welded together with one welding gasket during a single step of electromagnetic energy.

AE = AG k

Figure 8-89 Flat-to-flat joint

0.171 0.040

8.10.4

Available Welding Gasket Shapes and Forms

Most welding gaskets are supplied by Emabond® Systems as a key component. The Emaweld® material formula is generally based on the same injection molding thermoplastic resins to be joined. In some cases, a similar melt index is used to ensure efﬁcient melt mixing in the welding joints.

0.161

0.097

0.062

Figure 8-90 Step joint

Emaweld® materials are compounded with ﬁne particles of iron, stainless steel, or other magnetic materials. The ﬁller type, the amount of ﬁller, and the ﬁller particle size determine the required or optimum cycle time. These are related to the matrix material and the joint conﬁguration as the latter often predicts the coil size and the distance of the coil to the weld area. Emabond® will select and provide the best performing and most economical weld material for each application. The Emaweld® welding gaskets are shown in Figure 8-92.

0.017

0.190

0.015

0.015 0.171

0.017

Figure 8-91 Shear joint for welding three components

Figure 8-92 Emaweld® welding gaskets in various standard shapes

8.11 Vibration Welding Method

8.11

Vibration Welding Method

Vibration welding is based on the principle of friction welding for joining thermoplastic materials. Spin welding, the most common method of friction welding has been used successfully for over 75 years. In vibration welding, the heat necessary to melt the thermoplastic is generated by pressing one of the parts against the other and vibrating it through a small relative displacement in the plane of the joint. Heat generated by friction melts the thermoplastic at the interface. Vibratory motion is then stopped and the parts are automatically aligned. Pressure is maintained until the thermoplastic solidiﬁes to bond the parts permanently with a strength approaching that of the parent material. Vibration welding should not be considered in competition with existing and well known assembly welding methods. It provides a means to join parts for which none of the other methods are suitable and gives the designer considerably more liberty in the dimensioning and appearance of the ﬁnal products. The melt time generally ranges from 2 to 3 seconds and the hold time is usually 1 second or less. Total cycle time averages between 6 to 15 seconds, including loading and unloading. Multiple cavity ﬁxtures or an automated welder system can also be used with the vibration welding process. The vibration welding method can join most of the thermoplastic materials, with small as well as with very large thermoplastic components. The change-over from one product to another takes very little time with the vibration welding process, even though the products can vary greatly in size, geometry, or number of components. The upper and lower nest plate sets are replaced with units designed for the new product, the clamp load adjustment is changed, and the weld and hold times are reset. Once these values have been determined for a product, they can be quickly set again when the plates are changed.

8.11.1

High Frequency Vibration Welding

Vibration welding was developed as a 120 Hz process with one part moving in relation to the second part at amplitudes to 0.16 in. Both mechanically driven systems and spring mass systems, driven at resonance, have been developed as production equipment at this frequency. In spring mass systems, a ﬁxed frequency operation requires that the mass of the vibrating parts be kept constant to maintain resonance at the operating frequency. A more recent development is variable high frequency vibration welding, which reduces the required amplitude of motion and eliminates the need to tune to a ﬁxed frequency by controlling the mass of the tooling. Typical vibration frequencies for this process range between 250 and 300 Hz, with vibration amplitudes for effective welding ranging between 0.03 and 0.06 in. Several important beneﬁts are realized at higher vibration frequencies: ﬁrst, for the same velocity of relative motion between parts, the higher frequencies allow smaller displacement amplitudes. Smaller displacement means heat generated by the friction is conﬁned to a narrower region and quicker melting results. In a number of instances, it has been found that this restriction on the heated area has proven to create welds of superior quality at lower velocities than required at the original 120 Hz.

465

466

8 Thermoplastic Assembly Methods Secondly, when wall thicknesses of the parts to be joined are comparable to the displacement amplitudes, reduced displacement amplitudes yield better coverage. In particular, when thin, mating walls have the same thickness, smaller vibration amplitudes can be especially beneﬁcial. With a 0.09 in wall and a 0.05 in displacement typically used at 270 Hz, nearly 50% of the weld zone is always covered. At the larger amplitudes used at lower frequencies, the area at the ends of the part is exposed twice during each cycle for more than double the time. Here again, the smaller amplitude can concentrate the frictional heat in a limited area, resulting in a sound, leak-tight weld that is produced in a shorter time. Weld joint areas under walls that are parallel to the direction of relative motion are not uncovered by even the large amplitude motion. The small amplitude of the high frequency welder is beneﬁcial here also, because the melted material stays in place near the joint and forms a more solid ﬂash. Large amplitudes tend to move the melted material further and form a ﬁne thread-like ﬂash in some products. This may result in objectionable appearance and special ﬂash traps must be designed into the weld joint. The appearance of ﬂash resulting from small amplitude vibration is frequently acceptable, even with simple butt joints.Very simple design changes can be made to the butt joint for situations where no visible ﬂash is acceptable. Walls can be as thin as 0.09 in, because of the small clearance required for high frequency, small amplitude motion. Additional weld area can be obtained if a small ﬂange is added. The latter has the additional advantage of improved loading of the weld joint.

8.11.2

Vibration Welding Modes

Vibrational displacement of the parts is produced by a reciprocating motion and is referred to as the linear mode. Angular displacement is produced by a back and forth rotational motion around an axis in the center of the part. The angular mode is not as versatile as the linear mode and in most cases can be used to weld only one part at a time. With the linear mode, several parts can be welded simultaneously. Angular Vibration Welding Modes Angular welding means that the oscillations are generated around a center point, which can either be located inside or outside the joint area. For circular shaped parts, especially with large diameters, they are oscillated around their own center. Thus they can be provided with “V” shaped joint proﬁles (like spin welded parts), which provide the highest strength. In this case, only one part can be welded at a time. Since this is not always the most economical way, circular parts of small diameters are often provided with butt joints in order to allow for the welding of several parts simultaneously. Since all noncircular parts require a butt joint, it does not actually matter whether they are welded by angular or linear motions. Figures 8-93 and 8-94 show the angular vibration welding for circular and noncircular shaped applications.

467

8.11 Vibration Welding Method

Noncircular

Noncircular

Circular

Circular

Noncircular

Figure 8-93 Angular vibration welding applications

Motorcycle gas tank

Container

Automotive radiator reservoir

Automotive water pump

Figure 8-94 Angular vibration welding applications

468

8 Thermoplastic Assembly Methods

Figure 8-95 Linear vibration welding applications

Linear Vibration Welding Methods If the vibrations take place in a straight line, the procedure is called linear welding. Theoretically, all parts can be welded this way, except when previously assembled internal parts do not allow a linear displacement. A machine built for linear welding cannot be used for angular vibrations; whereas on angular welding machines, the parts can be located on the center or outside the center.

Head lamp

Container

It is important to know that about 95% of all vibration welded parts studied are suitable for linear welding. This method has proven to be more suitable for the speciﬁc ﬁelds in which vibration welding is applied. Figures 8-95 and 8-96 show the linear vibration welding for various shape geometries and applications. Part Geometry, Size, Shape, and Material

Motorcycle gas tank

One of the major advantages of linear vibration welding is the ability to weld larger parts. Components as long as 22.00 in and as wide as 12.00 in are being successfully welded on standard equipment, even extremely large parts, such as injection molded thermoplastic automobile body components, reservoirs, emission control canisters, and gas tanks. For automatic welding processes, two or more power units can be grouped together with a single ﬁxture. Complex, irregularly shaped parts can be joined by linear vibration welding; it is not restricted to circular parts. The outer conﬁguration of a part is not limiting, as long as the parts can be properly ﬁxtured to hold them securely and to provide the necessary force at the joint. Welding is not limited to the periphery of the part. Inside surfaces and dividers within a hollow part can be welded to produce several separate compartments, each completely sealed from the other, as in automobile battery cases.

Automotive brake fluid reservoir

Automotive emission control canister

Figure 8-96 Linear Vibration Welding Applications

Ideally, the welding joint is in a single plane. However, parts with more than one plane can also be welded.Variation in the joint must be limited to two dimensions and all surfaces in the third dimension must be parallel. Excessive joint plane variation should be avoided. All injection molding thermoplastics (with only minor exceptions) can be vibration welded, whether they are processed by injection molding, extrusion, blow molding, or thermoforming. The vibration welding method is particularly advantageous for semi-crystalline resins, such as acetal homopolymer, nylon 6/6, thermoplastic polyester, polyethylene, and polypropylene. These resins are not as easily joined by ultrasonic welding or solvents as are amorphous thermoplastics, such as polystyrene, ABS, acrylic, and polycarbonate.

8.11 Vibration Welding Method For hygroscopic resins, such as nylon, the molded parts do not have to be dry for vibration welding. Filled, reinforced, and foamed resins are easily welded. The reciprocating action also pushes contaminants out of the way. Even a coated surface can be welded without extra coating removal operations. Masking is unnecessary, because the vibration welding can be done through metalized material, through painted material, even through carbon and dust. Vibration Weld Quality Vibration welding produces high strength and, where needed, pressure-tight seals. In many cases, the strength of the weld approaches that of the parent materials. Pressure-tight seals of such high quality have been produced that, when subjected to burst tests, the parts failed outside of the joint, leaving the weld intact. In transparent materials, the weld is optically clear with homogeneous bonding of the material. Required Relative Motion The most important design requirement for vibration welding is that the vibrated part be free to move relative to the stationary one in the plane of the joint interface through an amplitude (stroke) of at least 0.120 in. This is the peak-to-peak amplitude. For best results and fastest cycles, the design should allow for an amplitude of up to 0.160 in.

8.11.3

Comparing Vibration Welding to Other Assembly Methods

Ultrasonic Welding Method There is very little practical overlap with ultrasonic welding. Applications that can be welded by either ultrasonic or vibration welding are more economically and rapidly joined by ultrasonic welding, because of shorter cycle times and lower equipment costs. However, with vibration welding, it is possible to weld larger and more complex parts. Crystalline resins, low modulus resins, ﬂuoropolymers, cellulosic, and other resins that are more difﬁcult to weld ultrasonically can be easily welded by vibration welding. Ultrasonic welding is: • Less expensive • Has geometry limitations • Has size limitations • Has shape limitations • Has difﬁculty with some thermoplastic materials Spin Welding Method Spin welding is faster and less expensive than vibrational welding, but it is limited to joining parts that have circular joint areas. Injection molded thermoplastic components requiring a precise orientation at the end of the weld cycle, such as those with mounting brackets or inlet and outlet connections are difﬁcult to orient accurately with spin welding. Such orientation is easily achieved with

469

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8 Thermoplastic Assembly Methods vibration welding. Unlike spin welding, more than one part can be welded at a time. Spin welding is: • Faster • Less expensive • Is for circular components only • Difﬁcult to register • Capable of welding one or two components per cycle Adhesive and Solvent Welding Method Characteristics of adhesive and solvent welding: • Requires low initial cost • Requires adhesive surface preparation • Requires UV curing light energy • Provides temperature stability • Results in longer welding cycles • Is a labor intensive operation • Difﬁcult to automate the process • The process could be toxic • Handling problems with the adhesives Hot Plate Welding Method Hot plate welding is relatively expensive and welding cycles are considerably longer than with vibration welding. Because of the absence of heated coils and fewer components requiring alignment, an application can be set up faster and easier. Vibration welders operate at a peak power of 1,800 W (only during the weld cycle), while a comparable hot plate welder uses 5,000 W. Control of temperature and fouling of heating plates are problems of hot plate welding that are nonexistent with vibration welding. Characteristics of hot plate welding: • Greater size/shape capabilities • Welding problems with high temperature crystalline thermoplastics • Material degradation • Requires longer cycle times for welding • Higher equipment cost for same production volume • Higher ﬁxture cost • Higher electrical energy consumption • The process requires exhaust ventilation • Higher maintenance requirements

8.11 Vibration Welding Method

8.11.4

Vibration Welding Equipment

Vibration welding equipment consists of a linear motion generating system, a vibrating element, and a clamping mechanism. Vibration may be initiated by activating electromagnets, while a set of ﬂat springs control the relative position of the components being welded. Alternatively, the entire system may be controlled by hydraulics. Injection molded thermoplastic components to be vibration welded are limited in width to about 24.00 in and depths of 12.00 in on conventional vibration welding equipment. However, with side-to-side ganging in the direction of motion, this length may be adjustable to vibration weld long components, such as automotive bumpers. The necessary parameters for obtaining a good vibration joint are: • Amplitude

• Vibration frequency

• Thrust pressure on the joint interface • Weld time

An ideal machine designed for welding parts of various sizes should provide for adjustment of all of these parameters within reasonable limits. Amplitude Amplitude is a function of the joint width and should be adjusted to maintain some overlap in the extreme positions. The peak-to-peak amplitude of the welder ranges from a minimum of 0.120 in to a maximum of 0.160 in. High amplitude results in a shorter weld time. The amplitude of the welder is adjusted by changing the mass of the vibrating ﬁxture rather than the power to the coils. This makes it possible to have full power available throughout the range of amplitudes. Vibration Frequency If the vibrations are generated by a mechanical gear box; the frequency can be adjusted by varying the driving motor speed. There are, however, limits due to the acceleration forces acting on the mechanical parts. If the vibrations are produced by electromagnetic systems, frequency is twice that in the power supply, e.g., 120 or 240 Hz. This does not allow for adjustment without rather expensive additional equipment. However, experience has shown that this frequency is quite convenient for all practical purposes and that an adjustment is not really required. The frequency of vibration and linear displacement are: • Frequency of 120 Hz; linear displacement from 0.080 to 0.160 in • Frequency of 240 Hz; linear displacement from 0.030 to 0.060 in Thrust Pressure on the Joint Interface Experience has shown that a correct weld for an injection molded thermoplastic component requires a joint pressure (force exerted per square inch of joint area) from 100 to 200 psi. The joint pressure depends on the materials, coefﬁcient of friction, and the energy required to melt. Too high or too low a pressure may produce unsatisfactory welds and extensive preliminary tests are therefore required.

471

472

8 Thermoplastic Assembly Methods Weld Time

1.25 T

0.5 T

28˚

Weld time (duration of vibration) is generally 2 to 3 s with 0.5 to 1 s hold time. Typically, the total cycle including loading and unloading varies between 6 and 15 s, depending on the degree of automation and the size and complexity of the injection molded thermoplastic components. Because of the relatively high weld pressure, the joint faces can be brought into contact even if the parts are slightly warped. However, very rigid components, especially in glass reinforced thermoplastic materials, are too rigid to deform under pressure. Any warpage requires additional melting of material to achieve tight and strong joints, thus increasing weld cycle time considerably and producing more ﬂash than is normal for perfectly ﬂat parts. It is, therefore, advisable to design and mold the parts in a way as to eliminate warpage as much as possible.

8.11.5

Vibration Welding Joint Design

The basic joint design for vibration welding is a simple butt joint. A ﬂange is generally required unless the wall is sufﬁciently rigid or supported to prevent ﬂexure. A ﬂange makes it easier to grip the parts and apply uniform pressure close to the weld. It also increases the strength of the joint above that of the surrounding structure. A ﬂange width of 2 to 3.5 times the wall thickness is desirable for maximum strength and support. The minimum weld surface width is 60% of the total relative motion (peak-to-peak amplitude).

32˚

T

For thin or long, unsupported walls, special provisions may be necessary, such as an “H” ﬂange. With this type of ﬂange, walls as thin as 0.032 in have been welded successfully. An alternative is the use of knurling or serrations on the surface of the ﬁxture plates in conjunction with a basic ﬂange.

10% T

10% T

Internal walls or dividers may require special features to prevent deﬂection during welding. This can be accomplished with short ribs on either side of the wall and a number of small tapered tabs to guide the wall into the groove for pre-assembly. As in any method where joint surface melting takes place, some of the melted material may be displaced from the joint. This is commonly referred to as ﬂash. If it is objectionable for functional or aesthetic reasons, suitable traps can be designed in the joint to completely contain ﬂash. Basic ﬂash trap designs along with variations and typical dimensions are shown in Figure 8-98. The traps should be volumetrically sized to the amount of material displaced during welding (multiplied by 1.25 for minor variations). For large size, circular shaped injection molded thermoplastic components, it is advisable to design and mold the parts in a way as to eliminate warpage as much as possible. For components, which can be welded angularly, this type of joint gives the best strength and requires a smaller protruding bead on the outside diameter.

Figure 8-98 Circular “V” joint with flash trap designs

Figure 8-98 shows a few circular joint designs with various ﬂash trap arrangements. If, as it happens on industrial parts, the protruding ﬂash can be tolerated or removed afterwards, the ﬂash traps can be ignored and the total joint width correspondingly reduced.

473

8.11 Vibration Welding Method

2-3T T 0.03 R.

T T T

3T

1.2 T 10% T

3 - 3.5 T Figure 8-99 Noncircular butt joint with flash trap designs

Vibration Welding Butt Joints Since “V” joints are limited to angular welded circular shaped parts, most of the applications are provided with butt joints. Typical noncircular joint designs are shown in Figure 8-99. Because stresses on loaded joints are highest on the inside and decrease toward the outside, strength is not signiﬁcantly improved if the joint width is more than 2 to 3.5 times the wall thickness. Thin wall vessels and all large parts must be provided with a joint as shown in Figure 8-99, lower right, to prevent bending of the wall. Butt joints can also be designed with various ﬂash traps. However, the ﬂange width is increased, which may not be acceptable.

8.11.6

Vibration Welding Aligning and Fixturing

When designing thermoplastic components that are assembled by vibration welding, a few basic requirements must be applied to obtain proper welds. • The part must be provided with a correct joint proﬁle and a sufﬁciently large weld area to achieve a strong and tight weld. This is especially important for all types of butt joints because their strength, compared to the tensile strength of the resin, has certain limits. • The parts must be designed to allow easy ﬁtting and ejection from the welding ﬁxture. • Vibration welding requires a relatively high speciﬁc pressure on the whole joint area, which is applied through the ﬁxture. Inadequately designed parts do not always fulﬁl this requirement. This problem causes poor joint strength and leakage.

474

8 Thermoplastic Assembly Methods • Transmission of the vibration energy to the components produces high acceleration forces that tend to bend the walls rather than generating the desired frictional energy. This condition is very important with thin walled vessels and on components with internal ribs. Injection molded thermoplastic components should ﬁt in the ﬁxture closely to prevent wall ﬂexure and motion (slippage between the part and the ﬁxture), which may mark the parts, result in the loss of amplitude, and produce excessive noise, especially with very rigid thermoplastic resins. Because of dimensional variations, large molded thermoplastic components should be located and driven by a small feature, such as a boss or recess close to the center of the parts rather than at the extremes. For ease of loading the parts in the linear vibration welder and for higher productivity, the two molded thermoplastic components to be joined are placed together in the stationary ﬁxture. As the two ﬁxture plates come together during the operating cycle, the part to be vibrated is fed into the vibrator ﬁxture plate. To facilitate this method of loading, alignment devices can be included in the stationary ﬁxture. However, an easier method would be to design the molded thermoplastic components with a break-away pin or similar feature which would align the parts prior to placing them in the ﬁxture and would break away when vibrations begin. Much of the success of an application depends on the molded thermoplastic component holding ﬁxtures. They are generally aluminum plates with cut-outs to accept the molded thermoplastic components. Contoured cavities are used for parts that cannot be supported by a ﬂange and are generally cast of aluminum or epoxy. Larger parts may require more complex designs.

8.11.7

Vibration Welding Tolerances

Due to the ﬂow of material from the joint during welding, the height of the welded assembly is reduced by 0.016 to 0.032 in, depending on the thermoplastic material, the ﬂatness of the parts, and the vibration welding parameters. For a speciﬁc set of welding conditions, the tolerances should be 0.004 in or less. The horizontal alignment of parts in the direction of welding should be accurate to less than 0.010 in if the parts ﬁt fairly closely in the ﬁxtures with little loss of relative motion. Perpendicular to the direction of welding, alignment should be within 0.005 in.

8.11.8

Vibration Welding Equipment

The vibration welding method employs an exceedingly simple, maintenance free mechanism, consisting of only one moving element with no bearing surfaces to wear or require lubrication. Reciprocating motion is achieved by magnetic force alternating at 120 cycles per second (100 cps for 50 cycle current), acting directly on a resonant mechanical suspension. This electromagnetic system overcomes the limitations of previous attempts at implementing vibration welding where rotary energy is converted to reciprocating motion (100 cps) and the mechanical drive is subject to complexity and abuse inherent in rapidly oscillating mechanical systems. Figure 8-100 shows a linear vibration welder mechanism cross section.

475

8.11 Vibration Welding Method Springs

Vibrating element

Electromagnet

Vibrating part

Electromagnet

Weld joint Stationary element

Stationary part

Figure 8-100 Linear vibration welder mechanism cross section

Vibration welding single station

Vibrator Welder Components The major components of the vibrator are shown in Figure 8-100. They are a set of ﬂat plate springs, two electromagnets, a vibrating element, a stationary element, and a clamping mechanism. The springs have three functions. They are the resonating members, they support the vibrating element against vertical welding pressure, and they return the vibrating element to the aligned position when the magnets at the end of the weld time are de-energized. The vibrating element engages and holds the plastic part to be vibrated and the stationary element holds the other part of the assembly.

Rotary turntable

Pressure is applied to the injection molded thermoplastic components by a pneumatically operated clamping mechanism that engages the stationary element and pulls it toward the vibrating element during the welding cycle. Vibration Welding Process Automation Because only one of the two thermoplastic components to be welded is vibrated, the vibration welding process can be easily automated. Semi-automatic operation is possible with minor modiﬁcations to the standard vibration welder. For an automatic vibration welding operation, the vibration welder can be mounted into either a rotary or in-line system with a remote power supply. One type of automatic vibration welding system, an adaptation of the basic machine, employs a rotary turntable for remote load/unload operations. Another system employs one or more welders that can be mounted over a transport device such as a conveyor and lowered onto the parts for the welding cycle.

Figure 8-101 Single vibration welder with rotary turntable

Three vibrator welding stations Automatic loader system

A single station, manual process vibration welding system has a typical cycle time between 12 and 15 s, including loading and unloading. Automated cycle times drop to approx. 6 s for single part vibration welding operation. Figure 8-101 shows a single vibrator welding station mounted over a rotary turntable for remote load/unload operations. Figure 8-102 shows a fully automatic in-line conveyor line system with three vibrator welding stations. A similar system has been employed for many years to assemble automotive emission control canisters made of nylon 6/6 in a total vibration welding cycle time of 6 s.

In-line conveyor Figure 8-102 Three vibration welder stations with automatic loader and in-line feeding conveyor

476

8 Thermoplastic Assembly Methods

8.12

Spin Welding Method

Spin welding is a rapid and efﬁcient assembly method for producing circular surface weld joints of most injection molding thermoplastic resins. Spin welding should be considered for thermoplastic components molded of similar polymers having circular surface weld joints, when the application requires strong, permanent, leak-free weld joints. Spin welding thermoplastic components molded of dissimilar materials generally results in low weld joint strength. Spin weld joints are made by rotating the surface of the thermoplastic end cap to be welded at high speeds while the other part (base) is supported in line with the ﬁxed chuck. The frictional heat is generated by the differential speed at the joint between the surfaces when they are held together. After a ﬁlm of melted polymer has formed, relative motion is stopped and the weld is allowed to solidify under pressure. The spin welding cycle can be automated and the complete welding operation performed in 1 to 2 s.

8.12.1

Applications

Figure 8-103 shows some typical spin welding applications. Figure 8-104 shows the basic spin welding principles.

Automotive emission control canister

8.12.2

Basic Spin Welding Equipment

The spin welding process can be performed on a high speed drill press with a driving tool and a vise or clamp provided for holding a prototype or low production volume injection molded thermoplastic part. High volume applications require either a drill press equipped with a spin welding tool, air cylinder, valves and timer, or specially designed spin welding equipment.

Seal container

Driving tool spins & descends Plastic top trip by tool

Lighter end cap

Plastic base clamp by chuck

Spin, downward pressure stop

Automotive intake manifold end cap

Figure 8-103 Typical spin welding applications

Flash outside weld

Driving tool moves up

Welding completed

Figure 8-104 Basic Spin Welding Process Sequence

477

8.12 Spin Welding Method Basic equipment requirements for a spin welder are shown in Figure 8-105. Parts to be welded are held in the driving tool and a ﬁxed chuck. A spinning tool is attached to a clutch, an electric motor, a motor speed control, and motor brake may be required. An air cylinder advances the tool and applies force to the parts during welding. A timer and air valve are used to control the air cylinder.

Brake Motor

Clutch

Timer

8.12.3

Spin Welding Variables

There are three basic variables in the spin welding process: rotational speed, joint thrust pressure, and spin time. Rotational speed depends on the circular joint diameter. The average lineal velocity at the joint should be about 20 to 40 ft/s. Speeds greater than 3,000 rpm are not recommended for standard spin welding. For joint surface diameters less than 1 in, where higher speeds would be indicated based on 20 ft/s lineal velocity, greater joint pressure and/or longer cycle times may be used to generate the necessary frictional heat. Joint thrust pressure during the rotation for spin welding should be approximately 200 psi, applied directly to the joint welding surfaces. Polyethylene and polypropylene resins can be spin welded using lower joint thrust pressures, approx. 50 to 100 psi. Adjustments in joint pressure are made to suit speciﬁc applications and joint conﬁgurations. Spin time is the time that the joint pressure is supplied by an air cylinder during the forward stroke of the rotating spindle. The spin welding cycle starts when a switch activates the air valve to the cylinder and the rotating driving spindle chuck advances toward the stationary mating part that has been positioned and retained in alignment. An electric timer begins as the switch is activated. The rotating driving tool engages and spins with one part against the other under pressure, creating a layer of melt by frictional heat. The excess melt is pushed out forming a ﬂash at the joint. When the electric timer runs out, the driving tool is retracted to complete the spin welding cycle. Once melting has occurred, it is essential that the relative motion between the parts be stopped to allow the melt to solidify under pressure. Failure to stop this motion will cause tearing of the weld and result in low weld strength. The rotational motion between the parts may be stopped by an electric motor brake, through the driving tool.

8.12.4

Types of Spin Welding Processes

Two processes have been developed for the spin welding method: spin pivot welding and spin inertia welding. Either type can be designed for a given joint; however, the spin pivot welding process is usually better suited for joints having a projected area of 0.50 in2 or less, while the spin inertia welding process is better suited for joints with larger joint surface areas. 8.12.4.1

Spin Pivot Welding Process

The spin pivot welding process requires the use of a special driving tool mounted on a rotating and vertical press, as shown in Figure 8-106. However, this welding mechanism is not commonly used in the fabrication of commercial spin welders.

Air cylinder Air valve

Grip teeth

Driving tool

Plastic base clamp

Spinning plastic cap Centering chuck

Figure 8-105 Basic spin welding components

478

8 Thermoplastic Assembly Methods

Pressure ring adjustment Thrust spring Driving tool Cap

Pivot pin depresses the cap Grip tooth

Body

Centering chuck

Driving tool spins and the pivot pin depresses the cap

Elements of the spin pivot welding cycle are: • The rotational pivot pin spindle should be moving between 3,600 to 5,000 rpm and with a spring thrust load of 20 lb directly on the center pivot. The stroke timer is set at about 0.5 s. • A toothed crown is provided for driving the rotating thermoplastic part, whereas the threaded ring allows for adjustment of the spring load on the center pivot. As the driving tool advances, the pivot pin engages at the center location for alignment. • The pivot pin is compressed as the driving tool continues to advance until the teeth or rubber liner holder grips the outside surface of the rotating thermoplastic part. The driving tool spins until the timer runs out, causing the tool to retract. The frictional heat during spinning forms a melt layer on the joint surface to weld both contact surfaces. • As the driving tool retracts, the grip teeth disengage, the pivot pin stops spinning instantly because of joint friction (self braking). The joint remains pressured until the pivot pin disengages, which is sufﬁcient time for the thin melt ﬁlm in the joint to solidify.

Cap spins press body

Grip teeth spin andpress down the cap

Spinning cap presses against body causing frictional heat

When the part design does not permit the use of a pivot pin, the pivoting feature is accomplished by means of a thrust bearing design. The driving tool crown teeth could leave indentations on the surface of the thermoplastic cap. This may not be desirable in some applications, a rubber liner (holder) inside the driving tool is used to grip the peripheral surface of the thermoplastic spinning cap to transmit the maximum torque. To achieve a tight and strong weld joint, the pivot thrust load is very important and should be determined experimentally for each application. For this reason, an adjustable spring is recommended to provide the variation in load necessary for different parts. To reach the melting point, the product of pressure and velocity must exceed a certain minimum. If this value is too low, no melt will be generated and the joint surfaces will not be welded. If the value is too high, melting temperature is reached so quickly that it becomes difﬁcult to adjust the weld time. As a basic rule, joint velocity should range between 20 and 40 ft/s. The thrust pressure is then adjusted to obtain correct joint weld strength with a cycle time between 0.5 to 1.0 s. A spin pivot welding device must be equipped with the following controls: • Rotating speed • Thrust load

Tool retracts and stops spinning

Figure 8-106 Spin pivot welding process sequence

• Weld time • Pressure hold down time • Total cycle time • Vertical driving tool speed Due to these requirements, the spin pivot welding method is not as practical for automatic assembly processes as the spin inertia welding method. The pivot tools can be used on hand operated drill presses for short runs, prototypes, and repair work. However, the spin pivot method is applicable for the following special welding applications:

479

8.12 Spin Welding Method • Parts having small diameters that would require excessive rotating speeds with the spin inertia method • Very thin walled parts, unable to withstand the high thrust load required for the spin inertia method 8.12.4.2

Spin Inertia Welding Process

The spin inertia welding process is based on the principle of accelerating a precalculated steel mass (ﬂying wheel) to a speed that generates sufﬁcient kinetic energy to obtain the required weld strength. The process starts by spinning the driving tool (fast) to a desired speed and then disengaging it from the motor and moving it downwards at high speed. The kinetic energy stored in the ﬂying wheel and the thrust load pressing the two thermoplastic parts together is converted into frictional heat at the joint surfaces during the spinning motion. It is essential that the thrust pressure be sufﬁcient to stop the rotating ﬂying wheel mass within a fraction of a second. Too low a thrust pressure causes the ﬂying wheel mass to rotate longer, shearing off the solidifying molten resin and producing poor weld strength. The driving tool is attached to the piston rod and engaged to a running motor by means of a clutch. The piston stops in an intermediate position while retracting, allowing time to load the thermoplastic components. An automatic operation requires a quick acting clutch to bring the driving tool (ﬂywheel) up to speed. This may be accomplished by placing the clutch above the air actuator so that, as the driving tool advances, the clutch disengages allowing the driving tool to spin freely. Retraction of the driving tool after welding will reengage the clutch. Basic spin inertia welding process guidelines are listed in the following: • Rotating speed is a function of the diameter which obtains a linear velocity on the joint between 20 to 40 ft/s. In some cases (for small diameters), lower values have been used; however, it is of advantage to work at higher speeds whenever possible. • The rotational speed can be adjusted to balance the spin inertia welding process. • Considering the rotating speed and the joint surface, the driving tool (ﬂywheel) weight can be approximated by using 50 to 200 pounds of weight for each square inch of weld area (within 2.0 lbs) The mass of the ﬂywheel should be kept to a minimum for quick braking.

Air actuator

Flywheel

Driving tool Plastic cap Plastic body Centering chuck

Spin inertia welding tool

• The time to stop the ﬂywheel mass from rotation should be about 0.10 to 0.30 s by means of a brake forming a small amount of ﬂash around the weld. First, estimate a speciﬁc pressure to achieve 200 psi on the projected joint area. If the ﬁrst trials do not give satisfactory results, the kinetic energy can easily be adjusted by varying the rotating speed. Too much energy will only produce more ﬂash without affecting joint strength if the thrust load is sufﬁciently high. Too low a thrust load would give poor results, because it would increase the weld time. • Combine several spin welding stations with turntables and feeding devices; the process is most adaptable to fully automatic assembly lines. • For short runs, trials, and prototypes, a simple tool can be ﬁtted onto a high speed drill press table and hand operated.

Single station spin inertia welder

• For high volume production, the use of a commercial spin inertia welding machine is highly recommended.

Figure 8-107 Spin inertia welding assembly process equipment

480

8 Thermoplastic Assembly Methods Figure 8-107 shows several types of spin inertia welding assembly process equipment used for the assembly of injection molded thermoplastic components.

Automotive intake manifold Centering fixture

8.12.5

Joint area End cap

Driving tools

Work table

Spin Welding Joint Designs

Spin welding joint size and geometry is a function of the wall thickness, weld strength, and appearance of the product.

Air actuators

Figure 8-108 shows the recommended circular joint design for an injection molded thermoplastic thin-walled cylinder. Figure 8-109 shows a circular joint design for an injection molded thermoplastic cylinder with thick walls.

Controller T1

T2

T3

T1

T2

T3

0 0

0 1 0

1 1

00

0 1 0

1 1

Two spin inertia welder stations for simultaneous assembly of end caps

The presence of ﬂash at the joint after welding is essential to insure a good weld. In cases where ﬂash is objectionable from a functional or appearance standpoint, Figure 8-110 shows four spin welding joints with ﬂash trap designs. If it is not possible to use a spin welding joint with ﬂash trap, the ﬂash can be removed manually by means of ﬂash trimming knives. For small molded parts, this can be done by conveying the plastic part through an automatic vibrating screen, where using a sharp edge the ﬂash and burrs are sheared off.

8.12.6

Spin Welding Process Suggestions

Setting Up the Cycle • Minimum speed of 20 ft/s at the joint surface area • Driving tool rotation should stop within 0.1 to 0.3 s of welding Single spin inertia welder station with six positions rotary turntable

Figure 8-107 (continued)

• A uniform external ﬂash should be visible around the weld joint Flash Removal • A ﬂy knife attachment on the driving tool can be used to remove ﬂash simultaneously during the spin welding process • Providing a shielding skirt may hide the ﬂash (without trimming) • Conveying the small plastic parts through a vibrating screen. Evaluation of the Spin Weld • For transparent resins or translucent semi-crystalline molded parts having thin walls it is practical to inspect the weld visually. The weld should appear smooth; any change in coloration, marks, spots, or voids are indications of poor quality weld. • A reduction in axial length of 0.015 to 0.025 in measured from two points, one on either side of the weld, usually indicates that a good weld has been formed. • Some type of end use test is suggested for checking the weld quality. For instance, a pressure vessel should be burst- and leak-tested. Other parts may need to be impacted, pulled apart (tension), or sheared off at the weld. Once the proper welding conditions have been found, normal quality control sampling methods may be employed.

481

8.12 Spin Welding Method

5˚

T

0.2 T 1.5 T 32˚

5˚

1.8 T

14˚ 0.5 T 0.4 T 0.5 T 28˚

1.9 T

16˚ T

Figure 8-108 Spin welding joint design for thin-walled cylinders

T 0.6 T 0.6 T

1.8 T

0.4 T 1.1 T 0.1 T

0.9 T

28˚

T

0.1 T

0.8 T

32˚

Figure 8-109 Spin welding joint design for thick-walled cylinders

Figure 8-110 Spin welding joint with flash trap designs

482

8 Thermoplastic Assembly Methods

8.13

Steam iron and pump housing

Automotive dash panel

Ultrasonic Welding Method

Over the past seventy ﬁve years, the ultrasonic welding technology has become an important method for countless applications in virtually every industry where injection molded thermoplastic resins are used. The most signiﬁcant recent developments have been in microprocessor technology that have further enhanced the controls for automation, while resulting in quantum leaps in precision and the quality of the ﬁnal assembled product. Ultrasonic assembly is a fast, clean, and efﬁcient method for assembling injection molded rigid thermoplastic components.Various ultrasonic assembly techniques are used by all segments of the industry to join plastic to plastic and plastic to nonplastic materials thereby replacing or precluding the use of adhesives and solvents, mechanical fasteners, or other welding methods. Advantages and beneﬁts in ultrasonic welding process include: • Energy efﬁciency • No need for elaborate ventilation systems to remove fumes or heat

Wrist watch cases

• High productivity with lower cost than many other assembly methods • Ease of interface with automated assembly line production • Immediate start-up and shut-down with no residual heat. Ultrasonic welding applications are shown in Figure 8-111.

8.13.1 Battery housing

Audio and video cassette and computer microdisk

Computer dip switch

Figure 8-111 Ultrasonic welding thermoplastic assembly applications

Ultrasonic Welding Basic Principles

The ultrasonic welding equipment used to assemble thermoplastics, converts 50/60 Hz current to 20 kHz or 40 kHz electrical energy through a solid state power supply. This high frequency electrical energy is supplied to the converter, a component that changes electrical energy into mechanical vibratory energy at ultrasonic frequencies. The vibratory energy is then transmitted to the horn through an amplitude modifying device called a booster. The horn is an acoustic tool that transfers this vibratory energy directly to the thermoplastic components being assembled. A ﬁxture supports and aligns the thermoplastic components to be welded. Generally, it is fabricated from aluminum or steel and sometimes it is lined with cast urethane or other resilient materials. The actuator contains the converter, booster, horn, and pneumatic controls. Its function is to bring the horn into contact with the work piece, apply appropriate force, and retract the horn after the weld cycle is completed. The vibrations are transmitted through one of the mating thermoplastic components to the joint area, where vibratory energy is converted to heat by friction that melts the thermoplastic. A combination of applied force, surface friction, and intermolecular friction at the joint interface elevates the temperature until the melting point of the material is attained. Force is maintained after vibration ceases, and a molecular bond (weld) is produced. Cycle times usually are less than 1 s and weld strength obtained approaches that of the parent material. In ultrasonic welding, the energy required to accomplish a weld is the product of the average power dissipated in the joint and the weld time. Microprocessor

483

8.13 Ultrasonic Welding Method control now enables welding by constant energy, varying the time to deliver the predetermined amount of energy required to produce the best quality welds.

8.13.2

Microprocessor

Ultrasonic Welding Basic Components

Figure 8-112 shows the various components required for ultrasonic welding systems. Vibrations are introduced into the thermoplastic parts by the ultrasonic welding process and the horn causes several types of waves. In a longitudinal wave, the direction of vibration of individual particles is the same as the direction of wave movement. In some solids, it is also possible for a sound wave to have a vibration amplitude perpendicular to the direction of propagation of the sound wave. Such a wave is called a shear wave. In a third type of wave, the particles on the surface of a solid move in an elliptical path. Thermoplastic components transmit longitudinal waves with greater ease than they transmit shear waves; ultrasonic welding is the result of the introduction of longitudinal vibrations into the parts to be welded. Vibratory energy is transformed into frictional heat at the joint or interface between two parts. Scientiﬁc disagreement exists in the deﬁnition of how the heat is created at the joint by the ultrasonic mechanism. One theory states that when two rigid parts or two layers are vibrated against each other, surface friction develops, causing localized surface frictional heating. The longitudinal waves introduced into the parts are converted into surface waves in the joint and the energy caused by rubbing the joint surfaces of the thermoplastic parts generates frictional heat. When sufﬁcient heat is generated at the joint interface, softening and melting of the thermoplastic contacting surfaces occurs in the parts. Welding is the result of the thermal and mechanical energy agitating the molecules during the melting stage, forming bonds at the joint surfaces when pressure is applied. A weld is formed at the joint once the melt cools down when the energy stops.

8.13.3

Power supply

Ultrasonic Welding Equipment

Equipment required for ultrasonic welding is complex and sophisticated in comparison with the equipment needed for other welding processes. An ultrasonic welding system is composed of an electronic power supply, timers to control the cycle, an electrical to mechanical energy transducer, a welding horn, and a part holding ﬁxture, which may be automated. Commercially available ultrasonic welding equipment is shown in Figures 8-113 and 8-114. Standard ultrasonic welders are designed in a variety of conﬁgurations to meet speciﬁc needs, such as versatility, durability, reliability, and ease of service. Integrated units are composed of a power supply and pneumatic actuator assembly and stand in a self-contained bench model unit. They are available in 20 kHz with power outputs of 450, 900, 1,500 and 1,800 W. Component systems allow the power supply to be located away from the work station. Power supply and actuator combinations are available in both 20 kHz and 40 kHz frequencies, power ranges from 150 to 3,200 W, depending on the application, allowing maximum setup ﬂexibility and system integration.

Bench press structure Transducer Mounting piece Welding horn Plastic cap Plastic body Air eject

Holding fixture

Figure 8-112 Ultrasonic welder basic components

484

8 Thermoplastic Assembly Methods

Figure 8-113 Commercially available ultrasonic welders

Figure 8-114 Commercially available ultrasonic welders

Ultrasonic welder with transducer magnetostrictive

Ultrasonic welder branson controller microprocessor 900M

Ultrasonic welder dukane press system

Ultrasonic welder branson bench rotary dial

Hand-held welding systems in 20 and 40 kHz are available for use with power supplies ranging from 150 to 1,000 W. The tools are lightweight and allow ultrasonic assembly of large parts or assembly of parts with hard to reach joint areas. Ultrasonic Welder Power Supply In most commercial ultrasonic welders, the power supply generates a 20 kHz or 40 kHz electrical output, ranging from a hundred to a thousand or more watts

485

8.13 Ultrasonic Welding Method of rated average power. The power supplies are solid-state operation in an openloop format; the introduction of the microprocessor greatly enhances the welding process. The power supplies operate at lower voltages and have impedances close to those transducers to which the power supply is connected. Ultrasonic Welder Microprocessor Controller During ultrasonic welding, many variables can affect the strength and the quality of the weld. If the ultrasonic welder has been properly set up for the application, the signiﬁcant variables are limited to the ultrasonic equipment. These may include weld time, hold time, operating pressure, stack (converter, booster, horn), and power draw. Conventional welding equipment operates in an open-loop format. That is, a desired weld time, hold time, and operating pressure are selected, whether the parameters are correct or the welding equipment is in good operating condition. The equipment is not sophisticated enough to operate with the feedback control, nor can it conﬁrm if the desired operating conditions are being met with any degree of precision. The introduction of the microprocessor control to ultrasonic welding of thermoplastic had a signiﬁcant impact on the technology and enabled quantum leaps in ultrasonic assembly as a manufacturing tool. The user now has control over electronic and mechanical functions that determine part to part weld consistency and often enable the successful assembly of those applications that otherwise may not have been possible to join due to the injection molding process or part dimension tolerance problems. Manufacturers now can control the tolerances between thermoplastic parts for assembly much better than ever before and control the weld as precisely they can control the injection molded thermoplastic parts. The ultrasonic welding microprocessor provides time or energy welding with ﬁnite timing. The user sets times accurately in ﬁve millisecond increments or energy in one joule increments and determines the precise force required before ultrasonic waves are triggered on and off as well as the duration of ultrasonic waves. In addition, ﬁnite force buildup before triggering of ultrasonics is provided by a load cell, with air pressure readings on a digital display. Welding by distance (meltdown) with a resolution of 0.0001 in is achievable. Along with this absolute control, the user also can measure and plot energy, force, and velocity for each cycle and weld in modes with limits that best suit the application. It is possible to apply as little as 10 ms or 1 J of energy to a molded thermoplastic part with a 20 kHz or a 40 kHz welder, or to apply 2,000 W for 10 s with the same welder. All of these controls and limits are now accurate and repeatable, cycle to cycle.

20 or 40 kHz input

Cooling fan Piezoelectric elements

Ultrasonic Welder Transducer The transducers used in ultrasonic welding are electro-mechanical devices used to convert high frequency electrical oscillations into high frequency mechanical vibrations through either piezoelectric or magnetostrictive effects. The piezoelectric material changes length when an electric voltage is applied across it. The material can exert a force on anything that tries to keep it from changing dimensions, such as the inertia of some structure in contact with the material. These transducers have a conversion efﬁciency near 95%. They are cooled only by forced air. A typical piezoelectric transducer is shown in Figure 8-115.

Mounting piece

Welding horn

Figure 8-115 Ultrasonic welder, piezoelectric transducer

486

8 Thermoplastic Assembly Methods Elastomeric pad "OUT" cooling liquid

20 or 40 kHz input Magnet and coil

Transducer core Mounting cone

Seal

"IN" cooling liquid

Welding horn

Figure 8-116 Ultrasonic welder, magnetostrictive transducer

Magnetostrictive material exhibits a similar behavior, but under the inﬂuence of a magnetic ﬁeld rather than an electrical ﬁeld. Natural crystals such as quartz and specially treated ceramic materials are piezoelectric. All magnetic materials are magnetostrictive; however, only those generally based on nickel alloys exhibit this effect. Magnetostrictive materials will expand (or contract) independently of the direction of the magnetic ﬁeld. Hence, a biasing ﬁeld provided either by a permanent magnet in the transducer core or by a direct current in the coil is necessary to keep the core vibrations in frequency with the driving electrical input. The transducer core is bonded by silver soldering or bolting to a mounting bar that has an approximate quarter wavelength. The total assembly is one-half wavelength and therefore resonant at the 20 kHz or 40 kHz frequencies. Heat is produced in the transducer core as a result of the mechanical and magnetic hysteresis, the eddy currents, and the resistive losses in the coil. Magnetostrictive transducers have a conversion efﬁciency of less than 50%. Therefore, a liquid cooling system is required to remove the heat from the transducer. By the nature of this construction, magnetostrictive transducers are rugged. A typical magnetostrictive transducer is shown in Figure 8-116. Ultrasonic Welding Horn An ultrasonic welding horn is attached to the output end of the transducer. The ultrasonic welding horn has two functions: it introduces ultrasonic vibrations into parts being welded and it applies the pressure necessary to form a weld once the joint surface has been melted.

Figure 8-117 Typical ultrasonic welding horns

The thermoplastic parts represent a load or impedance to the ultrasonic transducer. The welding horn serves as a means to match and to increase the amplitude of vibrations from the transducer (impedance matching transformer) to the load. As a measure of ampliﬁcation, total movement or double amplitude of the transducer output may be approximately 0.0005 in, while the vibrations suitable for ultrasonic welding range from 0.002 to 0.006 in. Ampliﬁcation or gain is one factor in establishing the design of the welding horns. Typical ultrasonic welding horns are shown in Figure 8-117.

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8.13 Ultrasonic Welding Method Interconnecting horns will increase or decrease the amplitude of vibrations of the last horn in the series. The middle horn, located between transducer and welding horn, is called a booster horn and is a convenient way to alter amplitude, an important variable in ultrasonic welding. Vibration horn material requirements are the ability to sustain large motions without failure. High strength titanium alloys are the best; other suitable materials are Monel metal, stainless steel, and aluminum. Materials that dissipate acoustic energy, such as copper, lead, nickel, and cast iron, are not suitable. For an efﬁcient ultrasonic welding process, the horns must resonate at a frequency very near the nominal 20 kHz or 40 kHz operating frequency of the welding system. The ultrasonic welder suppliers manufacture a unit for electronically tuning the welding horns, to make variations in horn dimensions to achieve optimum performance. The design and fabrication of complex horns using sophisticated materials should be left to equipment manufacturers with experience and capabilities in analytical design of welding horns. The use of simple aluminum step horns should be limited to the evaluation of prototype ultrasonic welds in a laboratory. Ultrasonic Welding Holding Fixture Fixtures for aligning and holding the thermoplastic parts stationary during welding are an important aspect of the ultrasonic welding equipment. Parts must be held in alignment around the end of the welding horn so that uniform pressure between parts is maintained during welding. If the bottom part to be welded is simply placed on the welding table, both parts may slide out from under the horn during welding. High frequency vibrations reduce the effect of nominal frictional forces that might be generated if the part was held stationary. A typical ﬁxture is shown in Figure 8-118. Ultrasonic welding holding ﬁxtures are machined or cast so that the ﬁxture engages the lower part and holds it securely in the desired location. The holding ﬁxture should be rigid so that the relative motion is developed between the thermoplastic base and the vibrating cap. This can be achieved by making the holding ﬁxture massive and tune to a quarter wavelength. Flatness or wall thickness variations in some molded thermoplastic parts could prevent consistent welding; the use of an elastomeric liner inside the holding ﬁxture is used to overcome this joint surface alignment problem. High production ultrasonic welding applications require the use of automated part handling equipment and holding ﬁxtures. For small thermoplastic parts, vibrating hoppers and feeders are used to feed parts onto an indexing table equipped with multiple ﬁxtures for holding parts. Several ultrasonic welding stations are used at sequential positions around the indexing table.

8.13.4

Ultrasonic Welding Process Variables

The major ultrasonic welding process variables are: weld time, hold time, amplitude of vibration, and thrust pressure.

Welding horn Plastic base Holding fixture

Plastic cap Air ejection

Figure 8-118 Ultrasonic welding holding fixture

488

8 Thermoplastic Assembly Methods Weld Time Weld time is the period during which the ultrasonic vibrations are applied. The correct weld time for each ultrasonic welding application is determined by trial and error. It is important to avoid over-welding. In addition to creating an excessive amount of ﬂash that may require trimming, over-welding can degrade the quality of the ultrasonic weld and create leaks in applications requiring a hermetic seal. The welding horn can mark the thermoplastic part surface. Also, melting and fracture of portions of the parts away from the joint area may occur when using longer weld times, especially at holes, weld lines, and sharp corners in the injection molded thermoplastic parts. Hold Time Hold time is a nominal period after ultrasonic welding during which the thermoplastic parts are held together and allowed to solidify under thrust pressure without vibrations. Hold time is not a critical variable in most applications. A hold time of 0.3 to 0.5 s is usually sufﬁcient for most ultrasonic welding applications, unless an internal load tends to disassemble welded parts, such as a coil spring compressed before welding. Amplitude of Vibrations The physical amplitude of vibrations applied to the thermoplastic parts being welded is an important ultrasonic welding process variable. High amplitude of vibration (approximately 0.004 to 0.006 in, peak-to-peak) is necessary to achieve efﬁcient and rapid energy input into the thermoplastic parts. Because the basic transducer delivers its power at high force and low amplitude, the amplitude must be stepped up before reaching the welding horn tip. The welding horn design usually includes amplitude transformation inherent in tapering or stepping its proﬁle down to a small diameter. Where the thermoplastic part geometry requires a large or complex tip shape, this ampliﬁcation may not be possible in the welding horn. In this case, ampliﬁcation can be conveniently achieved in most commercial systems by use of an intermediate tuned section called a booster horn. Booster horns up to an ampliﬁcation ratio of 2.5 : 1.0 are commercially available. Negative boosters of 0.4 : 1.0 for horns having a high amplitude for a given application are also available. Boosters that provide a 2.0 : 1.0 or 2.5 : 1.0 ampliﬁcation are typically required for semi-crystalline thermoplastic parts, except for small parts that permit the use of a high gain welding horn. Increasing amplitude improves weld quality in semi-crystalline parts designed with shear joints. For amorphous parts with energy director type joints, weld quality is increased and weld time is reduced by increasing the amplitude. Thrust Pressure Trust pressure provides the static force necessary to couple the welding horn to the thermoplastic parts so that vibrations may be introduced between the two joint surfaces of the parts. This same static load ensures that parts are held together as the melted polymer in the joint solidiﬁes during the hold portion of the welding cycle. The determination of the optimum thrust pressure is essential for good ultrasonic welding. If the thrust pressure is too low, the equipment is inefﬁcient and results in unnecessarily long welding cycles. If it is too high with the welding horn tip amplitude, it can overload and stall the welding horn and dampen the vibrations. The overall amplitude gain provided by the booster and the welding horn is analogous to the load matching provided by the gear

8.13 Ultrasonic Welding Method ratio between an automobile engine and its rear wheels. In ultrasonic welding, low thrust pressure is required with high amplitude and high thrust pressure is required with low amplitude.

8.13.5

Ultrasonic Welding Joint Designs

The most critical detail of an injection molded thermoplastic part design for ultrasonic welding is the joint design (the conﬁguration of the two mating surfaces). There are a variety of joint designs, each with speciﬁc features and advantages. Their selection is determined by such factors as type of thermoplastic, part geometry and size, weld requirements, molding capabilities, moisture, and cosmetic appearance. Proper joint design is essential for optimum ultrasonic welding results. The thermoplastic material, part geometry, and weld requirements all are design factors that must be considered, particularly with materials that have a semicrystalline structure and a high melting point. Joint design is less critical when ultrasonically welding amorphous thermoplastic materials. A joint should have a small and uniform initial contact self-alignment area to concentrate the ultrasonic energy for rapid and localized energy dissipation. There are two basic types of ultrasonic joint designs, the shear joint for semicrystalline resins and the energy director joint for amorphous resins. The shear joint is used when a high strength hermetic seal is required. During welding, the interfaces melt and telescope together producing an in-weld shear. This design is particularly recommended for semi-crystalline materials. The energy director is used for amorphous resins; it consists of a small triangular bead in one of the joint surfaces. Good weld strength can be achieved with amorphous resins; however, poor weld strength is obtained with semi-crystalline resins, because of their fast crystallization rate, that causes the molten energy director to freeze-off before the weld is completed. 8.13.5.1

Ultrasonic Welding Shear Joint Design

Shear joints are the recommended design for ultrasonic welding applications with semi-crystalline thermoplastic polymers. The shear joint design was developed by Du Pont in 1967 and has been used worldwide very successfully in many applications. Shear joints provide high strength hermetic seals of various conﬁgurations with high melting point semi-crystalline thermoplastic resins. The basic shear joint, chamfer shear joint, and variations of the basic shear joint for large parts to be ultrasonically welded are shown in Figure 8-119. The initial contact is limited to a small area that is usually a recess or step on either or both sides of the thermoplastic parts. The contacting surfaces melt ﬁrst; then, as the parts telescope together, they continue to melt along the vertical walls. The smearing action of the two melt surfaces eliminates leaks and voids, making this the best joint for strong hermetic seals. Several important aspects of the shear joint should be considered: • The top thermoplastic part should be as shallow as possible • The outer walls should be well supported by a holding ﬁxture • The design should allow for a clearance ﬁt (0.005–0.008 in)

489

490

8 Thermoplastic Assembly Methods

T

5°

B

C

A

T Holding fixture

Basic shear joint T C

5° B

A

T

30°- 45° Holding fixture

Chamfer shear joint

A

Holding fixture

Shear joint for large parts

A

Holding fixture

Shear joint for large parts Figure 8-119 Ultrasonic welding basic shear joint design variations

Part size

Interface “A”

≤ 0.75 in 0.75–1.50 in ≥ 1.50 in

0.008–0.012 in 0.012–0.016 in 0.016–0.020 in

Part wall thickness = “T” Lead-in “B” = 0.016–0.024 in Weld depth “C” = (1.25–1.50) T

491

8.13 Ultrasonic Welding Method • A lead-in “B” should be incorporated in the product design. A lead-in is required for self locating • A ﬂash trap can be incorporated if necessary The chamfer shear joint requires lower vibration energy and the shortest welding time of the various shear joints. This is due to the small initial contact area of the top part and tapered self-locating lead-in wall in the bottom part, and the uniform progression of the weld as the thermoplastic melts and the parts telescope together. Heat generated at the joint is retained until the vibrations cease because, during the telescoping and smearing action, the melted thermoplastic is not exposed to air, which would cool it too rapidly. Weld strength is therefore determined by the depth of the telescoped section, which is a function of the weld time and part design. Joints can be made stronger than the adjacent walls by designing the weld depth “C” of telescoping as 1.25 to 1.5 times the part wall thickness “T” to accommodate minor variations in the injection molded thermoplastic parts. Several important aspects of the shear joint must be considered; the top part should be as shallow as possible (just a lid). The walls of the bottom section must be supported at the joint by a holding ﬁxture that conforms closely to the outside conﬁguration of the part to avoid expansion under the welding pressure. Noncontinuous or inferior welds result if the upper part slips to one side or off the lower part, or if the stepped contact area is too small. 8.13.5.2

Chamfer Shear Joint Design Variations

Figure 8-120 shows variations of the chamfer shear joint design.

Holding Fixture

Holding Fixture

Holding Fixture

Figure 8-120 Chamfer shear joint designs variations

The ﬁt between the two thermoplastic parts should be close before welding, but not tight. The shear joint requires 300 to 400% more weld time than other ultrasonic joint designs, because larger amounts of polymer are melted to form a shear joint weld. A certain amount of ﬂash will also be visible on the shear joint after welding. 8.13.5.3

Basic Shear Joint Design with Flash Trap

Allowance should be made in the design of the shear joint for the ﬂow of molten material displaced during the ultrasonic welding. When ﬂash cannot be tolerated for aesthetic or functional reasons, a ﬂash trap as shown in Figure 8-121 can be designed into the shear joint.

492

8 Thermoplastic Assembly Methods 0.005 inch 0.008 inch Flash trap Flash trap

T

Flash trap Flash Trap

0.25T

Figure 8-121 Basic shear joint designs with flash trap

0.12T

90˚

8.13.6

Ultrasonic Welding Energy Director Butt Joint

The butt joint with an energy director is the most common joint design used with the amorphous thermoplastic resins in ultrasonic welding. The main feature of this joint is a small 90° or 60° triangular shaped bead molded into one of the mating surfaces. This energy director limits the initial contact to a very small area and focuses the ultrasonic energy at the apex of the triangle. During the welding cycle, the concentrated ultrasonic energy causes the thermoplastic bead to melt and the thermoplastic to ﬂow throughout the joint area, bonding or welding the parts together.

90˚ Energy director for thick walls T

0.25T

0.22T

60˚ 60˚ Energy director for hermetic seal

T

For easy to weld amorphous polymers, such as ABS, SAN, acrylic, polystyrene, and others, the size of the energy director is dependent on the area to be joined. Practical considerations suggest a minimum height of 12% of the molded part wall thickness for the 90° energy director and a minimum height of 22% of the wall thickness for the 60° energy director. Semi-crystalline polymers with high melting points, such as nylon, thermoplastic polyesters, acetal, polyethylene, polypropylene, polyphenylene sulphide, and others, as well as the high softening temperatures of amorphous resins such as polycarbonate and polysulfone are more difﬁcult to weld ultrasonically. For these high temperature resins, the 60° included angle energy directors with a minimum height of 22% of the part wall thickness is recommended. Energy Director Joint Design The 90° included angle energy director height should be at least 12% of the joint wall thickness and the width of the energy director should be at least 25% of the joint wall thickness. Figure 8-122 (top) shows a butt joint with a 90° included angle energy director. With thick-walled joints, two or more energy directors should be used and the sum of their heights should equal 12% of the joint wall thickness.

0.015T

Size reduction after welding Figure 8-122 Type of energy directors butt joint designs

To achieve hermetic seals when ultrasonically welding amorphous components, it is recommended to design a 60° angle energy director into the part. The energy director width should be 25% of the joint wall thickness. Figure 8-122 (middle) shows a butt joint with a 60° included angle energy director. Figure 8-122 (bottom) shows how the parts should be dimensioned to allow for the ﬂow of the molten material from the energy director throughout the joint area.

493

8.13 Ultrasonic Welding Method When ultrasonically welding components made of identical amorphous thermoplastic materials, the energy director can be designed into either side of the thermoplastic components’ joint surfaces. However, when designing the energy director into components made of different amorphous thermoplastic polymers; one component made of an amorphous copolymer or terpolymers, such as ABS, and the other component made of an amorphous homopolymer such as acrylic, to obtain the maximum strength in the welded joint, the energy director should always be located in the component made of the amorphous homopolymer material.

Before welding T

0.25 T

0.12 T

0.2 T Clearance 0.005 in. 0.008 in.

0.2 T

Energy Director Step Joint Design A step joint with 90° energy director is shown in Figure 8-123. The step joint molds without any problems. The step joint provides a strong, well aligned joint with minimum effort. This joint is usually stronger than a butt joint, because the polymer melt ﬂows into the vertical wall clearance, increasing the length of the weld by 20%. The step joint provides good strength in shear as well as in tension and is often recommended when cosmetic appearance is critical. When working with low melting temperature semi-crystalline materials, and when the end use welding requirements are not important, a 60° included angle energy director should be used instead of the 90° included angle energy director.

After welding

Joints with energy director are not recommended for use with high melting temperature semi-crystalline materials requiring high weld strength or repeatable hermetic seals. When these end use requirements are speciﬁed by the application, the basic shear joint or the chamfer shear joint should be considered. Figure 8-123 90° energy director step joint design

Energy Director Tongue and Groove Joint Design The ultrasonic welding tongue and groove joint with a 90° energy director is shown in Figure 8-124. This joint design provides an advantage in the welding process: the tongue and the energy director in the joint become a self-locating feature for ultrasonic welding. The melt ﬂash caused by ultrasonic welding does not over run internally or externally. The amount of melt produced during welding is controlled by the joint dimensions; the melt ﬂows through the channel and prevents the ﬂash from moving outside the joint area. This joint also provides the greatest weld strength of the ultrasonic welding energy director joints for amorphous thermoplastic resins. Figure 8-125 shows the ultrasonic welding basic 90° energy director step joint design variations. Before welding

After welding

T Clearance 0.005 in. 0.008 in.

0.5 T

10˚Taper walls 0.12 T 0.35 T

0.25 T

Figure 8-124 90° energy director tongue and groove joint design

494

8 Thermoplastic Assembly Methods Before welding

Before welding

Before welding

Before welding

After welding

After welding

After welding

After welding

Figure 8-125 90° energy director step joint variations

8.13.7

Ultrasonic Welding Method Design Limitations

The inﬂuence of the overall part design on ultrasonic welding feasibility is very signiﬁcant in the product development work. Some generalizations can be made about certain aspects of the part design and its effect on the success of the ultrasonic welding process. Determining the location at which the welding horn will contact the lid or top component surface to be welded is a very important aspect of part design. Welding Configurations Welding horn

Plastic body

Holding fixture

Plastic lid Joint

In the far ﬁeld welding conﬁguration, the welding horn contacts the main structure surface at a far distance from the joint where the lid is held. The ultrasonic vibrations are absorbed by the plastic without creating appreciable heat energy to transmit to the joint, therefore, making welding more difﬁcult.

Far field

Welding horn Plastic lid

Joint

Holding fixture

There are two ultrasonic welding conﬁgurations, the far ﬁeld and the near ﬁeld, which are deﬁned as the point of contact between the welding horn and the distance from the joint surface. The best ultrasonic welding results for all injection molded thermoplastic parts are obtained with the near ﬁeld welding conﬁguration. Therefore, whenever possible, parts should be designed for the lid top surface to directly contact the welding horn as close to the joint as possible.

Plastic body

Near field Figure 8-126 Far and near field ultrasonic welding configurations

These two welding conﬁgurations are shown in Figure 8-126. Commodity thermoplastics, such as polyethylene, can only be welded by the near ﬁeld welding conﬁguration. Because they have a high acoustic damping factor, they strongly attenuate the ultrasonic vibrations upon entry into the material. If the joint is too far from the welding horn, the energy is not transmitted to the joint and the thermoplastic melts only at the interface with the welding horn. Thermoplastic materials are poor transmitters of shearing waves. This fact makes ultrasonic welding more difﬁcult when the geometry of the plastic body is complex. Vibrations are partially attenuated or dissipated at bends, angles, or discontinuities, such as holes in the plastic body. To maximize transmission of vibrations, the thermoplastic parts should be designed with a ﬂat contacting surface for the welding horn. This surface should be

8.13 Ultrasonic Welding Method as broad as possible and continuous around the joint area. Interruptions in contact between the welding horn and the part may result in weld interruptions. Since the entire structure of both thermoplastic components being welded is subjected to vibrations, a very high level of stress occurs at sharp internal corners, resulting in fracture or sporadic melting of the ultrasonic weld. It is difﬁcult for the welding horn to transmit ultrasonic vibrations uniformly through the thermoplastic material to several joint areas with different sizes, located either in the same plane or in different planes. The welding horns are designed and tuned to provide only the speciﬁc amount of heat to a joint located close to the horn. Ultrasonic welding should be performed only one joint at a time; when several joint areas are welded at the same time, the vibrating energy is sporadic and different for each joint, resulting in poor quality welds and probably incomplete welds in all joint areas. Moisture Absorption Nylon resins absorb more moisture from the air after molding than most other thermoplastic resins. Wet nylon is more difﬁcult to weld and the moisture content can seriously affect weld quality. Other resins are also hygroscopic, for example, polymers such as ionomer, polyester, polycarbonate, and polysulfone. These materials are not as problematic as nylon, but they also require drying of the molded parts before ultrasonic welding. If hygroscopic parts are allowed to absorb moisture, during welding the water (steam) will evaporate at 212 °F, with the trapped gas creating porosity (foamy condition) and often degrading the polymer, producing a lower molecular weight and poor weld quality at the joint interface. If welding cannot be done immediately after molding, nylon parts should be kept in a dry-as-molded condition prior to welding. Exposure of one or two days to 50% relative humidity at 73 °F is sufﬁcient to degrade weld quality by 50% or more. Nylon parts may be kept dry for periods up to several weeks by sealing them in polyethylene bags immediately after molding. For longer periods, greater protective measures must be taken, such as the use of heat-sealable moisture barrier bags. Nylon parts that have absorbed moisture may be dried before welding in a dehumidifying drying oven; however, possible heat degradation of the parts must be avoided. Welding parts at longer than normal welding cycles may offset the quality of the weld, but often at the expense of heavy weld ﬂash, marring under the welding horn, and causing severe problems. Reground Material Scrap material formed during the molding process (sprues, runners, reject parts), can usually be recycled directly back into the process after the material has been reduced to a usable size. Control over moisture, the amount, and the quality of reground is necessary, because they can adversely affect the ultrasonic welding characteristics of the molded part. In some cases, the use of 100% virgin resin may be required to obtain the desired results. Welding Dissimilar Resins A similar melt temperature between the materials to be welded is a basic requirement for successful welding of rigid parts. A temperature difference of 40 °F can be sufﬁcient to affect the quality of the weld. The lower melt temperature material

495

496

8 Thermoplastic Assembly Methods melts and ﬂows, thus preventing the generation of sufﬁcient heat to melt the higher melt temperature material. For example, with an energy director on a part composed of high temperature acrylic opposing a parallel surface composed of a low temperature acrylic, the weld surface of the high temperature acrylic part will not reach the necessary temperature to melt. The opposing surface will be in a molten state before the energy director begins to soften, and if the energy director fails to melt, bond strength will be impossible to obtain. In order to successfully weld dissimilar resins, both thermoplastics to be welded must have the same molecular structure and they should be chemically compatible. The percentages of the chemical radicals present will determine the molecular compatibility and bond strength between dissimilar plastics. Compatibility exists only among amorphous resins or blends containing amorphous resins. Typical examples are polycarbonate to acrylic, ABS to acrylic, and polystyrene to modiﬁed polyphenylene oxide. Grades of Resins The different grades of polymers can have a signiﬁcant inﬂuence on weldability because of melt temperature and other property differences. An example is the difference between injection/extrusion grades and cast grades of acrylic. The cast grade has a higher molecular weight and melt temperature, is often brittle, and forms a skin that gives it greater surface hardness, which reduces the weldability to the injection grade. A general rule of thumb is that both materials to be welded should have similar molecular weights and melt temperatures within 40 °F of each other.

8.13.8

Weldability of Thermoplastic Materials

There are a number of factors that affect the transmission of ultrasonic energy and, therefore, the efﬁciency for welding various types of thermoplastic materials. Because polymers differ in their physical structure, melt characteristics, and stiffness, ultrasonic energy requirements are determined on an individual basis. The major factors include polymer structure, melt temperature, modulus of elasticity (stiffness), and chemical characteristics. Thermoplastic polymers are categorized according to their physical structure: the amorphous thermoplastic resins and the semi-crystalline thermoplastic resins. Amorphous Thermoplastic Resins Amorphous polymers (polystyrene, ABS, PPO, polycarbonate, etc.) are very energy efﬁcient regarding their ability to transmit ultrasonic vibrations and can be welded under a wide range of force/amplitude combinations. They are characterized by a random molecular arrangement and a broad range of softening temperatures (Tg, glass transition temperature) that allow the material to soften gradually. This allows the material to ﬂow easily without premature solidiﬁcation of the polymer melt. Semi-Crystalline Thermoplastic Resins Semi-crystalline polymers (acetal, PE, PP, nylon, PBT, PET, etc.) require higher ultrasonic energy levels because of their very orderly molecular structure, sharp melting point, and high heat of fusion. The molecules of the resin, when in the solid state, are spring-like and internally absorb a percentage of the high frequency mechanical vibrations, thus making it more difﬁcult to transmit

497

Figure 8-127 shows the thermal behavior differences between amorphous and semi-crystalline resins. Amorphous resins become softer with temperature, they do not have a melting point, while semi-crystalline resins need a large amount of heat fusion energy to change from solid to liquid stage for molding.

100

Semi-crystalline resin

Tg

83

Tg

66

Amorphous resin

50 33

Tm

16 0 0

16

33

50

66

83

100

Temperature, (%) 100

Heat of fusion, (%)

ultrasonic energy to the joint interface. For this reason, high amplitude is usually required to satisfy the high heat of fusion. The sharp melting point is the result of a very high energy requirement (high heat of fusion) necessary to break down the semi-crystalline structure to allow the melt to ﬂow. Once the molten polymer leaves the heated area, the melt solidiﬁes rapidly (crystallization) with only a small reduction of the melt temperature. Successful ultrasonic welding results are obtained by using higher amplitudes, applying the basic shear or the chamfered shear joint designs, establishing the near ﬁeld horn contact distance of 0.25 in or less, and a close tolerance support ﬁxture.

Modulus of elasticity, (%)

8.13 Ultrasonic Welding Method

Semi-crystalline resin

83 66

Amorphous resin

50 33 16 0

Melt Temperature The glass transition temperature (Tg) is a measure of the temperature at which the noncrystalline portions of the polymer change from a glassy state (at low temperature) to a rubbery state (at higher temperatures). The Tg represents roughly the center of a transition region. However, if the polymer molecules are bonded by strong secondary bonds, the bonding will interfere with such motion. Semi-crystalline polymers in the temperature range between Tg and Tm (melting point) are referred to as leathery, because they are made up of a combination of the rubbery noncrystalline regions and the stiff, crystalline regions. Those thermoplastics with the highest Tg have stiff, bulky chains and strong, intermolecular hydrogen bonding between chains. The higher the temperature required to melt a polymer, the more ultrasonic energy required for welding. Modulus of Elasticity The crystallinity and the secondary bond strength inﬂuence the stiffness of the thermoplastic materials. The stiffness helps in the transmission of the ultrasonic energy for welding the joint interface. The stiffer a polymer, the greater the capacity to transmit ultrasonic energy.

8.13.9

Effects Caused by Thermoplastic Additives on Ultrasonic Welding

Compounding various additives or processing aids with the polymer matrix improves or creates new properties not inherent in the base polymer. These additives, which can enhance certain properties and processing capabilities of the compound material, may in some cases create problems in ultrasonic welding. Mold Release Lubricants The purpose of this additive is to provide the surface of the mold cavity a release coating that facilitates removal of the molded parts. The resin pellets are coated with an external lubricant, such as zinc stearate, aluminum stearate, ﬂuorocarbon, or silicone. This lubricant can be transferred to the joint interface and interfere with the vibrating energy. Silicone is the most detrimental in welding. If it is absolutely necessary to use an external lubricant, paintable/printable (nontransferring) resin grades should be used. These grades prevent the resin from

0

16

33

50

66

83

Temperature, (%)

Figure 8-127 Thermal behavior of amorphous vs semi-crystalline resins

100

498

8 Thermoplastic Assembly Methods wetting the surface of the mold, with no transfer to the molded part itself allowing the least amount of interference with ultrasonic welding. Detrimental mold release lubricants can in some cases be removed by using a solvent cleaner, such as ﬂuoropolymer. When small amounts of internal mold release lubricants are used in the compounding process; these lubricants are uniformly dispersed with the matrix, and usually have minimal effect on the ultrasonic welding process. Melt Flow Modifiers These additives enhance the ﬂow of the polymer melt against itself or against other materials. Examples of these lubricants include waxes, zinc, stearate, stearic acid, and esters. These melt ﬂow modiﬁers reduce intermolecular friction (melt viscosity) within the polymer or reduce melt ﬂow friction against the surfaces of the mold runner and cavities. Since molecular friction is a basis for ultrasonically induced temperature elevation, melt ﬂow modiﬁers can inhibit the ultrasonic assembly process. However, since they are compounded and dispersed internally, like internal mold release lubricants, their effect is usually minimal on the ultrasonic welding process. Plasticizers These additives are high temperature boiling organic liquids or low temperature melting solids that are added to resins to improve the ﬂexibility of the resin. Plasticizers can reduce the intermolecular attractive forces of the polymer matrix. They can also interfere with the ability of the resin to transmit ultrasonic vibratory energy. Attempting to transmit ultrasonic vibrations through a highly plasticized material (such as vinyl) is like transmitting energy through a sponge. Even though plasticizers are considered an internal additive, they do migrate to the surface over time and the combination of internal as well as surface lubricity makes it an impossibility to ultrasonically weld the plasticized material. FDA approved plasticizers do not present as much of a problem as metallic plasticizers, but a prototype of the ultrasonic weld is recommended. Impact Modifiers The impact modiﬁers used for compounding with polymers are elastomeric materials such as thermoset rubber. These impact modiﬁers can affect the ultrasonic weldability of the compound material by reducing the amount of thermoplastic available at the joint interface. They can also reduce the ability to transmit ultrasonic vibrations through the compound material, making it necessary to increase the amplitude of vibrations to generate a melt, causing a weak weld. A prototype of the ultrasonic weld is recommended. Foaming Agents The density of the foamed structure is reduced by using internal cellular voids dispersed in the polymer. The foamed structures reduce the ability to transmit ultrasonic energy. Voids in the cellular structure interrupt the energy ﬂow, reducing the amount of energy reaching the joint area, depending on the density and ﬂexibility of the polymer. A prototype of the ultrasonic weld is recommended. Flame Retardants These additives are added to a polymer to inhibit ignition or modify the burning characteristics of the compound material. The ﬂame retardant resin formulations

8.13 Ultrasonic Welding Method can adversely affect the ultrasonic welding characteristics. Flame retardant additives are generally inorganic oxides or halogenated organic chemical elements that do not permit ultrasonic welding of some ﬂame retardant parts. Typical ﬂame retardant additives are aluminum, antimony, boron, chlorine, bromine, sulfur, nitrogen, and phosphorus. The amount of ﬂame retardant additives required to meet certain burn test requirements may vary from a low percentage to 50% or more by weight of the total matrix, thus reducing the amount of available weldable material. This reduction must be compensated for by modifying the joint design to increase the amount of weldable material at the joint interface and by increasing ultrasonic energy levels for some ﬂame retardant resins. Colorants Pigments may have a detrimental inﬂuence on ultrasonic welding. Most pigments are inorganic compounds and are typically used in concentrations ranging from 0.5% to 2%. Ultrasonic welding equipment is set up to produce quality welds in natural color parts, but, when pigmented parts are ultrasonically welded, the quality of the welds are inferior. Poor quality is reﬂected in welds of lower strength and greater brittleness. The mechanism by which pigments inﬂuence welding quality has not been understood. The presence of pigments appears to inﬂuence the means of heat generation at the joint. Often, lower weld quality may be offset by welding pigmented parts for longer times than anticipated for the natural color parts. Weld times may have to be increased by 50% or more. However, these longer weld times may produce undesirable effects, such as the formation of excess weld ﬂash and marking the surface under the welding horn. Titanium dioxide (TiO2) is the main pigment used in white parts. Titanium dioxide is inorganic, chemically inert, and can act as a lubricant. If used in high loadings (greater than 5%), it can inhibit weldability. Black parts that have been pigmented with carbon black can also inhibit weldability. When designing colored injection molded thermoplastic products that need to be ultrasonically welded, welding tests of prototype parts are recommended to establish welding feasibility in this application. The use of dye coloring systems, which do not signiﬁcantly affect ultrasonic welding, may offer an alternative solution to pigments. Reinforcements and Fillers Basic polymers are compounded with ﬁllers to modify the physical properties or to reduce the cost of the resins. Typical ﬁllers are calcium carbonate, kaolin, talc, alumina trihydrate, silica, wollastonite, and mica. These ﬁllers can increase the weldability of the resin considerably; they enhance the ability of some resins to transmit ultrasonic energy by imparting higher rigidity (stiffness). However, it is very important to recognize that a direct ratio between the percentage of ﬁllers and the improvement of weldability exists only within a prescribed quantitative range. Reinforcements such as ﬁber glass, glass spheres, chopped glass, polymeric ﬁbers, and mineral ﬁbers are compounded with a basic polymer to improve the mechanical, electrical, and thermal properties of the resins. Reinforcements also improve the ultrasonic welding performance of the resins.

499

500

8 Thermoplastic Assembly Methods Resins with a reinforcement content up to 10% can be welded normally without special procedures and equipment. However, when the reinforcement content exceeds 10%, the presence of abrasive ﬁbers at the thermoplastic part surface can cause wear problems in the welding horn contact surface. In this situation, the use of hardened steel or carbide-faced (coated) titanium welding horns is recommended. When the reinforcement content approaches 35%, there may be insufﬁcient polymer at the joint interface to obtain reliable hermetic seals. When the reinforcement content exceeds 40%, tracking, or the accumulation of ﬁbers, can become so severe that insufﬁcient base polymer is present at the joint interface to form a consistent ultrasonic weld. When the ﬁber glass reinforcement used to compound the basic resin is longer than 0.125 in, the long ﬁber glass causes an injection molding problem. The long ﬁbers dispersed in the molten polymer become separated from the melt by the shear forces produced by the volume reduction of the mold runner at the gate area during the molding process. These long ﬁbers agglomerate and cluster at the mold gate. The melt ﬂow forces the long ﬁbers through the gate to ﬁll the cavity in lumps of long ﬁber glass rather than uniformly dispersing them within the polymer melt. This agglomeration of long ﬁber glass could be injected and packed in the cavity joint surface during the injection molding process, causing the molded part to have a much higher percentage of long ﬁber glass exposed at the joint surface.

Valve body

Automotive intake manifold

Cooling fans

Figure 8-128 Ultrasonic insertion thermoplastic applications

If this were to occur, no appreciable weld strength could be achieved, because the ultrasonic energy would embed itself in the adjoining surface without producing the required molten polymer to cover the joint area. This problem can be eliminated by selecting a compounded thermoplastic resin reinforced with a well dispersed short ﬁber glass (less than 35% content).

8.14

Ultrasonic Insertion

8.14.1

Applications

Ultrasonic insertion is the process of embedding a metal insert in a boss of a thermoplastic molded part. Thrust pressure and ultrasonic vibration energy drive and secure the insert into a molded hole. This assembly process replaces the difﬁcult method of injection over-molding or encapsulation around the metal insert with thermoplastic. Figure 8-128 shows several ultrasonic insertion applications. In ultrasonic insertion, a hole, slightly smaller than the metal insert, allows the installation of the metal insert into the molded thermoplastic part. This hole provides eight degrees (compounded) of interference and guides the insert into location. The metal insert is usually designed with exterior knurls, undercuts, or threads to resist torsion and pull-out loads imposed on the ﬁnished assembly. Figure 8-129 shows the ultrasonic insertion process sequence.

501

8.14 Ultrasonic Insertion

Welding horn

Thermoplastic part boss

Metal insert

Loading

Driving

Anchored

Figure 8-129 Ultrasonic insertion process sequence

8.14.2

Ultrasonic Insertion Configurations

Figure 8-130 shows two ultrasonic insertions, one by driving the metal insert into the thermoplastic hole and the other by driving the thermoplastic component over the metal insert; the latter conﬁguration it is not recommended. A variety of ultrasonic metal inserts is commercially available, all very similar in design. Thrust pressure and vibration energy on the metal insert causes the thermoplastic to melt, while the horn drives the insert into the boss. The molten polymer is displaced by the large diameters of the insert ﬂowing around the insert grooves, the melt solidiﬁes and the insert is permanently secured. Flats, notches, or axial serrations prevent pull-out and rotation in the insert. The volume of the displaced melt should be equal to or slightly more than the volume of free space corresponding to the groove and the serration of the insert. While the majority of applications are the ultrasonic insertion of threaded bore or stud inserts, other metal components can be also inserted, including eyeglass hinges, machine screws, threaded rods, metal roll pins, metal shafts, metal screens, decorative trim, electrical contacts, and terminal connectors. Ultrasonic Insertion Advantages The ultrasonic insertion process has several advantages:

Prefer configuration Welding horn

Metal insert

Plastic boss

Driving insert into part hole

No recommended Welding horn

• Elimination of possible mold damage and downtime should inserts fall into the mold • Elimination of preheating, cleaning, and hand loading of inserts • Faster injection molding cycles

Plastic boss

• Less critical insert dimensional tolerances in the mold cavity • No induced molded-in stresses around the metal insert • Multiple inserts can be driven at one time, using several horns • Ideal for automated, high production operations • Repeatability and control over the assembly process

Metal insert

Driving part on top of insert Figure 8-130 Ultrasonic insertion configurations

502

8 Thermoplastic Assembly Methods

8.14.3

8°

Insert minimum depth length, plus 0.030 inch Figure 8-131 Tapered insert for maximum pull-out

D + 0.030 in. Threaded bore Interference Undercut Lead-in

D + 0.015 in. D - 0.045 in.

D

The functional characteristics or requirements of an application determine the type and size of the metal insert and the required hole and thermoplastic boss dimensions. Approximately 0.015 in interference ﬁt between insert outside diameter and the molded hole inside diameter is required. A sufﬁcient volume of molten thermoplastic must be displaced to ﬁll the exterior undercuts, knurls, and/or threads of the insert to anchor it in place, producing the pull-out and torsional strength required for the application. Figure 8-131 shows a typical tapered threaded metal insert designed primarily for pull-out strength. It has multiple axial undercuts to provide maximum resistance. Figure 8-132 shows a metal insert with long axial knurls for maximum torque strength. For metal inserts requiring pull-out and torque strength, these inserts should exhibit a combination of both axial undercuts and knurls. The insert acceptance hole should be designed so that the insert does not bottom out before the top surface reaches the required height. It should also be of sufﬁcient depth to prevent the screw from reaching the bottom wall of the hole during assembly and driving out the insert. The hole may be tapered; especially if it is to accept a tapered metal insert, this facilitates accurate positioning of the metal insert and usually reduces the ultrasonic insertion process cycle time.

8.14.4 Figure 8-132 Long axial knurl insert for maximum torque

Ultrasonic Insertion Product Design

Ultrasonic Insertion Equipment Requirements

To ensure an efﬁcient ultrasonic insertion, following these basic rules concerning the equipment is recommended: • Because extremely high initial vibration energy is required before a melt starts, sufﬁcient power should be available. – Insert outside diameter less than 0.25 in:, 300 to 450 W are required – Insert outside diameter from 0.25 to 0.50 in:, 700 to 900 W are required – Insert outside diameter greater than 0.50 in or multiple inserts: 1200 W minimum are required • Because of the high wear situation encountered in applications in which the horn contacts the insert, the welding horn should be made of hardened steel or carbide-faced titanium for minimum wear of the horn face. For low volume applications, titanium horns with replaceable tips may be used. The horn face should be three to four times the diameter of the insert when possible, to minimize the effect of coupling of the horn to the insert. • When the horn contacts the insert, a converter protection circuit should be installed in the welder. This prevents possible damage to the converter, associated wiring, and connectors from electrical feedback caused by reverse piezoelectric effect resulting from metal-to-metal mechanical shock. • To maintain an accurate depth of insertion, the total distance the horn travels should be either mechanically limited by a positive stop or electronically limited by a lower limit switch. • A pre-trigger switch is recommended in some cases. With very small inserts (#2 or #4) pre-triggering prevents cold pressing the insert in place; for large (greater than 0.375 in) or multiple inserts, pre-triggering prevents the high starting loads that would cause overloading. • The component to be ﬁxtured should be rigidly supported directly under the insert acceptance hole.

503

8.15 Ultrasonic Stud Staking Method

8.14.5

Ultrasonic Insertion Process Guidelines

The following guidelines for ultrasonic insertion are recommended:

Screw

• Relatively low amplitude is recommended with the total gain of the horn/ booster combination being between 1.5 to 3.0. Medium thrust pressure (20–50 psi) should be used, with the pressure increased accordingly for large or multiple inserts.

Insert flash

• Slow strokes, or slow down speed of the carriage assembly should be used to prevent cold pressing the insert into the hole. • After seating, the top of the insert should be ﬂush or slightly above the surface of the part for maximum pull-out strength and torque resistance as shown in Figure 8-133. This will also prevent the possibility of a “jack-out” condition. • To prevent mechanical or electrical feedback that could damage the piezoelectric converter, the welder should be equipped with a converter protection kit. In most applications the thermoplastic component is ﬁxtured and the metal insert is driven into the hole. In special applications, where the thermoplastic component is driven over the metal insert, the previous process guidelines apply except for horn material selection (chrome-plated aluminum or titanium horns may be used), amplitude (noise level is greatly reduced), and the use of the converter protection kit is not necessary.

8.15

Ultrasonic Stud Staking Method

Ultrasonic stud staking is a variation of the ultrasonic welding shear joint design developed by Du Pont. This method is ideally suited for use with most thermoplastic resins. It is an economical and reliable assembly method that can be used to join molded thermoplastic parts at a single point or numerous locations. It is used when resin selection, size, or complexity of the molded parts preclude the use of other assembly processes. In many applications requiring permanent assembly of molded components made of similar materials, where a continuous weld is not required, the size and complexity of the molded thermoplastic parts seriously limit attachment points or weld location, the ultrasonic stud staking assembly method can be an effective and economic process. The power requirement is low because of the small weld area and the welding cycle is short, usually less than half a second. Among the applications of ultrasonic stud staking are clock frames, timers, electro-mechanical devices, electrical connectors, pump impellers, etc.

8.15.1

Ultrasonic Stud Staking Joint Design

Figure 8-134 shows the basic ultrasonic stud staking sequence: before, during, and after welding. The stud shear staking occurs along the circumference of the stud. The strength of the shear weld is a function of the stud diameter and the depth of the weld. Maximum strength in tension is achieved when the depth of the weld is equal to half the diameter of the stud. The radial interference “A” must be uniform and should generally be 0.008 to 0.012 in for studs

Spacer

Plastic part

Poor design Insert above

Screw

Plate Plastic boss

Metal insert

Recommended design Figure 8-133 Insert top surface alignment for maximum strength

504

8 Thermoplastic Assembly Methods D

B

C A Before welding

having a diameter of 0.50 in or less. Tests show that greater interference does not increase joint strength but does increase weld time. For example, studs of 0.20 in diameter with 0.016 in interference required four times the weld cycle of studs with 0.010 in interference welded to the same depth. The hole should have a sufﬁcient distance from the edge to prevent breakout with a minimum of 0.125 in recommended. In the joint, the recess can be on the end of the stud or in the mouth of the hole. When using the latter design, a small chamfer can be used for rapid alignment. To reduce stress concentration during welding and in use, an ample ﬁllet radius should be incorporated at the base of the stud. Recessing the ﬁllet below the surface serves as a ﬂash trap that allows ﬂush contact of the parts. Stud diameter = “D” Radial interference “A” = 0.008 to 0.012 in for “D” up to 0.50 in Depth of weld “B” = 0.5 × “D” for maximum strength Minimum lead-in “C” = 0.016 in (step at end of stud, or top of hole) Figure 8-135 shows variations of the basic ultrasonic shear stud staking. A third component of a dissimilar material can be locked in place. These assembly applications replace other methods, such as self tapping screws or metal rivets. Unlike metal fasteners, the ultrasonic stud staking produces a relatively stressfree assembly.

During welding

Figure 8-136 shows a variation that can be used where appearance is important, or where an uninterrupted surface is required. The stud is welded into a boss. The outside diameter of the boss should be no less than two times the stud diameter. When welding into a blind hole, it may be necessary to provide an outlet for air. Ultrasonic staking modiﬁcations are recommended, such as a center hole through the stud, or a small, narrow slot in the interior wall of the boss.

Flash trap

Weld

Flash After welding

Figure 8-134 Ultrasonic stud staking process sequence

When the amount of relative movement during welding between two parts being assembled is limited, such as when locating other internal components between the parts, a double step ultrasonic stud welding should be considered. This assembly procedure reduces the movement by 50%, while the area and strength of the weld remain the same. An example of this type of welding is shown in Figure 8-137. Figure 8-138 shows single step and double step ultrasonic stud staking variations that are very useful when welding studs into a thin wall component of 0.060 in thickness. With these process variations, there is less of a problem with trapped air. With the single step stud joint, the required lead-in (0.015 in) reduces the available area and strength of the weld; with the double step ultrasonic stud joint the weld strength is increased by 25%.

Before

After

Standard welding horns with no special tip conﬁguration are used for ultrasonic stud staking. High amplitude horns or horn and booster combinations are generally required. Best results are obtained when the welding horn contacts the part directly over the stud and on the side closest to the joint. Figure 8-139 shows an optimum ultrasonic stud staking conﬁguration, where the welding horn is located close to the stud welding interface.

Before

After

Figure 8-135 Ultrasonic stud staking variations

When ultrasonic stud staking a number of studs in a single large part, one welding horn can often be used. If the studs are widely spaced (more than 3.00 in between the furthest studs), small individual welding horns energized simultaneously should be used. In these cases, design variations of the studs, such as the rectangular plate and the square plate shown in Figure 8-140 should be considered.

505

8.15 Ultrasonic Stud Staking Method

Before

Before

After

After

0.015 0.060 0.060

Weld

0.010

Weld

Stud staking single step

Single step interface Vent hole

Before

Flash 0.030

Weld

0.016

Before

Double step interface

After

0.008

Figure 8-136 Ultrasonic stud staking in a blind hole

After

B

Stud staking with double step Figure 8-138 Ultrasonic stud staking plugin for thin-walled parts Weld

2B

Before After Welding horn closest to joint

Flash

Figure 8-137 Double step ultrasonic stud staking

Weld

Figure 8-139 Optimum ultrasonic stud staking configuration

Rectangular plate B A

A Square plate B

Stud Plate

Stud Plate

0.002/0.004 in. Assembly section “A-A”

Figure 8-140 Ultrasonic stud staking for large parts

Assembly section “B-B”

506

8 Thermoplastic Assembly Methods

8.16

Ultrasonic Stud Heading Method

Ultrasonic stud heading is an assembly procedure used to join dissimilar materials, usually dissimilar thermoplastics or metals. A hole in the metal part receives the thermoplastic stud and a special welding horn with different head proﬁles contacts the thermoplastic stud to capture or lock another material in place. Examples are printed circuit boards or metal components of an assembly in place. High frequency vibrations of the ultrasonic horn are imparted to the top of the thermoplastic stud, which melts and ﬁlls the volume of the horn cavity to produce a head. The progressive melting of the thermoplastic stud under continuous but generally light pressure melts and reforms the stud to create a locking head over the metal or thermoplastic part. Heading horn

90˚

Melt here Small area

0.5 D

L

As in any process involving localized heating by the efﬁcient dissipation of ultrasonic vibrations to control where and how fast the temperature melts the thermoplastic stud, geometry plays an important role in determining the location of high strain that results in desirable localized heating. An energy director is incorporated either in the horn for ﬂat studs, or in the pointed studs when concave proﬁle horns are used, employing the ultrasonic stud heading method. That is, the cross sectional area/height ratio of the material at the location where the initial dissipation is to occur is drastically reduced compared to the adjacent segments. The ultrasonic stud heading process requires that out-of-phase vibrations be generated between the horn and thermoplastic surfaces. Light initial contact pressure is therefore a requirement for out-of-phase vibratory activity within the limited contact area.

D

Pointed stud Heading horn

The advantages of ultrasonic stud heading include short cycle time (generally less than 1 s), tight assemblies with virtually no tendency for recovery, the ability to perform multiple stud heading with one horn, repeatability and control over the process, design simplicity, and the elimination of metal fasteners.

Small area Melt here

8.16.1 L

Thermoplastic Stud Profiles for Ultrasonic Heading

D

The integrity of an ultrasonic stud heading assembly depends on the volumetric relationship between the thermoplastic stud, horn cavity, and the ultrasonic parameters used when forming the thermoplastic stud (e.g., amplitude of the horn, weld time, pressure). Proper stud heading joint design produces optimum strength and appearance with minimum ﬂash.

Flat stud Heading horn

Melt here

Small area

L D

Hollow stud Figure 8-141 Ultrasonic heading thermoplastic stud profiles

Three basic designs are used to produce the needed geometry. The ﬁrst proﬁle is the pointed thermoplastic stud that has the energy director designed into the pointed stud itself. The pointed thermoplastic stud design yields excellent results with broad transitional materials (e.g., ABS). It provides for easier alignment of parts (looser alignment tolerances) and can be used with a wide variety of head conﬁgurations to meet special design requirements. The pointed thermoplastic stud design also reduces horn wear problems with fiber glass reinforced thermoplastics. The second proﬁle is the ﬂat thermoplastic stud that makes use of the ﬂat stud area as the energy director on the horn contacts the stud. The third proﬁle is the hollow thermoplastic stud that uses the circular pointed surface as the energy director designed in the hollow stud itself. Figure 8-141 shows these basic ultrasonic stud heading proﬁles.

507

8.16 Ultrasonic Stud Heading Method

2.25 D

2.0 D 0.5 D

D

2.0 D 0.5 D

0.5 D r.

D

D

1.5 D

Simple head profile

Flat head profile

0.2 D

Pan head profile

Figure 8-142 Ultrasonic heading stud head profiles

90˚

2D

0.5 D

0.5 D D

D

Flash head profile

Knurled head profile 2D

1.5 D

D

0.25 D

Rosette (low) head profile 0.5 D r.

0.5 D

D

Rosette (standard) head profile

0.25 D

D

Hollow head profile Figure 8-143 Ultrasonic heading stud head profiles

Figures 8-142 and 8-143 show several stud heading proﬁles using the three types of thermoplastic studs and various horn cavity designs that have been developed for the ultrasonic stud heading process. The requirements of the application and the physical size of the stud(s) being headed determine the design to be utilized. The principle of ultrasonic stud heading is the same for each: the initial contact between the welding horn and the thermoplastic stud should be kept to a minimum, thus concentrating the energy to produce a rapid melt. Simple Head, Flat Head, and Pan Head Profiles Simple head, ﬂat head, and pan head proﬁles are recommended for pointed thermoplastic studs with a diameter of 0.060 in or less. The top of the pointed stud should be tapered (cone-shaped); it is the point (energy director) that initiates material melting, reducing energy being transmitted through the stud. The alignment between the welding horn and the thermoplastic stud is not as critical as with the rosette proﬁles and the appearances of the formed heads are based on the proﬁles used. These welding horn tip proﬁles are less susceptible

508

8 Thermoplastic Assembly Methods to wear than the rosette proﬁle tips when abrasive materials are being stud headed. Flush Head Profile For applications requiring a ﬂush surface and having sufﬁcient thickness in the contained countersunk recess, the ﬂush head proﬁle is ideal. The pointed thermoplastic stud design used for the simple head is recommended and a ﬂat faced welding horn or tip is utilized. The welding horn for the ﬂush head has a smooth face. The ﬂush head proﬁle may be used for all thermoplastic materials. Knurled Head Profile The knurled head proﬁle available in both male and female patterns is designed for simplicity and rapid rate of assembly. The knurl is available in ﬁne, medium, and coarse patterns. It may be used for all thermoplastics for applications where appearance is not critical. There is no dimensioned horn cavity and multiple heads may be made without concern for precise horn alignment or stud diameter. A hand-held ultrasonic welding gun may be utilized. Rosette Head Profiles Rosette head proﬁles are most commonly used for ﬂat thermoplastic studs having a diameter between 0.060 and 0.187 in. The top of the stud is ﬂat, and melt is initiated by the point (energy director) in the welding horn cavity. The head produced is twice the diameter of the stud and satisﬁes the requirements of the majority of ultrasonic stud heading applications. It is ideal for stud heading of non-abrasive thermoplastics, both rigid and ﬂexible. The rosette (low) proﬁle and the rosette (standard) proﬁle welding horn tips are available for studs with diameters of 0.030 and 0.187 in. The rosette style is used primarily in press operations. Hollow Head Profile Hollow thermoplastic studs are generally used when the studs are greater than 0.187 in in diameter. Hollow thermoplastic studs offer advantages in molding, because they prevent surface sinks and internal voids. Heading a hollow thermoplastic stud produces a large, strong head without having to melt a large amount of material. Hollow studs can also be used with hand-held ultrasonic spot welders, depending on material and size of application. Also, where disassembly for repairs is a primary requirement of the application, repairs can be made by breaking away the formed stud head for access and driving a self-tapping screw into the inside diameter of the stud for reassembly. The hollow stud reduces the material volume of the stud and the cycle time is shortened because there is less material to melt and solidify. Multiple Ultrasonic Stud Heading More than one stud may be headed in a single operation. The welding horn used for multiple stud heading can be half or full wavelength in design. If the studs are on the same plane and within 0.50 in of each other, a half wave horn is recommended. Large parts having thermoplastic studs widely spaced on the same plane would require a full wave composite welding horn to provide the necessary amplitude for stud heading. Table 8-14 provides dimensions for different stud heads and proﬁles.

509

8.17 Ultrasonic Spot Welding Method Table 8-14 Profiles and Dimensions of Stud Heads

Head form proﬁle

Stud type

Stud dia.

Head dia.

Head height

Stud height

Simple

Pointed

D

2.25 D

0.5 D

1.6 D

Flat head

Pointed

D

2.00 D

0.5 D

1.85 D

Pan head

Pointed

D

2.00 D

0.5 D

1.7 D

Flash

Pointed

D

1.50 D

–

–

Knurled

Flat

D

2.00 D

0.5 D

1.9 D

Rosette (low)

Flat

D

1.50 D

0.25 D

0.6 D

Rosette (stand.)

Flat

D

2.00 D

0.5 D

1.65 D

Hollow

Hollow

D

1.50 D

0.25 D

0.5 D

8.17

Ultrasonic Spot Welding Method

Ultrasonic spot welding is an assembly method for joining two thermoplastic components at localized points without a preformed hole or the need of an energy director in the joint interface. Ultrasonic spot welding produces a strong, structural weld and lends itself to large thermoplastic parts, sheets of extruded or cast thermoplastic, complicated geometry parts, and hard to reach joint surfaces. Vibrating ultrasonically, the pilot tip of the spot welding horn passes through the top section. The molten thermoplastic displaced is shaped by the horn cavity in the tip and forms a neat raised ring on the surface. Simultaneously, energy is released at the interface producing essential frictional heat. As penetration of the bottom section is made, displaced molten thermoplastic ﬂows between the two surfaces into the surrounding interface area and forms a permanent molecular bond. The spot welding process is recommended for large thermoplastic parts or sheets applications, by using an ultrasonic hand-held gun and power supply. The standard horn tip produces a head having a diameter up to three times the wall thickness of the top thermoplastic part. The length of the ultrasonic spot welding horn tip is 1.50 times the wall thickness of the top layer. Figure 8-144 shows an ultrasonic spot welding joint proﬁle of two thermoplastic components.

T x 3.00 T x 1.50

Welding horn

Blind taper hole Raised ring top layer

Molecular bond joint interface

Horn cavity Pilot tip

Top plastic wall Base plastic structure

T Unmarked surface

Before welding

Figure 8-144 Ultrasonic spot welding joint profile

After welding

Displaced base layer

510

8 Thermoplastic Assembly Methods Ultrasonic spot welding produces a clean and strong structural weld to assemble two thermoplastic components made of the same polymer. The base layer bottom side has a cosmetic unmarked surface and the outer side of the top layer has a neat, raised ring and a blind taper hole. The process is fast, requires no extra fasteners, and generally no special ﬁxturing. Most thermoplastic materials can be ultrasonically spot welded.

8.17.1

Hand-Held Ultrasonic Spot Welder

The hand held ultrasonic spot welders are compact, portable, and lightweight tools for ultrasonic assembly of large parts and those with hard to reach joint areas. With the hand-held ultrasonic spot welder, pressure is controlled manually and a trigger switch located in the grip permits the operator to control the duration of applied ultrasonic energy for spot welding.

Hand held welding horn Power supply Welding gun

Ultrasonic spot welder

Figure 8-145 Hand-held ultrasonic spot welder

On the hand-held ultrasonic spot welders, the operator grips a sliding, spring loaded sleeve and by applying pressure against the top surface of the two parts to be welded, starts the ultrasonic spot welding cycle. The weld time is controlled by the power supply. An automatic switch energizes the power supply whenever the ultrasonic gun is pressed against the top surface of the two thermoplastic components being spot welded. By lifting the ultrasonic gun away from the thermoplastic components, the switch will override the weld time setting, causing the power supply to de-energize. A screw adjustment varies the depth of the stroke required before ultrasonic triggering occurs. The cavity that forms a ring at the base of the ultrasonic welding horn activates a switch, causing the ultrasonic gun’s vibrating energy to ﬁnish the spot welding operation. The simplicity of the spot welder design enables operators to be trained in minutes to perform high volume, high integrity ultrasonic spot welding. Figure 8-145 shows two types of hand-held ultrasonic spot welder guns and a portable power supply.

511

9

Thermoplastic Effects on Product Design

9.1

Polymer Melt Behavior

When an injection molded thermoplastic component has been produced, the conversion process subjects the thermoplastic material to severe physical conditions involving elevated temperatures, high pressures, and high shear rate ﬂow, as well as chemical changes. These drastic processing conditions place some limitations on the design of thermoplastic parts and on their performance. The walls and other sections of a thermoplastic part represent to the designer the things required to make the part functional for its intended use. To the mold designer and molding operator, they represent the ﬂow path for the thermoplastic material. This ﬂow moves rapidly under the complicating condition of a passage much cooler than the resin melt, through an oriﬁce (the gate) whose dimensions are severely restricted to ensure acceptable appearance of the part. With these complications in mind, it may not be possible, or it may be very difﬁcult, to mold some shapes. Large area parts with thin walls represent one class of parts that present difﬁculties. Because of the heat exchange between the ﬂowing resin melt and the mold cavity surfaces, the melt ﬂow may freeze before the part is completely ﬁlled. Parts that have alternate sections of thick and thin walls cause problems in ﬂow and cooling that make it difﬁcult to ﬁll the cavity. In some cases, the resins that have been selected for the end use requirement are too viscous to ﬂow properly in a part and this makes manufacturing difﬁcult. Figure 9-1 shows the classic Newtonian characteristics of thermoplastic melt ﬂow. The melt ﬂow is laminar with the maximum velocity in the middle of the melt channel and with zero velocity on the melt channel surface. This external stationary surface becomes an insulation skin that is approximately 0.040 in thick.

Shear force between layer "A" & "B" Finite velocity layer "A"

Maximum velocity

Melt front Melt channel Zero velocity skin layer "B"

Figure 9-1 Newtonian laminar polymer melt ﬂow behavior

Flow restriction problems do not only affect ﬁlling the runners and mold cavities with thermoplastic melt. There are a number of other consequences of melt ﬂow limitations in molding that are less obvious and more signiﬁcant to the performance of the part. Figure 9-2 shows dimensional changes after the mold cooling stage, caused by the orientation of the melt ﬂow. Restricted melt ﬂow causes high shear rates in the polymer as it ﬁlls the mold cavity. Therefore, higher injection pressures, the use of higher melt temperatures, and higher melt ﬂow speed are required to ﬁll the part. Higher melt ﬂow speeds generally have lower impact resistance and lower strength properties. The use of higher temperatures usually results in degradation of the thermoplastic material’s properties by shearing off the polymer to lower molecular weight materials. The high shear rates encountered in molding also result in degradation of the molecular weight of the material just from the shear. The shear rate is directly dependent on the pressure drop in the melt channel, which is a cubic function of the melt channel diameter. The high shear rate produces two other effects that signiﬁcantly affect part performance. The thermoplastics’ molecules become aligned as a result of the high shear ﬂow so that the material in the walls is highly oriented in the ﬂow direction. This may be a desirable effect. For example, a polypropylene cover is molded to generate a living hinge effect by orienting the material. In some materials, such as nylon, the single direction orientation results in improved

Flow orientation

Edge gate

Effects only of flow orientation without shrinkage after cooling

Edge gate

Figure 9-2 Effects of ﬂow orientation on dimensions

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9 Thermoplastic Effects on Product Design strength in the ﬂow and perpendicular to the ﬂow direction. In most cases, the effect is undesirable, because the strength in the perpendicular to the ﬂow direction is reduced and the part has a tendency to crack along a ﬂow line. In addition, oriented materials exhibit reduced end-use temperature properties and cause warpage in the molded part. Molded-in stress is another condition that develops in injection molded parts, particularly in thin-walled parts. The thermoplastic melt ﬂowing along the cavity walls of the mold is chilled by heat transferring to the cold mold cavity walls so that a thin layer is essentially frozen at the mold walls. The melt between the two chilled skins continues to ﬂow and as a result, it will shear the chilled skins of thermoplastics and subject them to molded-in stresses. When the ﬂow ceases, the surfaces (skins) of the part are in tension and the inside material is in compression resulting in a molded-in stress condition. This stress level is added to any externally applied load so that a part with molded-in stress is likely to fail at reduced load levels. There are other conditions that result from the molded-in stresses. In materials such as polystyrene, which have low elongation rates and are in the glassy state at room temperature, a frequent result is crazing, the appearance of many ﬁne micro cracks across the surface in a direction perpendicular to the molded-in stress direction. This result may not appear at the beginning and may be triggered by exposure to either a mildly solvent liquid or vapor. Polystyrene parts dipped in kerosene will craze quickly in areas with molded-in stresses. The crazing effect always leads to premature part failure. A post-molding annealing operation may minimize these molded-in stress problems. Molded-in stresses are real loads applied to the materials, and when even slightly elevated temperatures are applied to the part, the molded-in stresses will cause the part to deform severely. This may affect an impact grade of material or a semi-crystalline polymer even more drastically than a glassy material. The molded-in stress problems are interactions between the molding process and the product design. Poor control in the molding process can produce severe orientation and frozen-in stress in parts that are properly designed. On the other hand, there are product designs used particularly with certain materials where it is impossible to avoid the frozen-in stress and orientation problems. It should be pointed out that in the development of difﬁcult to mold parts, the condition can be observed in transparent parts by means of the photo stress effect. Examination of a transparent amorphous molded part by polarized light will show the combined effect of the orientation and the molded-in stresses. It is difﬁcult to determine which effect is being observed, because both have the same birefringence effect on polarized light. The use of reﬂected polarized light from the surface of the part gives a somewhat different reading of the effect. It is desirable to differentiate between the two effects, because the resulting part performance will be different for each condition. To avoid these molded-in stress problems in new product development, it is recommended to make a prototype mold and perform a “melt ﬂow” stress analysis on the new product to simulate the stress formation during the injection molding process. Physical testing of the product, melt ﬂow stress analysis, photo stress analysis, or exposure to selected solvents to check for stress crack characteristics may initiate changes in product design or molding process conditions to minimize molded-on effects and improve product performance.

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9.2 General Characteristics of Polymers

9.1.1

Thermoplastics Glass Transition Temperature

The viscoelastic behavior of thermoplastics is modiﬁed by material-dependent factors. These factors indicate the chemical, physical, and/or compositional nature of the thermoplastic itself. It is important to be aware of the manner in which these factors affect the viscoelasticity of a thermoplastic. Therefore, it is helpful to categorize polymers into thermoplastic versus thermosetting, amorphous versus semi-crystalline, and reinforced versus unﬁlled polymers. The polymer molecules within thermoplastic materials are made up of many thousands of smaller molecules, linked together by strong bonds. In the case of thermoplastic materials, the individual polymer molecules are not interconnected and these macro-molecules are potentially capable of being quite mobile within the structure. The thermosetting materials consist of macro-molecules that are extensively interconnected to form a network system of similar linked molecules.

9.2

General Characteristics of Polymers

Descriptions of numerous material properties that should be considered will be covered in this section and precede the discussion of thermoplastic and thermosetting process effects on the characteristics of the product. The toughness of polymers, or impact resistance, varies with molecular structure and the type of stress application. Care must be taken in relating ﬂexibility to toughness, but generally, a more rubbery character gives higher elongation to break and better impact resistance values, although such materials would have lower modulus of elasticity. Some stiffness can be recovered by adding ﬁbrous reinforcement, but the reinforcement can also affect impact resistance. The relationship between toughness and stiffness must also be considered during material selection. A longer, higher molecular weight thermoplastic polymer will be tougher than a shorter and lower molecular weight polymer of the same chemical structure. The higher molecular weight polymer will also

Semi-crystalline polymer

Flexural modulus

The mobility of polymer chains within thermoplastic resins is additionally affected by the internal arrangement of the linear macro-molecules within the bulk thermoplastic. In their solid state, thermoplastics can be considered to behave either as glassy solids or as crystalline solids. The glassy solid state is reached when the polymeric melt has been cooled below a certain mold temperature, known as the glass transition temperature (Tg). The Tg value and the end-use temperature depend on the type of thermoplastic. When the end-use temperature of the thermoplastic is below its Tg, the molecules of the polymer do not move or display any ﬂuid-like characteristics. When the Tg of the thermoplastic increases, only the crystalline micro-structure responds to the ﬂuid-like behavior. The crystallinity in the polymer will exist as long as the end-use temperature remains below the melting point (Tm). The crystallites act as effective crosslinks between the polymer molecules, thereby limiting their mobility at temperatures above Tg. The viscoelastic nature of a thermoplastic above its Tg usually requires careful consideration, as does the typical creep behavior of an amorphous thermoplastic (below its Tg) compared to a semi-crystalline thermoplastic (above its Tg). Figure 9-3 illustrates the thermal differences between an amorphous and a semi-crystal line thermoplastic.

Tg Tg Amorphous polymer Tm

Temperature

Figure 9-3 Comparison of thermal characteristics of polymers

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9 Thermoplastic Effects on Product Design provide higher melt viscosity or lower melt ﬂow rate and will be more difﬁcult to injection mold. Such high melt viscosity polymers have an advantage in such processes as extrusion, blow molding, and vacuum forming. The toughness of a polymer can be increased by compounding with impact modiﬁers to form a second noncontinuous phase used to disrupt crack propagation. Butadiene, EPDM, acrylic, and silicone are some of the impact modiﬁers used in toughened PP, nylon, polystyrene, and ABS. Some high molecular weight thermoplastics exhibit natural toughness because of their entangled molecular structures. The properties of these materials are reduced when annealed or by secondary operations such as ultrasonic welding, which apparently relaxes the molecules so that they lose their impact resistance. In thermosetting resins, which are cross-linked in their ﬁnal form, the molecule is inﬁnitely large. During processing, however, its size can be as small as that of the monomer. Some of the liquid monomers used to produce thermosets can have very low viscosities and can be ideal for processes such as transfer molding and pultrusion, in which the compound must ﬂow through preplaced glass reinforcements. For other processes, such as extra high strength molding compound and pre-oriented compression molding, the compound must be thickened by reaction to stay on the glass reinforcements. With sheet molding and bulk molding, the compound or specialty formulations must be chemically thickened so that it can carry ﬁllers and glass reinforcement during process ﬂow. Mold shrinkage effects that occur during the processing of both thermoplastic and thermosetting resins are discussed later in this chapter. Reinforcements and ﬁllers also increase the ﬂexural modulus, but in most cases they also signiﬁcantly reduce impact resistance. The geometry of these ingredients is very important; the ﬂake-like or ﬁbrous materials have greater stiffening effects and usually worsen impact resistance. They can also orient in the melt ﬂow direction, giving anisotropic physical properties, which can lead to poor impact resistance in the cross ﬂow direction. Rounded (particulate) ﬁllers tend to have less signiﬁcant effects. Some round glass reinforcements have been shown to increase both stiffness and impact resistance in many polymers. The use of coupling agents to reinforce and increase the adhesion of the polymer to the reinforcement particles can improve impact resistance in comparison to reinforcements without coupling agent.

9.2.1

Critical Properties of Thermoplastics

Certain thermal and related properties of the polymers are critical in several application areas, including ﬁlm formation and packaging, health care products, production of precision injection molded articles, optical devices, ﬁber industry, automotive and transportation industries, construction industry, electronics and electrical industries. • The effect of shearing forces on melt viscosity and molecular chain during melt processing causes degradation of the polymer. • The heat capacity and thermal conductivity of solid and liquid polymers control the processing parameters to heat-up and cool-off characteristics in injection molding and extrusion.

9.3 Polymer Reinforcements • The crystallization rate of the polymer, when cooling the melt from a liquid state to a solid shape at room temperature, directly affects the cycle time and other process parameters in the injection molding process. The crystallization rate upon heating the polymer and a fast freeze-off of the melt affects the molded-in stresses and the product dimensional control. • Nucleating agents affect the crystallization kinetics of the polymer and control the processing cycle and the product dimensional control. • The ﬁllers and reinforcements used to compound the polymers improve the thermal properties, such as UL temperature index, Tg, crystallization rate, heat-deﬂection temperature, thermal conductivity, mechanical properties, the product dimensional control, and the processing cycles. • The thermal effects of the polymers when chemical additives are used to improve other properties of the compounded polymers. • Desirable and undesirable effects of inorganic ﬁllers and plasticizers on polymer ﬂame resistance properties. • The effect of thermal properties of individual components of polymer blends on phase separation or degree of phase mixing and hence on the ultimate end-use performance. • Changes in tensile or ﬂexural strength as a function of temperature and the characteristics these compounds have on the end-use temperature and the product dimensional control or the temperature required for secondary operation treatments, such as annealing or heat sterilization. • The tendency of a polymer to creep at different temperatures, which is pertinent to post-molding operations, end-use temperature, storage conditions, and aging of the compound.

9.3

Polymer Reinforcements

Thermoplastics are often compounded with various reinforcements and ﬁlled to improve thermal and mechanical properties. The viscoelastic properties of a thermoplastic can be modiﬁed by compounding ﬁbrous reinforcements with the base polymer. If there is good inter-facial adhesion between the polymer and the ﬁber, the ﬁber will impart its non-creep behavior to the polymer matrix. Glass, carbon, boron, mineral, and certain organic ﬁbers, such as aramid, are among the more commonly used ﬁbers. Relatively short ﬁbers are usually compounded with thermoplastics, whereas very long ﬁbers may be incorporated into thermosetting polymers. Reinforcements affect the viscoelastic properties of semi-crystalline thermoplastics above their Tg. Another material factor that has to be considered is whether the material is homogeneous and/or isotropic. The presence of isotropic viscoelastic properties implies that there will be a variation in these properties depending on the actual melt ﬂow direction. The addition of ﬁber reinforcements to the polymers and the conﬁgurations of the long chain polymer molecules leads to property improvements of the injection molded thermoplastic part. The property improvements are obtained as a function of the melt ﬂow direction of the ﬁbers. The isotropy characteristic is caused by ﬁbers and/or polymer molecular chains that may become preferentially aligned and oriented during the injection molding process.

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9.3.1

Types of Fiber Reinforcements

The addition of glass, carbon, inorganic, or high tensile organic ﬁbers to compound a polymer will affect its physical properties. These properties can vary from being similar to those of the base polymer, at low loadings, to approaching those of the reinforcement, at high loading conditions. The diameter and length of the ﬁber glass are very important; they affect the ﬁnal physical properties of the compound. They can be very short, such as milled glass ﬁbers, which is less than 0.020 in long. Short cut ﬁber glass is about 0.125 in long; long cut ﬁber glass is about 0.25 to 0.50 in long; or it can be continuous ﬁber glass for composite lay-out applications. Glass and mineral reinforcement can also be used in ﬂake form. Mica is one natural form of mineral ﬂake. We will discuss primarily the use of ﬁber glass as a polymer reinforcement, although other materials are also used to improve the physical properties of the polymers, such as talc, calcium carbonate, wollastonite and others. Adding glass increases the stiffness: the higher the glass content, the higher the ﬂexural modulus. This effect is true no matter what form of glass is used. The advantage of using long glass ﬁbers is that with higher glass loadings, some physical properties, such as tensile and ﬂexural strength, become more related to those of the reinforcement. However, the thermoplastics injection molding process does not allow the long ﬁber glass to ﬂow uniformly. The small gate opening shears off (agglomerates) the long ﬁber glass at the gate and separates it from the melt. The long ﬁber glass retained by the gate is eventually injected inside the cavity, causing the molded part to have poor strength distribution and poor surface ﬁnishing. Therefore, long ﬁber glass is not recommended for the thermoplastics injection molding process. When thermosetting compounds are loaded with more than 60% glass, the resin becomes simply a binder to hold the glass together. At lower glass contents and with short glass lengths, physical properties such as tensile and ﬂexural strengths still relate largely to the base polymer. With shorter glass lengths, the impact resistance of the product varies inversely with stiffness. The use of polyurethane reaction injection molding for large automotive panels has always been a problem, because adding milled or ﬂaked glass to increase stiffness becomes effective only at about 12% loading, exactly the glass ratio at which the impact resistance property decreases dramatically. With long glass ﬁbers, an apparent increase in impact resistance can be measured. This depends to some extent on the measurement system, but incorporation of long, chopped or continuous glass into thermosetting polymers can result in products with very high impact resistance. This occurs only if the elongation to break of the polymer matrix is high enough for the forces to be transferred to the reinforcement. With most amorphous thermoplastics, the ﬂexural modulus stays relatively constant up to the region of the glass transition temperature or softening point, so that the material is useful up to that point. The addition of glass increases stiffness, but at the region of the polymer softening point, the properties of the material drop dramatically whether the amorphous thermoplastic is reinforced or not.

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9.3 Polymer Reinforcements On the other hand, semi-crystalline thermoplastics compounded with short ﬁber glass tend to decrease slowly in stiffness with temperature and tend to be very susceptible to creep under load. Unreinforced semi-crystalline thermoplastics, therefore, have a useful range ending considerably below their melting points. Reinforcement has a dramatic effect in that it increases the overall stiffness and creep resistance right up to the melting point. Illustrating this characteristic is the fact that the heat deﬂection temperature of unreinforced polypropylene is 60 °C (140 °F), while that of ﬁber glass reinforced polypropylene is 150 °C (300 °F). Reinforcement is therefore much more useful with semi-crystalline thermoplastics than with amorphous thermoplastics.

9.3.2

Isotropic Warpage of Fiber Reinforced Resins

Many processes that rely on the polymer melt to carry the reinforcement as it ﬂows into the mold will result in oriented reinforcing ﬁbers. This can cause part distortion and weakness in the transverse direction of the ﬁber glass. Figure 9-4 shows a thin-walled round disk molded with a single edge gate; the left illustration shows the ﬁber glass orientation in both directions, the right illustration shows the warpage of the molded disk caused by ﬁber glass isotropic shrinkage. In structural components requiring maximum strength from the ﬁber glass reinforcement, the ﬁbers need to be oriented in speciﬁc directions in the mold cavity. This is difﬁcult to achieve with processes that require ﬂow of the polymer melt and reinforcement orientation to ﬁll the cavity. For thermoplastic injection molding, the ﬁber glass must be oriented by the gates. For thermoset polymers, the ﬁber glass is oriented by using other techniques, such as ﬁlament winding or hand lay up in the transfer molding process.

9.3.3

Fiber Glass Reinforcement Limitations

The ability of various processes to control the orientation of the ﬁber glass reinforcement is very important to avoid process defects on the surface ﬁnish. Poor dimensional stability, low mechanical properties at elevated temperatures, and reduced impact resistance are the effects caused by improper ﬁber glass orientation.

9.3.4

Injection Molding Process Effects on Fiber Glass Orientation

Fiber glass reinforced thermoplastic molecules aligned in the melt ﬂow direction exhibit non-Newtonian behavior. The melt ﬂow orientation effects in thermoplastics injection molding are much greater than in thermoset processes, such as reaction injection molding. The thermoplastic melt structures are oriented when the melt solidiﬁes during the cooling cycle and the ﬁber glass orientation affects the properties of the product. The tensile strength in the melt ﬂow direction is much greater than in the transverse cross direction. Figure 9-5 shows the stress-strain curves of ﬁber glass reinforced thermoplastic polymers based on the ﬁber glass orientation obtained in the injection molding process. Structural molecules and ﬁbrous reinforcement line up in the direction of melt ﬂow when forced through an oriﬁce. The polymer melt is stretched from higher

Transverse direction

Edge gate

Flow direction Thin wall thickness

Thin walled disk isotropic warpage

Figure 9-4 Isotropic warpage behavior of glass reinforced resins

518

Stress

0°

-4

5°

0°

0°

-9 0°

-9 0°

9 Thermoplastic Effects on Product Design

45

°

90°

Strain 90° Transverse direction

45°

Flow direction

0° Glass orientation

Bi-directional

Flow direction

Random

Figure 9-5 Stress-strain curve caused by ﬁber glass orientation

pressure to lower pressure. Such effects occur in the plastiﬁcation unit, nozzle, mold, runner and gates, and thin long cavity walls. As the thermoplastic melt goes through a tubular nozzle, it is also possible to orient the molecules and reinforcements perpendicular to the ﬂow by constricting and then expanding the tubular melt channel inside the runner and gate while keeping the cavity wall thickness constant. The orientation of the ﬁber glass and of the polymer molecules is determined by the type, size, location, and number of gates used. When melt ﬂow into the cavity is basically in one direction, it will orient in that direction. This takes place when a rectangular plaque mold is ﬁlled using a ﬁlm type of gate located at the middle of the short side. However, if the melt is injected at the center of a disk through a sprue gate, a diaphragm gate, or pin point gate, the orientation of the polymer molecules and the ﬁber reinforcement will change. As the melt ﬂow front expands, there is some tendency for orientation across the ﬂow direction. In some cases, the orientation at the surface will be different from that at the core, where the ﬂow becomes more constrained as the mold cavity ﬁlls. Mold temperature, part wall thickness, polymer crystallinity rate, ﬁber glass reinforcement length, and gate design all affect the orientation and thus the mold shrinkage and the molded-in stresses. The more restricted the ﬂow, the greater the direction of melt ﬂow orientation. Thin-walled parts are more oriented than thick-walled parts. An example of orientation is found in the PET ﬁlm manufacturing process; the cast ﬁlm is gradually stretched 400 to 500% in the forward direction and 200% in the perpendicular direction. The ﬁlm becomes very strong in the forward direction, while the mechanical properties are reduced in the perpendicular direction. The resulting product is widely used in magnetic tapes and electrical insulation applications. Similar stretching is used in the production of high tensile polyamides and polyester ﬁbers to reinforce tires.

9.3.5

Tensile Stress Effects Caused by Fiber Glass Orientation

Fiber glass reinforced thermoplastics exhibit a linear relationship between stress and strain up to the point of elongation and failure at high strain. Typically, a ﬁber glass reinforced commodity thermoplastic will break at 350,000 psi at a strain of 3.50%. In contrast, the same unreinforced polymer and a relative majority of matrix materials are non-linear, having a low tensile strength (typically in the range 5,800 psi to 13,000 psi) and high strain rates (10% to 20%). The major difference between ﬁber glass reinforced thermoplastics and other more conventional unreinforced materials is that their properties can be isotropic. The direction of orientation of the ﬁber glass varies from unilaterally aligned ﬁbers, through bi-directional transverse ﬁbers, to completely random ﬁbers. The tensile behavior of such compounds depends on the direction of the applied stress and the melt ﬂow orientation of the polymer as illustrated in Figure 9-5. When the unidirectional ﬁbers are tested with the stress aligned with the ﬁbers (0°), maximum strength and ﬂexural modulus are achieved. When the stress is applied at 45° to the ﬁber orientation, the contribution from the reinforcement will be reduced to the resolved component of stress, with the result that strength and ﬂexural modulus suffer accordingly. When the stress is applied perpendicular to the ﬁbers, there is no enhancement of matrix properties and indeed the strength of the matrix may be reduced due to the presence of the ﬁber. The properties of randomly reinforced thermoplastics are approximately the same in all directions.

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9.3 Polymer Reinforcements Edge gate 15.000 psi (TD strength) 16.000 psi (TD strength) 13.000 psi (TD strength) 27.000 psi (MFD strength)

5. 00

in

0.125 in

3.00 in

13.600 psi (TD strength)

Figure 9-6 Tensile stress variations of PET 30% ﬁber glass reinforced

Figure 9-6 shows the tensile strength variations found in a PET 30% ﬁber glass reinforced molded plaque, based on ﬁber glass orientation and the area location tested.

9.3.6

Flexural Modulus Effects Caused by Fiber Glass Orientation

In many practical applications, the ﬂexural modulus (stiffness) of a thermoplastic material is just as important as its tensile strength. For tension, the stiffness per unit length is given by the product E × A, where E is the ﬂexural modulus, or modulus of elasticity, and A is the cross sectional area. However, stiffness is probably much more important in situations where the material is subjected to ﬂexure. In such cases, the stiffness per unit length is a function of the product E × I, where I is the moment of inertia of the area. Therefore, the ﬂexural modulus of the material increases proportionally to the amount of ﬁber glass reinforcement and the ﬁber glass orientation. However, the injection molded thermoplastic part stiffness can also be improved by many orders of magnitude by increasing the moment of inertia of the cross sectional area. Because the moment of inertia is proportional to the third power of the wall thickness, if the wall thickness is doubled, the stiffness is increased by a factor of eight. The drawback of this is that doubling the wall thickness will increase the weight of the molded product, increase the molding cycle, create possible molding process problems, and increase manufacturing costs of the component. To increase the moment of inertia of the cross sectional area without increasing the weight of an injection molded thermoplastic part requires the application of a new injection molding technology, known as double injection molding, which uses two separated thermoplastic materials. When the molded product is subjected to bending, the maximum ﬂexural stresses are at the outer surface, so it is sensible only to have the high ﬂexural strength thermoplastic material at the skin and use lower strength thermoplastic material at the center. This produces a sandwich-type composite, in which the ﬁber glass reinforced thermoplastic material forms the skin while the central region (core) consists of thermoplastic foam or a honeycomb structure. The most important property of the core is its

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9 Thermoplastic Effects on Product Design high shear strength-to-density ratio. Honeycomb structures are about ﬁve times better than foams in this respect, but it is also more expensive to encapsulate the honeycomb structure insert. In addition, this molding process requires the use of a special, higher cost dedicated molding machine and a complicated, more expensive mold.

9.4

Chemical and Environmental Resistance

The chemical resistance of a polymer, like other types of degradation resistance, is the ability to withstand attack by chemicals over time without a change in its surface appearance, part dimensions, weight, mechanical and electrical properties. Thermoplastics that differ in their solubility values will swell or dissolve, depending on the relative differences in solubility parameters. Polymers may also stress crack by absorbing different solvents. Solubility and stress cracking are inverse functions of molecular weight, chain branching, and cross linking. Thermoplastics consist of linear or branched polymer chains, which may be reversibly dissolved and solidiﬁed. Thermosets are cross-linked polymers, which will swell but not dissolve in solvents. Segmental movement at temperatures above the glass transition temperature (Tg) favors diffusion. Therefore, the rate of absorption of a corrosive molecule is dependent on the dimensional rate of change in the polymer chains. The rate of chemical attack is controlled by the morphology of the polymers. The rate of chemical attack is lower at temperatures below the Tg in crystalline, cross-linked, and rigid polymers. Ozone will attack unsaturated polymers, such as natural rubber, and will produce hydrolyzable ozonides. Ozone will also attack the surface of polyoleﬁns and form a barrier to further attack the polymers. When the volume of the environmental ﬂuid is ﬁnite, diffusion of a chemical into the polymer will proceed until equilibrium is attained. Water will diffuse into polyesters and will hydrolyze the ester groups. Considerable general information on the resistance of thermoplastics to environmental attack can be obtained from studying their structure. Aliphatic hydrocarbon polymers, such as HDPE and n-octane, will be resistant to attack by water, aqueous salt solutions, aqueous alkaline solutions, polar solvents such as ethanol, and non-oxidizing acids such as hydrochloric acid. Oxidizing acids, such as nitric or chromic acid, may attack HDPE and the rate of attack will increase with temperature. Crystalline regions will be more resistant to attack than amorphous regions of the polymer and the presence of low molecular weight polymers will increase the tendency toward stress cracking in the presence of solvents. Hydrocarbon polymers with branches, such as LDPE, or polypropylene, will be more susceptible to oxidative attack than HDPE. Cross-linked thermoset polymers will be more resistant to environmental attack than thermoplastics. Because polar pendant groups, such as chlorine, increase their solubility parameters, polymers such as PVC will be more resistant to aliphatic solvents than polyethylene. Resistance to solvents, heat, and corrosives will increase when ﬂuorine pendant groups are present and the presence of multiple ﬂuorine groups, as in polytetraﬂuoroethylene (PTFE), will increase the resistance dramatically.

9.4 Chemical and Environmental Resistance Unsaturated hydrocarbon polymers, such as polybutadiene, will be attacked by chloride, hydrogen chloride, hydrogen ﬂuoride, oxygen, and ozone. Polymers with hydroxyl groups, for example, polyvinyl alcohol, cellulose, incompletely substituted cellulosic such as secondary cellulose acetate and epoxy polymers, are susceptible to attack by oxidizing agents, such as nitric acid. Because the hydroxyl group in phenol is acidic, phenolic resins are not resistant to alkaline media. It is important to note that attack on pendant groups does not decrease the molecular weight of a polymer and may have only a minor effect on the integrity of the polymer. Therefore, the acid hydrolysis of polybutyl acrylate produces polyacrylic acid, but does not decrease the molecular weight of the polymer substantially. Because of the presence of the methyl group on the same carbon atom as the ester group in polymethyl methacrylate, the rate of hydrolysis of this ester is slower than that of polybutyl acrylate. Pendant amide and cyano groups are also hydrolyzed without affecting the molecular weight of polymers, such as polyacrylamide and polyacrylonitrile. In contrast, the molecular weight is decreased and the integrity of the polymer destroyed when ester, amide, or urethane groups are part of the polymer backbone. Therefore, the molecular weight of polyamide, such as nylon, polyethylene adipate, and polyurethane are decreased when these groups are hydrolyzed by acids. However, aromatic esters, such as polyethylene terephthalate (PET), are more resistant to hydrolysis than the aliphatic esters. Ether groups, such as those present in polyvinyl butyl ether and polyphenylene oxide, are resistant to attacks by corrosives. However, when the oxygen linkage is an acetal linkage, as in cellulose, the oxygen-carbon linkage is readily cleaved by mineral acids. Polymers with stiffening groups, such as sulfone, carbonyl, or phenylene groups, are more resistant to corrosives than those with impact resistance compound groups, such as ether groups and polymethylene groups. Filled and reinforced polymers are generally more resistant to corrosives than unﬁlled or plasticized polymers.

9.4.1

Effects of the Environment

Unlike metals, polymers do not undergo electrochemical corrosion. However, they can be affected by organic solvents and are subject to thermal, photo- and chemical degradation. Because it is not feasible to test every polymer in each speciﬁc environment over a long time, data for their potential lifetime performance is essential. Information on the mechanisms of degradation can be used to predict their service lifetimes, at least on a semi-empirical basis. Polymers primarily consist of long molecular chains of similar carbon atoms that are in constant motion at temperatures above their glass transition temperatures (Tg). Because the bond energy that holds these atoms together is relatively weak, single bonds are readily dispersed when subjected to a surge of energy greater than the bond energy.

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9.5

Types of Degradations

9.5.1

Oxidative Degradation

The degradation of polymers is an auto-catalytic process that usually leads to chain-bonding or cross-linking as a result of hydro peroxide formation and subsequent decay. Because the oxidative attack is on the hydrogen atoms, the stability of polymers will be inversely related to the number of hydrogen atoms on the carbon atoms present in the polymer chain. This degradation is a chain reaction that consists of initiation, propagation, and termination steps. The oxidative degradation of polymers is catalyzed by heavy metals, such as copper, in a fast reaction in which the peroxide groups are decomposed and accelerate the degradation reaction.

9.5.2

Radiation Degradation

Long wavelengths are not sufﬁciently energetic to cleave the molecular chain bonds in organic compounds. However, radiation with wavelengths lower than 300 nm will selectively excite electrons in the polymer chain. The excited electrons may lose energy by ﬂuorescing, phosphorescing, or increasing the vibrational and rotational energy of the molecular chain bonds.

9.5.3

Photo Oxidation

This is the most common type of radiation degradation, sensitized by the presence of chromophores. A sensitizer, which may be present, may be raised to an excited state by the absorption of photons to produce free radicals, which are produced by a hydrogen atom abstraction that can lead to chain-bonding and/or cross-linking. The basic chromophore groups help to produce free radicals. Benzophenone, which contains chromophore groups, can be added as a sensitizer to excite cross-linking and branching, promoting the photo oxidation of polymers. Although pigments, such as carbon blacks, retard photo oxidation; other pigments, such as titanium dioxide, may act as a photo sensitizer at a wavelength of 480 nm. This sensitization is dependent on the presence of water and oxygen. Considerable attention has been devoted to ultraviolet radiation degradation. However, ionizing radiation also generates free radicals, which undergo degradation. Aliphatic polymers, such as polyethylene, are degraded by gamma irradiation, but aromatic polymers, such as polystyrene, are relatively resistant to high energy radiation. A sterilization dose of gamma radiation (2.50 Mrad) causes polypropylene to become brittle as a result of cross-linking and branching. Nylon is more resistant to gamma radiation. However, polyethylene oxide is sensitive to gamma ray oxidative degradation.

9.5.4

Mechanical Degradation

When stress is imposed on polymer chains through grinding, milling, stretching, impacting, or ultrasonic bonding, softening may occur and macro radicals may be produced in this so-called mechano-chemical reaction. These macro radicals

9.6 Moisture Effects on Nylon Molded Parts can add oxygen and produce peroxy compounds, which will create degradative reactions.

9.5.5

Microbial Degradation

Some polyesters, polyurethanes, cellulosic, and plasticized PVC can be degraded by micro-organisms. It has been observed that enzymes attack amorphous regions preferentially; therefore, the resistance of susceptible polymers to microbial degradation is related directly to the degree of crystallinity of the polymers. Environmental attack on polymers is a complex problem and is dependent on the chemical structure of each speciﬁc polymer. Generalizations can be made, but are related to the molecular structure of the polymer molecule, the extent of crystallization, the strength of the intermolecular forces, the bond energies of the covalent bonds, the presence of reinforcements, and the extent of crosslinking. Fortunately, when the failure of a polymer due to environmental exposure is not catastrophic, the rate of attack is progressive. In addition, the degradation rate can be monitored and is extrapolated through physical, instrumental or chemical analyses. When tests are conducted to verify published performance data, they should simulate anticipated end-use environment and evaluations should be made periodically over an extended period, preferably at several different temperatures and locations. One should not overlook the signiﬁcance of simple tests, such as changes in appearance, hardness, and density. Differential scanning calorimetry provides considerable data on thermal properties and infrared spectroscopy provides data on changes in molecular structure with time. Despite occasional failures, most properly selected polymers are successful in their applications. However, successful uses of polymeric components in plumbing ﬁxtures, dishwashers, and cooking utensils are not publicized. Other, less widely used applications in hostile environments will be equally successful if the proper polymer is selected for the anticipated end-use conditions. The published and unpublished information that is now available will minimize failures due to long term exposure to hostile environments. Considerable information is available in technical books for thermoplastics, technical design manuals from thermoplastic suppliers, and much more information can be obtained from the resin supplier’s technical representatives.

9.6

Moisture Effects on Nylon Molded Parts

Dry nylon parts made by any process are subject to dimensional changes caused by water absorption, crystallization, stress relief, and thermal expansion or contraction. Nylon parts are, however, found in many applications where dimensional stability is important because the major factors, water absorption and stress relief, often offset each other. The humidity of the environment is cyclical and the moisture content of nylon parts will ﬂuctuate around an average value. Once this average value is approached, the subsequent dimensional change depends on how much the humidity varies and the time available at ambient temperature for the nylon part to respond. Experience has shown that under normal conditions, the change due to variations in humidity is small, usually less than 0.20% in any one direction.

523

524

9 Thermoplastic Effects on Product Design The equilibrium moisture content of nylon depends on the amide group concentration, crystallinity, relative humidity, and temperature. The expansion effect of the moisture can also vary, because the speciﬁc volume of water in nylon may depend on humidity. An exact prediction of the dimensional change resulting from exposure to a speciﬁed relative humidity involves knowledge not only of the kind of nylon and part geometry, but also of thermal history and absorption isotherms. Conservative ﬁrst approximations can be made, however, if it is assumed that the volume change equals the moisture absorption typical of a nylon molded part, the part is stress free, and the change is isotropic. Nylon is a hygroscopic polymer and will absorb moisture from the atmosphere. Absorption of moisture from the atmosphere occurs slowly and continues until an equilibrium results between the moisture in the nylon and the moisture in the atmosphere. At equilibrium, no further absorption takes place. However, if the humidity should drop after equilibrium is reached, the reverse reaction occurs and the nylon slowly loses moisture. As the moisture is absorbed, some changes in the nylon also take place. A slight increase in dimensions results from the moisture absorption. Many physical properties change: toughness and elongation increase, while the yield stress, tensile modulus and hardness decrease. The absorption of moisture from the air at room temperature is a slow process, especially for thick-walled parts. A nylon molded part with 0.125 in wall thickness takes nearly a year to come to equilibrium in a 50% relative humidity and 73 °F ambient temperature. Boiling the injection molded nylon part in water is used to provide quick toughening for assembly purposes. However, boiling it continuously to equilibrium will result in a higher moisture absorption than will occur at equilibrium in a 50% relative humidity atmosphere, which is representative of average conditions. If the boiling is discontinued before equilibrium is reached, the surface of the injection molded nylon part will contain more moisture than the inside. A molded part with 2.50% of moisture concentrated on the outside surface will have different properties and dimensions than another equal part having a 2.50% of moisture uniformly distributed throughout the molded nylon part. The dimensional changes due to water absorption can be speeded up by moisture conditioning the injection molded nylon parts in boiling water. The time required to condition nylon to 50% relative humidity and 100% relative humidity (saturation) is given in Figure 9-7. At moisture levels below saturation, the water is concentrated near the surface and redistributes itself with time. The combined dimensional changes due to moisture absorption and temperature for annealed parts molded of unreinforced nylon 6/6 are shown in Figure 9-8. In this graph, the coefﬁcient of linear thermal expansion has been assumed to be 0.00005 in/in/°F. The moisture absorption characteristics of nylon 6/6 with 33% ﬁber glass reinforced related to time are shown in Figure 9-9. The weight increase of several ﬁber glass reinforced nylons when exposed to different levels of relative humidity is shown in Figure 9-10. The maximum linear changes in dimensions for several ﬁber glass reinforced nylons, along with their ﬁber ﬂow orientations (ﬂow and transverse directions), when they are submitted to different levels of relative humidity are shown in Figures 9-11 and 9-12.

525

9.6 Moisture Effects on Nylon Molded Parts 0.5

R. H. 50%

Wall thickness (in)

0.4

0.3

0.2

0 10

0.1

%

R.

H.

Figure 9-7 Time to moisture condition Nylon 6/6 in boiling water (Courtesy: Du Pont)

0 0.1

10

1

100

1.000

Time, (hours) 100% (R H)

Dimensional changes (Length/in)

0.030

0.025

90% (R H)

0.020 70% (R H) 0.015

50% (R H)

0.010

10% (R H) Annealed - dry

0.005

30%

) (R H

0.0

Figure 9-8 Unreinforced nylon6/6 dimensional changes vs. temperature at various relative humidities (Courtesy: Du Pont)

-0.005 -4

32

68

104

176

140

Temperature (°F)

0.125 in wall thickness 6

Moisture absorption (%)

100% (R H) 5 4 3 2

50% (R H)

1 0 0

50

100

150

200

Time (days)

250

300

350

Figure 9-9 Nylon 6/6, 33% ﬁber glass, moisture absorption vs. time (Courtesy: Du Pont)

526

9 Thermoplastic Effects on Product Design 0.125 in wall thickness 10

Nylon 6/6 (13% F G) Nylon 6/6 (33% F G)

Weight gain (%)

Nylon 6/6 (43% F G) Nylon 6/12 (33% F G) 1

0.1

Figure 9-10 Fiber glass reinforced nylon, weight gain vs. relative humidity (Courtesy: Du Pont)

0

20

40

60

80

100

Relative humidity (%)

0.125 inch wall thickness 0.8 0.7

Nylon 6/6 (13% F G)

Length change (%)

0.6 0.5 Nylon 6/6 (33% F G)

0.4

Nylon 6/6 (43% F G)

0.3 0.2

Nylon 6/12 (33% F G)

0.1 0

Figure 9-11 Fiber glass reinforced nylon, ﬂow direction change vs. humidity (Courtesy: Du Pont)

0

20

40

60

80

100

Relative humidity (%)

0.125 inch wall thickness 1.6

Width change (%)

1.4

Nylon 6/6 (13% F G)

1.2

Nylon 6/6 (33% F G)

1.0 Nylon 6/6 (43% F G)

0.8 0.6 0.4

Nylon 6/12 (33% F G) 0.2 0

Figure 9-12 Fiber glass reinforced nylon, width changes vs. humidity (Courtesy: Du Pont)

0

20

40

60

80

Relative humidity (%)

100

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9.7 Aqueous Potassium Acetate for Moisture Conditioning Nylon The previous information is intended to provide an understanding of the factors affecting dimensions of nylon parts and to give estimates of the amount of change to be expected. For precise dimensional changes of nylon, the manufacturer’s technical manuals provide sufﬁcient data, but careful testing of the actual part is often necessary.

9.7

Aqueous Potassium Acetate for Moisture Conditioning Nylon

A procedure for rapidly conditioning nylon molded parts to a uniform moisture level equivalent to that attained in a 50% relative humidity, is conditioning the injection molded nylon parts in aqueous potassium acetate. This procedure consists basically of immersing the injection molded nylon parts in boiling aqueous potassium acetate solution at the prescribed concentration until equilibrium moisture is obtained. In a solution consisting of 125.00 g of potassium acetate per 100.00 g of water, the nylon parts placed in the potassium acetate solution will come to the same moisture equilibrium that they would attain in a 50% relative humidity condition. The procedure is useful to obtain uniform moisture content in nylon molded parts. When physical properties are important, the technique is usually employed for nylon molded parts with wall thicknesses of 0.25 in or less. This treatment is not recommended for thicker molded parts, because the prolonged boiling at 250 °F required to reach equilibration will result in some hydrolysis of the nylon. This results in a lower relative viscosity and reduced toughness of the injection molded nylon parts.

After immersion has been done for a long time, some discoloration may appear in the potassium acetate solution and staining may be noticed on the injection molded nylon parts conditioned in this solution. Although the discoloration may detract from the appearance, it does not appear to have any adverse effects on the physical properties. Although the cause of the discoloration is not known with certainty, it has been noted that the presence of even small amounts of iron and copper in the potassium acetate solution result in severe discoloration. Immersion in potassium acetate solution at 250 °F has an annealing effect on the nylon. Approximately 70% of the molded-in stresses are relieved with this procedure, compared with the results of annealing the nylon part in oil at 300 °F. Figure 9-13 shows the ratio of potassium acetate to water for conditioning nylon 6/6 to moisture levels between 2.0% and 6.0%.

Desired % of water content

The procedure is recommended for conditioning injection molded nylon parts that are to be subjected to various physical tests. If the treatment is limited to test pieces with thicknesses of 0.25 in or less, the mechanical properties should not be changed beyond what would be expected to result from the addition of water. However, depending on surface/volume ratio, a small amount of the potassium acetate may be absorbed. For 0.125 in test bars, the amount of potassium acetate absorbed is considerably less than 0.10%. Although the absorbed potassium acetate solution does not appear to affect the toughness and the tensile properties, it does have a harmful effect on such electrical properties such as the surface resistivity. 8 7 6 5 4 3 2 1 0 0 20 40 60 80 100 120 140 160 180 200

Potassium acetate gm/100 gm water

Figure 9-13 Effects of aqueous potassium acetate solution on the percentage of moisture absorbed by unreinforced nylon 6/6 (Courtesy: Du Pont)

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9 Thermoplastic Effects on Product Design

9.8

Injection Molding Cycles

An overall injection molding cycle may vary from a few seconds up to several minutes. The important thing to remember, however, is that once the molding conditions are set into the controlling instruments of the injection molding machine, the identical cycle will be repeated accurately for the duration of the production run or until some effective change has been made by the operator or setup person. In a typical injection molding cycle, action is started by closing the safety gate at the mold area. As it is closed, the safety gate trips two safety limit switches, one for the machine hydraulic system and the other for the overall electrical system. From this point on, the injection molding machine operation is automatic. If the machine is equipped with a hydraulic clamp, oil enters a booster housed within the main injection system, causing the movable platen to advance at high speed but under low pressure. Just before the mold halves make contact, a limit switch is tripped, causing the oil pressure to be diverted from the booster ram to the main cylinder. Platen speed immediately drops off, but pressure increases. As the mold halves make contact, pressure builds up, triggering a pressure switch that signals the injection unit to begin its cycle. At this point, the reciprocating screw is in its retracted position behind a previously plasticized shot. Upon receipt of the signal from the pressure switch, the shutoff valve opens and the injector’s hydraulic system is actuated. This causes the screw to advance as a plunger and to inject the shot at pressures as high as 30,000 psi. This initial high pressure to counter the chilling effect of the mold is controlled by an adjustable relief valve or other ﬂow control device. Length of screw stroke, and therefore shot sizes, is also variable and is controlled by a timer. When the mold cavities are ﬁlled, a signal from this timer to an injection speed controller reduces pressure on the screw, permitting injection to continue at a reduced rate (hold pressure). The purpose of maintaining melt pressure is to minimize the possibility of heat sinks, increasing the molded part weight, and to reduce mold shrinkage during the mold cooling cycle. The injection and hold pressure part of the cycle is called the screw forward time. At the end of the dwell time, the timer signals the shutoff valve to close and the screw to resume its back rotation. Plasticized melt again advances to the front of the screw check valve, forcing it to retract against a back pressure until the shot-control limit switch is contacted. At this point, the clamp is still closed and the molded parts are cooling. On a signal from the timer, set according to cooling time requirements, the clamp opens slowly under pressure. Slow movement, mandatory to protect the molded parts, continues until a fast return limit switch is contacted, at which point the machine platen increases the mold opening speed. Fast retraction continues until another limit switch is contacted by a control rod actuated by the platen, which reduces speed for slow, smooth ejection of the molded parts, which drop into a box, or water tank, or onto a conveyor. Thus the machine cycle and its subordinate injection unit cycles are completed. In some cases, automatic operation of an injection molding machine is interrupted for manual removal of parts that might be damaged by the automatic ejection system. The overall injection molding cycle may be described as the total time required to produce one complete shot of one or more parts, depending on the number of cavities in a given mold. The injection molding cycle is not merely the time

529

9.9 Mold Cavity Surface Temperature that the polymer melt remains in the mold. It includes the time necessary for the mold to close and clamp, any safety or delay time required at the start of the cycle, the injection time (time required to ﬁll the cavity), the hold time, the time required to cool the molten material, mold opening, and the ejection time. The sum of these elements is known as the total overall cycle. The injection molding cycle conditions for semi-crystalline thermoplastics are different than the conditions for the amorphous thermoplastics. The cycle differences for these two thermoplastic families are illustrated in Figure 9-14.

9.9

Mold opening part ejected mold closing

Injection time

Machine charging time

Screw stop

Screw forward time

Mold Cavity Surface Temperature

During any consideration of the injection molding process, the mold cooling must not be overlooked. Obviously, if hot polymer is to set up within the mold cavity, the material must be cooled sufﬁciently to solidify and to allow the molded part to be removed from the mold. This, in essence, is the mold cooling portion of any cycle and will vary in time and temperature, depending on the geometry of the part and the choice of thermoplastic material. Mold cavity cooling can be considered as taking place with the mold at any temperature below the process temperature of the thermoplastic melt. Most commonly, the mold cavity surface will be maintained at a temperature ranging from 30 – 40 °F up to 200 – 350 °F. The mold cavity surface temperature should be uniform to allow for consistent mold shrinkage within the mold cavity to obtain good dimensional control and reduce the tendency to warp. When the injection mold is being designed and constructed, particular consideration must be given to the proper layout of the necessary cooling channels in both halves of the tool. Often, this detail is overlooked and the channels are added as an afterthought, with little attention given to the important function of these cooling lines for providing temperature equilibrium in the operation of the mold. The fact that optimum cycle times and overall quality of the molded parts may be jeopardized is completely overlooked. Proper layouts of cooling lines and their size, commensurate with the necessary heat exchange, are of prime importance in the design and construction of any new mold. It may prove difﬁcult, if not impossible, to alter the mold later to provide better and more efﬁcient cooling. Remember that tool performance and the quality of the molded parts depend largely on the ability to transfer heat rapidly and uniformly. To control mold shrinkage and warpage, it may be necessary to operate the mold cavity at an elevated temperature, causing a longer cycle time. When the molded part has a relatively thin wall section and high speed production is necessary, the mold may be operated using closed-loop refrigerated cooling water. Obviously, the heat necessary to effectively melt the thermoplastic is the same quantity of heat that must be removed before the part can be ejected from the mold. As the heat content of the thermoplastic increases and as the melt temperature goes up, it becomes even more important that the cooling system be capable of removing this greater amount of heat rapidly and efﬁciently. The way a mold cavity is cooled is of great importance to the injection molder. Lower mold temperatures make for increased part stiffness and strength, improved clarity, minimum shrinkage, and substantially reduced cycle times, if the thermoplastic material allows this condition.

Screw retraction time

Packing time

Semi-crystalline thermoplastics Mold opening part ejected mold closing

Screw stop

Injection time

Machine charging time

Screw retraction time

Screw forward time

Packing time

Amorphous thermoplastics

Figure 9-14 Injection molding cycles for both thermoplastic families

530

9 Thermoplastic Effects on Product Design

86˚ F.

122˚ F.

158˚ F.

194˚ F.

Figure 9-15 Unﬁlled Nylon 6/6, surface ﬁnish effects caused by mold temperature

Molds are usually water-cooled. There is, however, a limit to the number of cooling channels that can be incorporated into a mold base. These channels should be positioned so that the heat transfer from the melt is as uniform as possible. To obtain balanced heat transfer, more cooling must often be provided around the thicker walls of the molded part. The mold temperature has a great effect on the surface ﬁnish of the molded part. Figure 9-15 shows the mold temperature effects on nylon 6/6 surface ﬁnish. A cold mold causes a thicker amorphous skin, while a hot mold provides a smoother and thinner amorphous ﬁnishing. Minimum injection molded part warpage requires a differential in high mold temperature between both halves of the mold. It also requires that the entire cavity be cooled at nearly the same rate. The cavity temperatures around the gate areas should be lower and less cooling should be provided at the cavity section areas farthest from the gate. To obtain fast cycles, it is also important that the hottest portion of the mold, adjacent to the sprue and runners, receive the most concentrated cooling to minimize the total time required for the part to cool in the mold. Keeping the mold too cold is risky for two reasons: it may result in short shots and it may yield molded parts with molded-in stresses that can result in later stress cracking of the parts.

9.10

Mold Cavity Temperature Control

For efﬁcient molding, the temperature of the cavities should be controlled, which is normally done by passing turbulent water through the cooling channel in the mold. The rate at which the mold cools affects the total cycle time as well as the surface ﬁnish, tolerances, distortion, and molded-in stresses of the molded parts. High mold temperatures improve surface gloss and tend to eliminate voids. However, the possibility of ﬂashing is increased and sink marks are likely to occur. If the mold temperature is too low, the melt may freeze in the cavity before it is ﬁlled. The mold temperatures used are a compromise based on experience. Cooling the thermoplastic melt requires cooling the cavity uniformly and quickly, while maintaining satisfactory appearance and properties of the molded parts. It is not possible to have the whole mold at a uniform temperature. The sprue and the gates are areas where the melt temperature is high so that with uniform cooling, these areas would cool slowly. In the best mold designs, maximum cooling is directed to where the thermoplastic melt is hottest. This can lead to

9.10 Mold Cavity Temperature Control some sophisticated cooling systems with complex channels. However, additional mold design efforts and higher tooling costs are the solution to reduce the temperature efﬁciently to obtain the beneﬁts of faster molding cycles and reduction of molded parts rejection rates. Following are the factors that affect the cooling conditions in a mold: • The thermal conductivity of the mold material. If heat removal from the melt is expected to be a problem, it may be worth considering a mold material other than steel. For example, the thermal conductivity of beryllium copper is four to ﬁve times that of steel and although it is expensive, the cost may be offset by higher production rates and fewer rejects. • The surface area and nature of the cooling channels. The larger the surface area of the cooling channel, the better is the heat transfer to the circulating water. In addition, if the surface of the channel is rough, or if it contains interruptions to ﬂow, this will promote turbulence in the circulating ﬂuid and markedly improve the heat transfer. • The number of cooling channels and their proximity to the mold cavity walls. To avoid temperature gradients over the surface of the thermoplastic melt as it cools, it is necessary to have as many cooling channels as possible close to the mold cavity. • The ﬂow rate of the circulating water. The higher the volume ﬂow rate, the greater will be the amount of heat taken away from the cavity. • The cooling channels ﬂow characteristics require turbulent ﬂow (Reynolds Number higher than 3,000) of the water to obtain a greater heat transfer. The use of higher amounts of glycol (50.00%) on the cooling water reduces the viscosity of the ﬂuid, decreases the ﬂow and produces a lower heat transfer characteristic. • Special quick-disconnect ﬁttings without restricting the internal ﬂow channel and braided stainless steel hoses with Teﬂon® liners could be used on the mold. • Sufﬁcient water cooling is critical to the cycle times and for maintaining good temperature control in the mold cavity. This is important for accurate process control. • Pipe ﬁttings should be located so that they do not interfere with the molding machine’s tie bars or clamps, particularly during installation. • Mold plates should be thick enough for the size of water lines running through them. A temperature insulation plate should be used between each external clamp surface of the mold and the machine clamping plates, to avoid mold heat dissipation through the machine. • Corrosion problems must be addressed either by the choice of steels or water channel plating when constructing the tool or by rust inhibitors in the water during tool operation.

9.10.1

Mold and Post-Mold Shrinkage

High semi-crystalline materials, such as acetal homopolymer, show relatively high mold shrinkage. The mold temperature inﬂuences the crystallization rate of the resin during the solidiﬁcation of the part in the mold cavity. Therefore, it is a decisive factor in determining dimensional stability of the molded parts.

531

532

9 Thermoplastic Effects on Product Design The total mold shrinkage, which can occur in a molded part, is the product of mold shrinkage and post-mold shrinkage. Mold shrinkage is deﬁned as the change in dimensions between the dimension of the mold cavity and the solidiﬁed molded part 24 hours after it was ejected from the mold. Post-mold shrinkage occurs after the part has been aged. Mold shrinkage increases proportionally with the increase of mold temperature and the viscosity of the polymer. The molecular weight and the length of the molecular chain are higher in the high viscosity polymers. These high molecular structure materials have excellent impact resistance, but their mold shrinkage rates are higher and the melt ﬂow rates are lower than the medium viscosity polymers. Figure 9-16 shows mold shrinkage rates versus mold temperatures of two acetal polymers with different melt viscosities. The mold shrinkage in both polymers is increased as the mold temperature is increased. As with all thermoplastic materials, mold shrinkage depends very much on the injection molding pressure. Increasing injection pressure will result in reducing mold shrinkage. This general characteristic for acetal homopolymers is illustrated in Figure 9-17. In this graph, mold shrinkage is plotted as a function of set injection pressure for mold temperatures of 176 °F, 212 °F, and 248 °F, using an injection molding machine with a high compression screw at a melt temperature of 420 °F.

Shrinkage (in/in)

0.028

High viscosity

0.024 0.020

ity

Medium viscos 0.016

Melt temperature 420° F. Mold temperature 200° F. Part wall thickness 0.125 in Screw forward time 26 sec.

0.012 0.008 140

160

175

200

215

230

250

Mold temperature (°F) Figure 9-16 Acetal homopolymer shrinkage vs. mold temperature (Courtesy: Du Pont)

Shrinkage (in/in)

0.024

0.020

175° F. mold temperature 212° F. mold temperature

0.016

250° F. mold temperature 0.012

Melt temperature 420° F. Injection pressure 13.000 psi Part thickness 0.125 inch Screw forward time 26 sec. 10.000

14.000

20.000

Injection pressure (psi) Figure 9-17 Acetal homopolymer shrinkage vs. injection pressure (Courtesy: Du Pont)

9.11 Process Condition Effects on Mold Shrinkage Higher injection pressures can be used if clamping forces permit, but this will not necessarily improve mechanical properties of the molded part. Use of high injection pressure may result in overworking the hydraulic system of the molding machine, leading to shot-to-shot variations. Injection pressure may be limited by the molding machine’s clamping force when the projected area of the mold cavities and runner is large. Exceeding the force will open the mold and ﬂash the molded part. Injection pressure may also be limited by mold construction. Insufﬁcient mold plate support will allow mold deformation and will also result in ﬂash problems. Insufﬁcient injection pressure will produce underpacked molded parts, voids, molded-in stresses, and therefore lower mechanical properties and poor dimensional stability of the molded parts. Abrupt variations in injection pressure during the screw forward time may cause changes in crystalline structure, which may lead to a decrease in mechanical properties and to poor dimensional control of the molded parts. If the screw forward time is less than optimum, mold shrinkage increases, warpage in the part, deformation, molded-in stresses, and voids (porosity) may also appear. To ensure that the screw forward time is effective, there has to be a good transmission of the hydraulic pressure of the machine, via the screw and check valve, to the melt in the mold cavity. This transmission is only possible if there is a melt cushion of 0.125 to 0.25 in in front of the check valve. Since this cushion is molten resin, it will act as a ﬂuid and thus transmit the hydraulic pressure to the cavity. It is very important to avoid leakage of the melt through the check valve while maintaining the injection pressure.

9.11

Process Condition Effects on Mold Shrinkage

Thermoplastic processing normally involves heating thermoplastic granules or powders to above their melting point and then forcing the melt into a cooled mold until it is solid enough to handle. During this process, three mold shrinkage steps may take place. First, as the polymeric material goes from melt (liquid) to solid, a change in contraction rate occurs. Second, some polymers crystallize below the melting point and there is a volume change on crystallization, because crystalline regions have a lower coefﬁcient of expansion than amorphous regions. Therefore, semi-crystalline polymers shrink more than amorphous polymers. To be crystalline, the molecules must have the correct shape to be able to line up physically with each other and lie parallel in the crystalline areas without the coupling of side chains. Crystallization starts below the melting point, but can continue long after the product cools to room temperature even for days or weeks if room temperature is above the glass transition temperature (Tg). This can cause unexpected warpage. Third, normal thermal contraction during cooling will also take place. Contraction is anisotropic because of orientation. If the part is constrained during cooling, as in injection molding, stresses can be built in that will dissipate elastically as the part is ejected from the mold or that will appear when the part is in service, causing part distortion. This is most likely to occur if the part is subjected to high temperatures in service. The rate of cooling and the temperature of the part upon extraction from the mold cavity will also affect the amount of mold shrinkage, especially with a semicrystalline polymer. Thick wall sections will shrink at a different rate than thin wall sections and will be at a higher temperature leaving the mold cavity.

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534

9 Thermoplastic Effects on Product Design Differential cooling in the mold cavity, wall thickness differences, and ribs can all cause stresses between different areas of the part. These factors can also cause part distortion on removal from the mold cavity. Cooling ﬁxtures can be used to constrain the part until it reaches room temperature after it is removed from the mold cavity. The part can also be left in the mold cavity longer, or the tool can be run at a different temperature to try to reduce warpage. For some semi-crystalline materials, nucleating agents can be used to increase the rate of crystallization. This will have the added beneﬁt of reducing crystal size and giving more, better dispersed crystals. The high pressures used in such processes as injection molding can both reduce some of the effects of mold shrinkage, by packing out the mold, causing distortion of the molded product because of increased internal stresses. With glass ﬁber reinforced materials, mold shrinkage of the resin away from the surface during molding can leave the glass ﬁbers proud (raised above a surrounding area), resulting in an unacceptable surface ﬁnish. A similar effect is found with thermosetting materials. Some of the worst effects of mold shrinkage can be overcome by using blowing agents to avoid shrinkage over ribs and bosses or by using processes that reduce pressure gradients in the polymer melt. Thermosetting reactions generally result in a volume loss. The degree of mold shrinkage varies with the type of reaction, the temperature at which the reaction takes place, and the type of bonds being formed. After high temperature reactions, the thermal contraction that occurs on cooling can cause internal stresses and sometimes cracking of the more brittle resins. This can also cause loss of adhesion to reinforcing ﬁbers. Mold shrinkage also leaves glass ﬁbers above the surface. By allowing the reaction to take place at the visible surface ﬁrst, the mold surface is replicated, shrinkage of the resin occurs progressively toward the back surface, which will have a very poor appearance. Heating the face of the mold to a higher temperature can result in this effect. Part distortion often occurs as one surface is still reacting and shrinking after the other has solidiﬁed. Mold shrinkage in thermosets is reduced by adding ﬁllers, by adding a thermoplastic resin to absorb the monomer and expand during the heating cycle (as in sheet molding compound processes), by gas formation during the reaction (as in polyurethane reaction injection molding), or by adding a blowing agent. The addition of ﬁllers can signiﬁcantly reduce the apparent shrinkage of both thermoplastic and thermosetting resins. This reduced shrinkage produces a much more stable part as molded. With higher ﬁller loading in thermoplastics, the part is much less likely to distort on removal from the tool or to warp in service. This increased stability is demonstrated with most processing methods. Most non-plastic materials show relatively small changes in linear expansion and contraction under the inﬂuence of temperature variation. On the other hand, most thermoplastics exhibit considerable dimensional change caused by environmental changes. Perhaps the most signiﬁcant change in thermoplastics dimensions occurs with phase change, causing the greatest concern to mold makers and mold designers. It also represents a major problem to the injection molder. When going through the normal cycle of solid state to thermoplastic melt and back to solid state, thermoplastics exhibit dimensional changes, which vary widely from one resin family to another.

9.11 Process Condition Effects on Mold Shrinkage There is no practical way of accurately predicting the exact mold shrinkage allowance for any thermoplastic material. A concerted effort has been expended, particularly by the thermoplastic resin suppliers, to develop meaningful and useful data on the subject of mold shrinkage. ASTM has set up a speciﬁcation for determining mold shrinkage (D-955). But as with most laboratory determinations, the data developed has little or no meaning to the molder, because the parts he must produce rarely conform to the dimensions of the test bars or discs. In addition, other variables, such as gating, part geometry, variations in wall thickness, and the efﬁciency of the machine itself, just to name a few, can make this data not only meaningless but also misleading. Most plastics engineers must rely on their past experience both with the material under consideration and the type of part they are designing for, in order to closely approximate the mold shrinkage allowance to be made for a particular mold design. Even with this experience, there is a possibility that this will not be the exact allowance needed. In such cases, one must rely on the ability to adjust the cycle to compensate for deviations or, if they are too great, the mold cavity itself must be reworked. The beginner is cautioned, not to rely too heavily on graphs and charts, which have found their way into trade literature and which proclaim to be accurate appraisals of the true mold shrinkage allowance for speciﬁc materials. Using such information can often lead to an improperly designed mold. When in doubt, the most logical source of data is the technical consultant of the thermoplastic resin supplier and it is advisable to consult more than one source. A certain amount of mold shrinkage is inevitable in any process that involves cooling of a material from an elevated temperature. Mold shrinkage must be taken into account when a mold is designed by allowing for a calculated mold shrinkage based on the properties of the resins. The injection molder can prevent excessive mold shrinkage, which will make close tolerance impossible, by controlling the operating conditions, such as molding at lower injection temperatures, running a cold mold, or by packing the mold. Packing can be achieved by molding at either moderate melt temperatures and high melt pressures, or by molding at fairly high temperatures and moderate pressures. However, excessive temperature or pressure may cause a mold to ﬂash. Another means of reducing mold shrinkage is through the use of high injection pressure and extended injection time, allowing additional resin to ﬂow into the mold as the material in the mold cavity cools and shrinks, again packing the mold as much as possible, short of causing the mold to ﬂash or the part to stick in the mold cavity. Before undertaking corrective action for solving this quality problem with the molded part, be sure the part is fully packed out. To be sure that the molded parts are packed out, measure and record the weight of several shots (parts only). Add time, normally ﬁve to ten seconds, to the screw forward time. Then repeat weighing and recording several shots. If there is no weight change, you have either a maximum weight or a premature gate freeze off. Premature gate freezing would be evident by the presence of voids near the gate of a part, cut in half through the gate area. Adjustment of the gate dimensions to retard freezing would be needed. If a weight increase is observed, repeat incremental increases in screw forward time until there is no further weight increase. Part weight is now optimized and other corrective actions may be considered if they are still needed.

535

536

9 Thermoplastic Effects on Product Design 0.0032

1.72

Part weight (gm)

Weight

1.68

0.0028

1.66

0.0026 0.0024

Shrinkage

1.64

0.0022

1.62 Melt temp. 420° F. Mold temp. 200° F. Wall thickness 0.12 in

1.60

Mold shrinkage (in/in)

0.003

1.70

0.002

1.58

0

5

10

15

20

25

30

35

Screw forward time (sec) Figure 9-18 Acetal weight, mold shrinkage vs. screw forward time (Courtesy: Du Pont)

The effect of screw forward time on part weight and mold shrinkage on an acetal homopolymer test bar (5.00 in long, 0.50 in wide, and 0.125 in thick) is shown in Figure 9-18. Maximum part weight and minimum mold shrinkage for an acetal homopolymer occur in about 24 seconds of screw forward time (injection plus packing time), which is the time when the gate freezes and maximum weight is obtained in the parts. It also assumes that the check valve functions properly and maintains a cushion of melt in front of the screw and check valve. No matter what remedies for mold shrinkage the molder chooses, whenever he works to close tolerances, all molding conditions must be controlled accurately. In addition, many of the means for reducing mold shrinkage can cause post-mold warpage and the two factors must be balanced against each other. Mold shrinkage is probably the main problem in precision molding. In addition to the volume change, there are also other contributory factors. For example, if there is orientation in the molding then mold shrinkage will occur due to relaxation of the stretched carbon to carbon linkages in the molecular chain. This tends to be greater in the alignment ﬂow direction than in the transverse direction. Another important consideration is whether the thermoplastic is semi-crystalline or amorphous. The difference between these two types of materials may be seen clearly when they are heated. For a semi-crystalline thermoplastic there is an abrupt change in viscosity at the melting temperature. Below this temperature the material is essentially solid, whereas at higher temperatures it is a relatively low viscosity liquid. It may also be seen that the slope of the characteristic at the high and low temperatures is small so that changes of temperature in regions away from the melt temperature has little effect on viscosity. For amorphous materials on the other hand, there is a more gradual change in ﬂuidity as temperature is increased. Figure 9-19 illustrates the effects of mold shrinkage versus screw forward time for acetal homopolymer of different wall thicknesses, to obtain the optimum mechanical properties and good surface appearance. The optimum injection speed for a part depends on its geometry, the size and location of the gate, and the melt temperature. When molding thin wall section parts, high injection speeds are usually required to ﬁll the part before it freezes.

537

9.11 Process Condition Effects on Mold Shrinkage 0.0035 Melt temp. 420˚ F. Mold temp. 200˚ F. Inj. press. 13.000 psi

Mold shrinkage (in/in)

0.312 in 0.003

0.25 in 0.203 in

0.0025

0.002 0.156 in

0.125 in

0.078 in

0 0

20

40

60

80

100

120

140

Screw forward time (s) Figure 9-19 Acetal homopolymer, mold shrinkage vs. screw forward time (Courtesy: Du Pont)

Screw forward time is the time during which hydraulic pressure is supplied to the screw check valve. Correct screw forward time is a function of the wall thickness of the part and changes slightly with mold temperature. As a result, in addition to the mold shrinkage effects mentioned earlier, for semi-crystalline thermoplastics there will be shrinkage due to the closer packing of the molecules in the crystalline state. Hence, the mold shrinkage of these materials is high, typically 1.0% to 4.0%, as compared with 0.3% to 0.7% for amorphous materials. Semi-crystalline materials also facilitate faster cycle times, because once the temperature drops below the melting temperature there is little change in viscosity and components may be ejected from the mold cavity while still relatively hot. When a material has a high mold shrinkage it is more difﬁcult to maintain close tolerances. An anomaly exists in that partially semi-crystalline materials, which exhibit large mold shrinkage, are more suited for precision engineering components that demand close tolerances, because of their physical properties, The successful molding of such molded products therefore needs considerable expertise on the part of the mold designer and machine setter in order to obtain a consistent acceptable quality. Mold shrinkage of semi-crystalline resins is dependent on the following factors: • Mold temperature • Injection pressure • Screw forward time • Melt temperature • Melt viscosity • Spue, runner, and gate sizes

538

9 Thermoplastic Effects on Product Design • Part wall thickness • Types and quantity of additives, reinforcements, or ﬁllers Good molding practices lead to the selection of the best set of molding conditions so that the mold shrinkage of the thermoplastic resin is near the recommended value obtained in test plaques in laboratory test conditions. For parts of uniform wall thickness, mold shrinkage tends to be uniform throughout the molded structure. For parts having variable wall thickness, mold shrinkage will tend to be nearly uniform, if the correct gate is properly sized, is located at the heaviest wall section of the part, and screw forward time equals or exceeds gate freeze time. If these criteria are not met, mold shrinkage will tend to be greater in the thick wall sections of the molded part. In practice, actual mold shrinkages may vary for the following reasons: • The molding machine is too small for shot size and production rate. • Insufﬁcient clamp capacity of the molding machine. • Melt temperatures too high or too low. • Screw forward time shorter than gate freeze time. • Freezing of the gate (small size) occurs too quickly. • Operating with mold temperature (mold cavity surface) other than recommended by the resin supplier. • Poor mold temperature control in each cavity or from cavity to cavity. • Injection pressure too low. • Poor gate location for adequate ﬁlling, may result in voids. • Polymer melt leakage through the check valve. Mold Shrinkage of Reinforced Thermoplastics Mold shrinkage of ﬁber glass reinforced thermoplastic compounds is less predictable because of the ﬁber orientation effects. The mold shrinkage in the direction of ﬂow tends to be signiﬁcantly less than that in the transverse direction. In addition, this difference is highly dependent on the part geometry, type, location, and number of gates.

9.12

Post-Mold Shrinkage

Mold shrinkage is basically the result of thermal changes in the thermoplastic melt while it is in the mold cavity. Post-mold shrinkage refers to the contraction of the solidiﬁed and ejected molded part caused by relaxation of molded-in stresses. Although in many injection molded thermoplastic parts, post-mold shrinkage does not play an important a role, in certain applications it can present a troublesome problem for the molder. In closures and in components which are to become components of an assembly, too much or too little post-mold shrinkage can be a valid cause for rejection of the molded part. In addition, uneven mold shrinkage causes warpage of the molded items.

9.12 Post-Mold Shrinkage Non-uniform shrinkage of a molded part after it is ejected from the mold causes bending or twisting out of shape and alters not only its dimensions but also its contours and angles. Warpage occurs mainly in large and ﬂat molded components and, though undesirable in any molding, occurs particularly in such items as container covers and closures. When an injection molded thermoplastic part warps after being ejected from the mold, it assumes its natural shape by relieving the stresses imposed on it while being shaped in the mold cavity in the viscous state. The problem for the molder – and it is often a difﬁcult one – is to minimize the molded-in stresses, which the part may later remember and relieve when cooling to room temperature. The molded-in stresses are generated in the mold cavity by such operating conditions as excessive injection molding pressures, uneven mold cooling, or too low a melt temperature, to mention only a few molding conditions. There is no single, clear-cut remedy for warpage. The internal stresses set up in the molded part during mold cooling may be reduced by adjusting the molding process conditions, redesigning the part or the mold, switching to another resin, or some combination of these steps. Generally, the best resistance to warpage is achieved by maximum melt temperature, high mold temperature, minimum injection pressure, and short screw forward time. Molding at high melt temperature tends to diminish the elastic memory of a resin and thus reduces the tendency to create stresses that might cause warping. Increasing the temperature of the mold cavity allows stresses to relieve themselves before the melt sets or freezes, thus reducing the tendency to warp. In addition, separated mold temperature controls are a must to produce warpfree parts. Part design and mold design have much to do with warping. Warpage may be increased for instance, if the part has greatly dissimilar wall thickness sections, if the gate is located in a thin wall section of the part, if the sprue is poorly placed, or if the mold is built up of inner surfaces with unequal heat dissipation. When mold shrinkage occurs in the mold cavity and the part reaches room temperature, further mold shrinkage may occur as time passes. The post-mold shrinkage is irreversible and it is determined by cooling rate, that is, by mold temperature, screw forward time, injection pressure, melt temperature, and part wall thickness. Figure 9-20 shows the post-mold shrinkage of an acetal homopolymer resin after exposure for 1,000 hours (after molding) at various temperatures. At exposure temperatures slightly above room temperature (100 °F), post-mold shrinkage is less than 0.002 in/in for parts with a wall thickness of 0.062 in molded in a cold mold. After exposure for one year, the post-mold shrinkage of parts molded n a cold mold will become signiﬁcant, while it is still minimal for parts molded at 200 °F or higher. Post-mold shrinkage after continuous exposure of parts made of acetal homopolymer with walls of 0.125 in thickness, molded at various mold temperatures, is shown in Figure 9-21. When injection molded thermoplastic parts requiring close tolerances will be exposed for prolonged periods at elevated temperatures, it may be necessary to anneal the parts after molding, using a secondary operation. However, for the moderate temperature exposures of most applications, good dimensional stability will be achieved by using higher mold temperatures.

539

540

9 Thermoplastic Effects on Product Design 0.012

1.000 hr exposure time Melt temp. 420° F. Wall thickness 0.062 in 100° F.

150° F.

0.008

200° F.

0.006

Mold temperature

Mold shrinkage (in/in)

0.010

0.004 250° F. Annealed 0.002

0 100

125

175

150

200

225

Exposure temperature (°F) Figure 9-20 Acetal homopolymer mold shrinkage at various mold conditions vs. exposure temperature (Courtesy: Du Pont)

Post mold shrinkage (in/in)

0.0035

Melt temp. 420° F. Wall thick. 0.125 in

0.003

Mold temp. 150° F. Exposure temp. 150° F.

0.0025 0.002

Mold temp. 250° F. Exposure temp. 150° F.

0.0015

Mold temp. 250° F. Exposure temp. 73° F.

0.001 0.0005

Annealed

0 1

10

100

1.000

104

Exposure time (hour) Figure 9-21 Acetal homopolymer post-mold shrinkage at various mold and exposure temperatures vs. exposure time (Courtesy: Du Pont)

5

10

541

9.13 Weld Lines Thermoplastics Annealing Annealing of a semi-crystalline thermoplastic molded part is occasionally useful to relax molded-in stresses and to stabilize dimensions. The amount of mold shrinkage that may occur depends on the variables that affect cooling rate, especially mold cavity temperature and part wall thickness. At a 100 °F mold temperature, a 0.125 in thick test bar made from acetal homopolymer resin was found to shrink an additional 0.011 in/in in length during the annealing process, while a similar bar molded at a 250 °F mold temperature had an annealing shrinkage of only 0.004 in/in. Since the change in dimensions during the annealing process is related to gate location, part wall thickness, mold shrinkage, melt ﬂow, ﬁber reinforcement orientation, process conditions, etc., it may not be uniform in all directions. The amount of change in any speciﬁc case must be determined experimentally.

9.13

Weld Lines

The area or plane in which two cold melt fronts meet as the cavity is ﬁlled is commonly called a weld line or knit line. In designing an injection molded thermoplastic part, a number of factors affect part and weld line strength. Figure 9-22 shows how the weld lines are formed during the injection molding process. Weld lines create a weakness and change material strength characteristics where cold melt fronts reunite. The extent of property change depends on the ability of the two melt fronts to knit together homogeneously. The following conditions affect weld line integrity: base resin type, part thickness, mold design, resin impact modiﬁers, resin mold released additives, reinforcements, molding process conditions (such as temperature and viscosity of the molten thermoplastic when they come together), and lubricants sprayed on the mold cavity surfaces. Different resins will exhibit different characteristics of tensile strength retention at the weld line. The effects of the inherent weld line integrity of the thermoplastic resin on property loss are indicated in Table 9.1. Polysulfone (PSU) exhibits little or no tensile strength loss at the weld line, while styrene acrylonitrile (SAN), PP, and polyphenylene sulﬁdes (PPS) exhibit some tensile strength loss in an unﬁlled state. Part wall thickness does not play a major role in weld line strength retention, except in thin-walled parts. The percentage of loss in strength at the weld line is greatly increased by ﬁbrous reinforcement. However, absolute weld line strength increases as more or stronger reinforcement is added. The large loss of strength at the weld with ﬁber glass reinforcement is caused by the ﬁber glass from both melts not blending with each other. As the melt fronts meet, the ﬁbers are turned 90° from the direction of ﬂow, but the glass ﬁbers do not inter-cross, because the skin of the cold melt front does not allow the ﬁber glass to be in front of the melt ﬂow. The ﬁber glass is 90° from the weld line direction, so the ﬁbers are aligned perpendicular to the applied stress and offer little reinforcement at the weld line. Particulate ﬁllers, such as talc, milled glass, polytetraﬂuoroethylene (PTFE), and others, have very low aspect (length/diameter) ratios. Therefore, the compounded resins using these ﬁllers have properties that are independent of the ﬁller orientation and they behave similarly to the unmodiﬁed base resin.

Core pin

Melt pattern

Gate

Weld line Melt pattern

Core pins

Gate

Weld lines

Figure 9-22 Weld line forms when melt ﬂow splits around core pins

542

9 Thermoplastic Effects on Product Design Table 9-1 Effects of Resins on Weld Line Strength

Resin

Reinforcement

Fillers

Tensile Strength 1 Gate Kpsi

Gate

Split flow

Split flow

Strong wall

None

PSU

% Retained

9.6

9.6

100

30% glass ﬁber

16.8

10.4

62

SAN

None

11.3

9.0

80

SAN

30% glass ﬁber

16.2

6.5

40

PP

None

5.4

4.7

86

PP

20% glass ﬁber

9.1

4.3

47

PP

15% glass ﬁber

15% glass beads

6.5

2.7

42

PP

30% glass ﬁber

15% PTFE

9.7

2.8

29

PPS

None

8.8

7.3

83

PPS

10% glass ﬁber

10.3

3.9

38

PPS

40% glass ﬁber

20.5

4.1

20

Strong weld lines are critical, because the properties in the weld line region decline signiﬁcantly compared to those in the rest of the part; these lines become likely points of part failure. Weld lines can also cause irregularities in the surface appearance of the molded part, making it more prone to wear. Therefore, weld lines should be located in less critical areas if possible.

Weld line Overflow tab pocket Melt flow pattern

PSU

2 Gates Kpsi

Guidelines to overcome these issues and to produce quality parts are given in the following:

Gate

Thermoplastic Resins • Provide the best resin with good weld line strength • Avoid contaminated material • Use only well dispersed color resins Core pin Strong wall Split flow

• Do not use reground material mixed with the virgin resin • Do not use resins with internal or external lubricants. Product Design

Weld line

• Increase the wall thickness to permit easier melt ﬂow

Overflow tab pocket

Figure 9-23 Improved weld line by using overﬂow tab pocket

• Use thick ribs to act as a conveying melt channel to improve and redirect the melt ﬂow in the cavity • Modify the part design to shift and/or eliminate obstructions to ﬂow

Hole (core pin)

Melt flow direction

Weld line Boss Boss rib

Figure 9-24 Improve boss strength using a rib on weld line location

• Holes may be molded partially to eliminate weld lines • Allow the use of an overﬂow weld line tab pocket in front of the weld line that will be removed after molding, the weld line melt will be transferred from the cavity to the tab pocket, as shown in Figure 9-23 • Add a rib on top of the weld line to increase the strength in this area as shown in Figure 9-24.

543

9.13 Weld Lines Countersunk hole

Example 9-1 Figure 9-25 shows a typical poor design: the screw countersunk hole is too close to the edge of the part; the edge gate is located at the opposite edge, in line with the hole; the countersunk screw produces radial forces directly on the weld line and causes the weld line to break. The following part design modiﬁcations are recommended: replace the countersunk screw head with a ﬂat head to eliminate radial forces; increase the hole- edge distance to one screw head diameter. Move the edge gate to relocate the weld line to a safer and bigger cross section area. Edge gate

Poor design

Injection Molding Machine Problems • Nozzle inside channel diameter too small and/or too long • Poor nozzle temperature control (heat variations)

Countersunk screw force

Weld line

• Check valve in poor condition can not maintain high pressures for injection and packing, the melt leaks back to the screw • Mold temperature unit capacity is too small; it cannot remove the heat quickly enough to maintain the uniform cavity temperature • Plastifying screw too short and of low compression ratio, cannot provide consistent high injection pressures • Interruptions of the screw injection speed during ﬁlling the cavity • Screw return stop control can not maintain uniform melt shot size. Mold Problems • Lack of venting or inadequate mold venting system; the use of proper venting located in front of the weld line areas is the most important parameter, because weld line strength is highly dependent on removal of gasses trapped inside the cavity

Melt flow pattern

Edge gate

Operational problems Weld line moved to a less critical area

Increased edge distance

Flat head screw

• Deﬂection of core support • Complicated runner lay-out and/or small runner diameters • Gate too small or poor gate design • Gate location and weld line too far from each other • Poor mold cavity temperature lay-out system.

Change gate location

Meltflow pattern Recommended design

Figure 9-25 Design modiﬁcations to improve weld line strength

544

9 Thermoplastic Effects on Product Design Weld lines

Example 9-2 By changing the gate location, the melt ﬂow direction changes and the weld line forms in a more ideal area of the cavity as shown in Figure 9-26.

Molding Process Parameters • High melt temperature is needed to ﬁll the cavities quickly and a hot melt front that moves fast insures a strong weld line Flow

• High mold temperatures slow down the crystallization rate of the melt, allowing the melt to pack and form a strong weld line during the cooling cycle

Flow

Poor weld line strength Weld lines Weld lines

• The right amount of clamp pressure should be used to close the mold; too much clamp pressure collapses the mold vent channels and the entrapped air cannot escape from the cavities; with too low a clamp pressure the melt pressure pushes the mold parting line apart, ﬂashing the cavities • Use slow clamp speed just before closing the mold to remove the entrapped air from the runner system and cavities

w Flo

Flo w

• Fast injection speed and high pressure are needed to allow the hot melt to ﬂow quickly, removing the entrapped air to form strong weld lines

Improved weld line strength Figure 9-26 Improve weld line strength by gate relocation

• Increase the shot size a little, to create a melt cushion in front of the check valve for packing the cavities • The nozzle should be without obstructions; the internal channel should be properly dimensioned, matching the mold sprue, without sharp corners to avoid shearing the melt and with a good temperature control system to avoid overheating the melt in the middle of the nozzle or freeze-off at the nozzle tip because of heat dissipation through the sprue • Do not use mold release spray on the cavities and/or runners; the melt front collects the lubricant from the surface walls and moves it to the middle of the weld line area when both melt fronts meet, causing a weak weld line.

545

10

Injection Mold Design

This chapter provides practical information for planning the construction of a thermoplastic injection production mold. It discusses the function of the molds, categories of molds, mold bases, types of steels used in molds, common types of mold designs, the effects caused by the product design on the injection molding process, mold speciﬁcations, mold design check lists, and the basic mold design principles. The thermoplastic injection molds may seem to be a very expensive piece of steel, with the cost of some molds higher than those for an injection molding machine. However in the long run, the mold represents only a very small fraction of the part cost. Good planning during the construction of the molds involves a lot of communication between the product designer, tooling engineer, process engineer, mold designer, mold builder, and resin supplier’s technical representative. The thermoplastic injection molding process obviously includes the use of the mold necessary to produce the end product. It is important to have some understanding of the product design, the thermoplastic resin, the injection molding process, and the technical information required to design and build the mold. What is a thermoplastic injection mold? Simply stated, it is a collection of steel plates and other mold components which, when properly assembled and installed within the injection molding machine, is capable of producing the required molded products from a given thermoplastic material. The majority of thermoplastic injection molds are made using high quality tool steel, designed and built by qualiﬁed mold makers, to exacting standards, capable of withstanding the high injection pressures and elevated temperatures of the process together with the usually fast cycles, which make injection molding an economically feasible manufacturing process. Thermoplastic injection molds require a sprue bushing to provide entry of the viscous melt into the cavity. Because the mold must be mounted into an injection molding machine, provision must be made for proper clamping of the mold to the machine platens and for the ejection of the molded part. Figure 10-1 shows a typical thermoplastic injection mold, together with the most common standard mold components.

10.1

Classiﬁcation of Injection Molds

The plastic industry classiﬁes thermoplastic injection molds in three general categories: prototype molds (25 to 1,000 parts), production molds (low volume from 1,000 to 10,000 parts) and high volume production molds (10,000 to 2,000,000 parts). Prototype Molds A prototype thermoplastic injection mold represents a preliminary step required in the development process of a new product. Prototype molding is used to investigate the molding characteristics of the resin, mold shrinkage, gating,

Figure 10-1 Typical mold for injection molding of thermoplastics

546

10 Injection Mold Design dimensional control of the molded part, molding process conditions, and molding cycle. The prototype molded parts are used for product quality control tests and occasionally can be used to meet initial market testing requirements. In simulating the production part, prototype molds become a relatively inexpensive learning device that the part designer can use to pinpoint and correct potential product design problems or material selection problems before expending on a production mold. A prototype mold may consist of an existing mold frame, interchangeable soft cavity inserts, simple cooling and ejection systems, manual insert loading and removal. Production Molds A production mold is built using a standard low cost mold base for the hardened tool steel cavities; this mold should be able to produce parts on demand at established production rates. The mold should allow easy repair and facilitate mold venting of the cavities to allow entrapped air and melt volatiles to escape during the molding cycle. The production mold must also have an automatic ejection system and a mold temperature control for consistent cooling of the thermoplastic melt to ensure minimum cycle time, lowest cost, and consistent quality. High Volume Production Molds These types of molds should have all the qualities of production tools, have multiple cavities and have fully interchangeable mold components. A high volume production mold should be designed to overcome any adverse outside force, with easy maintenance design. For example, how often do you start to take a mold apart only to ﬁnd a multitude of inserts without numbering or position marking? How about pry slots? How about jack screw holes for cavity removal? These obstacles can be overcome through a good planning process for the design and construction of the mold, a customized prevention maintenance program, and a protective surface coating on the steel to prevent corrosion and erosion.

10.2

Effects of Product Design on the Injection Molding Process

The product designer must have knowledge of the part geometry, because it may create problems during the molding process; he must also be familiar with the properties of thermoplastic materials, the injection molding process in general, mold design, and the quality of mold construction and product design required to produce functional thermoplastic molded parts. Producing quality thermoplastic parts requires converting the functional requirement of the application into a design. The product design’s geometrical conﬁgurations should not only satisfy functionally, but it also has to meet the conditions required by mold design and construction and operation of the mold in order to produce quality parts and guarantee efﬁcient molding operation. When designing injection molded thermoplastic parts, the product designer must also be aware of some part conﬁgurations that pose potential problems during the injection molding process. The part design requirements include uniform wall thickness, parting line location to balance the heat removal from both sides of the cavities, smooth internal corners, draft walls (to facilitate part

10.2 Effects of Product Design on the Injection Molding Process removal from the cavity), elimination of feather edges, elimination of fragile deep pockets (long thin cores), provide location for the gate, allow large permissible surface area for ejection, specify typical part dimension tolerances for plastics, and avoid the use of high-gloss surface ﬁnishing for the product.

10.2.1

Uniform Wall Thickness

Products that incorporate abrupt changes in wall thickness will create major mold design problems regarding the temperature control system of the mold. Abrupt changes in wall thickness make it difﬁcult to maintain a uniform temperature throughout the mold cavities during the molding cycle. After the thermoplastic melt has been injected, variations in part wall thickness do not allow the walls to cool at consistent rates. Thick walls will shrink more than thin walls, causing part warpage, voids on thicker wall cross sections, poor dimensional control, long cycle times, poor surface ﬁnish, and structural defects.

10.2.2

Balance Geometrical Conﬁguration

The positioning of the cavity should be balanced on both sides of the mold parting line. Both halves of the cavity should be subjected to the same volume of polymer melt for uniform cooling in the mold cavity. If one side of the cavity is injected with more melt than the other side, this side will become hotter. The hotter side of the cavity will have the tendency to stick on the deep hot spots, causing warpage, poor surface ﬁnish of the molded part, and long cycle times.

10.2.3

Smooth Internal Sharp Corners

Sharp corners create high stress concentrations on the thermoplastic part; they are also stress concentrators within the mold cavity. These sharp corner areas fail under high loads. Internal radii of at least 0.031 in should replace sharp corners in thermoplastic part design wherever possible. If a sharp corner is unavoidable, reduce the radii and polish this surface area; in addition, these mold cavity areas should be designed with removable inserts to facilitate ease of repair.

10.2.4

Draft Walls

Thermoplastic parts should be designed with positive draft walls. Minimum positive draft is required on all walls in the direction of mold opening or core pulling. Without draft, thermoplastic molded parts adhere to the mold cavity surface, causing drag marks and surface ﬁnishing defects. In many cases, the part will not be fully ejected so that the mold may close on it and cause damage. Lack of positive draft also increases cycle time and molding costs.

10.2.5

Feather Edges

Avoid the use of feather-shaped edges that require thin and fragile steel. Within the mold cavity, feather edges tend to break and chip, resulting in mold maintenance and downtime. Undetected broken and chipped feather edges will cause ﬂashing problems as the thermoplastic melt ﬁlls into the mold vents. Feather edges become extremely hot and take longer to cool because cooling

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10 Injection Mold Design water channels cannot be brought to the feather edge, thus increasing cycle time.

10.2.6

Proportional Boss Geometries

Avoid the use of long narrow cores. The height of the unsupported core should not exceed four times the core base thickness. During the molding process, the injection pressure will deﬂect long narrow cores, because they act as cantilever beams, causing parting line openings and possible early failure of the mold core insert. In critical cases, a structural analysis of the mold can be made based on expected forces and allowable deﬂection. Cores of greater height must be fully supported using core inserts to decrease the chance of failure and to ease repair.

10.2.7

Gate Type and Location

The gate is an important component in the injection molding process. The gate inﬂuences the type of mold needed for the application (two-plate, three-plate, hot runner, and automation). The location of the gate determines the mold shrinkage, the melt ﬂow, part dimension, warpage, and weld line strength. The gate functions as a thermo-valve between the runner and the cavity. The temperature is increased around the gate area by the melt injection speed, pressure, and temperature. The hot gate allows the melt to enter the cavity without shearing off the polymer; the gate cools off when the melt stops moving, closing the gate while the melt inside the cavity cools off under packing pressure.

10.2.8

Molded Product Ejection Surface Area

The mold ejection system automatically provides a uniform force to extract the molded product from the cavities. The ejection force breaks the vacuum between the internal surface wall of the part and the cavity core and ejects the parts from the mold. The ejection surface area of the product should be located in the direction of the moving half of the mold, where the ejector system is generally placed. The product designer should specify large ejection surface areas with heavy cross sections that are not critical for the functionality of the product. The type of ejection system depends on the molded product’s geometrical conﬁguration and the permissible ejection surface area wall thickness, the stiffness, crystallinity rate, and melt temperature of the thermoplastic resin. For example, punctuation holes or indentation defect marks on the external surface of a molded product may be produced by small diameter ejector pins pushing against a ﬂexible and thin-walled cross section of a thermoplastic molded product during the ejection cycle.

10.2.9

Molded Product Tolerances

A realistic view of the cost of tolerances often helps avoid high molding costs without affecting the performance of the part. It may be unreasonable to specify close production tolerances on a part when it is designed to operate within a wide range of environmental conditions. Temperature-induced dimensional changes alone can be three to four times as great as the speciﬁed tolerances.

10.3 Effects of Mold Design on the Injection Molding Process The tolerances for injection molded thermoplastic parts have been developed by the plastic molding industry. The purpose of these speciﬁcations is to assist the part designer in obtaining a quick and preliminary analysis for the different molding tolerance factors found in a generic injection molded thermoplastic part. Understanding the limitations of this process and knowing how to control ﬁne tolerances are the result of applying thermoplastic part design principles, working with molds and mold designers, being aware of the governing rules in polymer technology, and applying the latest technologies available for the injection molding process.

10.2.10 Surface Finish of Molded Product Surface ﬁnish affects part quality, mold cost, mold cycle, and delivery time. Surface ﬁnishing is used to enhance surface clarity for appearance of the molded product. The standard steel ﬁnishes range from a number one (mirror ﬁnish) to a number six (grit blast ﬁnish). Any ﬁnish speciﬁcations on the part print must reference the molded product and not the mold itself. Specifying the mold surface ﬁnish does not necessarily produce the expected result on the ﬁnished molded product. Although a requirement for a part with a high-gloss ﬁnish requires a high-gloss ﬁnish on the mold cavity, other factors, such as resin, gating, melt and mold temperature, injection speed, and mold venting affect the surface ﬁnishing of the part. For extremely high-gloss ﬁnishing, the types of steel used in the cavities may need to be speciﬁed to ensure reasonable life of the polished cavity in production.

10.3

Effects of Mold Design on the Injection Molding Process

Mold design is an important consideration in the injection molding process, the following are general concepts for mold design, more detailed information is presented later in this chapter.

10.3.1

Runner System

Good runner design includes not only the correct geometry, size, and layout of the runner, but cooling, ejectability, and minimizing regrinds. A balanced runner system is required for ﬁlling all cavities at the same time; this minimizes cycle time and gives the best dimensional integrity to the molded product. Long and skinny or half-moon runners require higher injection pressures so the mold does not cool off prematurely causing incomplete parts. Long and thick runners increase the amount of regrinds, decreasing the efﬁciency of the molding process. Ejector pins should be provided at the intersection of cold runners to provide sufﬁcient force to eject the runner. Runners should be placed on the ejector half of the mold, so that they can be pushed out by the ejector system.

10.3.2

Mold Cooling System

Mold cooling is one of the most important parameters for controlling dimensional integrity, physical properties, surface ﬁnishing, warpage, weld line strength,

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10 Injection Mold Design and cycle time. A series of cooling channels along the length of a long cavity provides poor warpage control. All core pins should be cooled, especially if the projected length to core diameter ratio exceeds four. Hot core pins cause surface defects and longer molding cycles. A cooled pin is a more efﬁcient system for extracting heat from the encapsulated surface area. Using water in contact with pins is far better than transferring heat across an air layer. Flexible resins need the ejector pins to be cooled for ejection. Temperature control of the sprue puller pin area reduces mold cycle time and ejection interruptions. High volume and turbulent ﬂow of the cooling liquid is critical to maintain good temperature control in the mold. To control the corrosion inside the water lines stainless steel mold plates are used; Corrosion can also be controlled by plating the cooling channel or by the use of rust inhibitors in the water. The mold plates should be thick enough to provide room for the proper size of the cooling channels.

10.3.3

Ejector System

Uniform ejection is critical to control part warpage. Ejector pins, sleeves, rings, or plates must operate without obstructions. A guided ejector system allows ejector pins and cores to be precision aligned and will also bear the weight of the plates so that the pins do not wear and misalign. Early return systems should be provided as a safety feature. They drive all ejector pins to their seated positions before the mold closes, which could collide with pins that did not fully retract. A protector pin should be provided on all molds where the ejector pins or sleeves are under any slides. This locks the ejection plate in its retracted position to prevent collisions between ejector components and slides. For ﬂexible, thinwalled, deep and box parts that are difﬁcult to eject, special ejection systems should be used.

10.3.4

Mold Venting

Mold venting of cavities and runner systems to remove the entrapped air and polymer melt volatiles is a must. A mold without venting causes many processing problems.

10.3.5

Other Mold Devices

Other mold design devices can also increase the cost of the mold, as well as create difﬁculty in injection molding the thermoplastic product. One example is the mechanically activated side cores or slides. Although sometimes necessary, side cores are expensive mold components and will have high maintenance costs.

10.4

Design Considerations for Injection Molds

The main parameters for the design of a thermoplastic injection mold are: type, size, number of cavities, tolerances, runner layout, gating, venting, parting line, ejection system, surface ﬁnishing, steel hardness, and mold cooling among others. The geometrical design of an injection mold, the type of resin, the dimensional tolerances of the product, the part quality, and the volume of part production inﬂuence the selection of the steel for the mold because of considerations involving cavity forming and difﬁculties in heat treatment. Too great a difference in the wall thicknesses of mold cavity inserts necessitates greater care during heat

10.4 Design Considerations for Injection Molds treatment, because the time to heat the thickest wall section uniformly may result in over heating the thinnest wall section. Quenching from high temperatures, cracking, or distortion may occur where a thin section adjoins a thick section. Injection mold design is also important from the standpoint of service performance. If mold walls are made too thin, excessive elastic deﬂection and even cracking can result from service stresses. Overcoming excessive mold deﬂection that results in ﬂashing problems, dimensional control problems, poor part surface ﬁnishing, and incomplete molded parts can only be accomplished by increasing the wall section size of the mold. It is of extreme importance that any injection mold be designed and built by a qualiﬁed mold maker to the exacting standards demanded by the industry. Typical mold development procedures are as follows: • The product design has been completed and approved for molding • Mold planning process, where the product designer, tooling engineer, process engineer, mold designer, mold builder, and resin supplier’s technical representative review the needs of the product and provide recommendations for the design and construction of the mold • Development of a preliminary mold design proposal is the responsibility of the mold designer • Review of the mold layout proposal by the product designer, tooling engineer, process engineer, mold designer, and purchasing, recommending mold changes if needed and providing authorization to ﬁnalize the mold design • Completing the mold design details • Reviewing the ﬁnal mold design by the product designer, tooling engineer, process engineer, mold designer, and mold builder • Construction of the mold • First molding run evaluation to produce samples at the mold maker’s shop, end user tool engineer and process engineer should be present. • Inspection and documentation of all product samples whether they meet print dimensions (end user) • Debugging of the mold if needed • Approval of molded products and pre-production run (mold maker’s shop) • Surface treatment and polishing of the cavities and cores.

10.4.1

Preliminary Mold Design

Preliminary mold design begins from the product dimensional drawing. This represents the part geometry and required dimensions after the thermoplastic injection molding process, when the molded part has completed the mold shrinkage process and has been stabilized at room temperature for at least 24 hours. The preliminary mold layout requirements are the following: • Mold base size, thickness of plates, and ejection traveling distance • Placement, number, and size of cavity insert blocks

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10 Injection Mold Design • Location, type, and size of water cooling systems • Runner layout and type of gates • Position and number of support pillars • Ejector pins sizes and locations • Parting lines layout • Position of injection machine tie bars • Injection machine stroke and shut height positions • Automation requirements • Hot runner mold feasibility After a review of the preliminary mold layout proposal, authorization is given for detailed mold design.

10.4.2

Detailed Mold Design

The cavity details are designed in accordance with the part geometry and the thermoplastic injection mold design, based on good mold manufacturing practices. The critical ﬁnal mold design requirements include: general mold layout, runner and gate system, cooling system, ejector system, cavity ﬁnishing and venting. General Mold Layout • The mold base speciﬁcations for optimizing the injection molding machine cost rates and cycle times • Weak, delicate cavities that wear should be designed with a cavity block for strength and cooling using segmented cavity inserts so that mold maintenance costs can be optimized • In a multi-cavity mold, the cavities should be positioned to balance the system, reducing the runner length. • Allow between 0.003 and 0.005 in projected height beyond the cavity plates on each cavity half, so that good contact of the cavity halves can be made under clamping pressure. • Eye bolt holes should be provided on both the top and back side of each half of the mold to ease mounting the mold into the injection molding machine • The appropriate type steels should be used for each mold component based on the function, machinability, dimensional stability, tolerances, wear resistance, toughness, and strength required for the mold • Mold components such as sprue bushing, core pins, inserts, and cavities should be keyed to prevent incorrect placement in the mold; movement of these components could cause problems in the process. • Mold mounting/removal should be easy so that the mold setup time can be minimized • Cavity inserts should be thick enough with smooth radii so that they are strong enough to survive during molding production

10.5 Types of Steels Required for Injection Molds • One leader pin and one return pin on the ejector system should be offset to prevent a 180° reversal of the mold halves’ orientation • Mold plates should be strong enough and thick enough for the cavity insert block size and cooling system • Sufﬁcient support pillars should not interfere with the ejector pins. Location under the cavity is the primary position to avoid deﬂections of the cavity during the injection of the melt inside the cavity • The ejector system should be designed so that travel is sufﬁcient. Too much travel increases the cycle time, wear of the ejector pins and sleeves. Sufﬁcient travel is such that the part can fall free from the mold • The mold must be vented properly to allow entrapped air and melt volatiles to escape during injection. Otherwise, the steel of the cavity’s deep pockets becomes corroded. The vent channels should be properly located to optimize venting without allowing ﬂash. The dimensions of the vents are determined by the viscosity of the thermoplastic material used in the molding process • The ﬁxed half side of the mold (top clamping plate and “A” cavity plate) need to be thick enough to support leader pins and to provide room for the water cooling lines and other needs.

10.5

Types of Steels Required for Injection Molds

This section provides information on different steel compositions available to manufacture molds used to produce thermoplastic injection molded products. These steels constitute the most commonly used types of materials in the construction mold industry. Other materials such as beryllium-copper, cast aluminum alloy, forged aluminum alloy, cobalt-nickel alloy, and kirksite are also used, but to a lesser extent. Steels are the workhorse of materials used in molds. No other materials offer comparable versatility for product applications. Steels are produced in the greatest variety of forms and ﬁnishes, have strengths ranging from 30,000 psi to over 300,000 psi, and can withstand a range of temperatures from cryogenic up to 2,000 °F.

10.5.1

Major Steel Families

Because of the great range of steel types, properties and applications, steels are categorized into many families based on the chemical composition, heat treatment, surface ﬁnishing, critical properties (mechanical, thermal, corrosion resistance, electrical, etc.), typical processing characteristics, end use applications, and other factors. The major families of steel are the following: Low Carbon Steels SAE 1008, 1010, 1015, and 1025 are the lowest carbon steels selected when cold forming ability is the primary prerequisite. These steels have relatively low tensile strength values. Strength and hardness increase with carbon addition and/or with cold work, but a decrease in toughness or the ability to withstand

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10 Injection Mold Design cold deformation is created. Low carbon steels are nearly pure iron in structure, they are readily welded, but do not machine freely, producing poor smooth ﬁnishes. SAE 1016 to 1025 provide increased strength and hardness and reduced cold forming ability. Carburizing or case hardening is possible in some grades. Increase in carbon gives greater core hardness in thicker sections. Increase in manganese improves the hardening ability and also improves machining. SAE 1025 is used for larger sections or where greater core hardness is needed. All of these steels may be readily welded or brazed. SAE 1020 is frequently used for welded tubing. These steels are used for forged parts; they usually machine better in the “as forged” condition without annealing, or after normalizing. Medium Carbon Steels The medium carbon steels, SAE 1030 to 1052, are selected for uses where higher mechanical properties are needed and frequent further hardening and strengthening by heat treatment or by cold work is done. The carbon and manganese levels selected increase the mechanical properties required in section thickness or in depth of hardening. The heat treatment preferred for any of the grades over 0.30% carbon allows selective hardening by induction or ﬂame methods. All of these groups of steels are used for forging, the section being governed by the section size and the physical properties desired after heat treatment. Medium carbon steels are popular for forging and general uses requiring greater strength than low carbon steels. High Carbon Steels Steels SAE 1055 to 1095 are the high carbon types, having more carbon than is required to achieve maximum as quenched hardness. They are used for applications where the higher carbon is needed to improve wear resistance, higher strength characteristics for cutting edges, for springs, and for special purposes. Selection of a particular grade is affected by the nature of the part, its end use, and the manufacturing methods available. Cold forming is not always suitable and most parts are heat treated before use. Free Machining Carbon Steels Low and medium carbon steels, in which sulfur, sulfur-phosphorus combinations, and/or lead are purposely added to improve machinability, are termed free-machining carbon steels. Their designations are SAE 1108 to 1151 for resulfurized grades and SAE 1211 to 1215 for rephosphorized and resulfurized grades. Leaded grades are indicated by the letter “L” with the number. Sulfur and phosphorus additions result in some sacriﬁce in cold forming ability, weldability, and forging ability. Lead additions have little effect on forming ability and forging ability, but impair weldability. Carburizing Carbon Steels This term is sometimes given to standard carbon steels, primarily low carbon grades, which are case hardened by various carburizing methods.

10.5 Types of Steels Required for Injection Molds Hardening Carbon Steels These heat treating grades are medium and high carbon steels, hardened by heat treatment before use. They also are referred to as water or oil hardening grades, depending on the quenching media used during heat treatment. Carbon Spring Steels This is a common term for medium and high carbon steels (SAE 1050 to 1095) used in spring applications. Annealed and pretempered strips and wires are the common conditions and forms. These steels also are used for music and piano wire, rope wire, and saw blades. Low Temperature Carbon Steels These are low carbon (0.20 to 0.30%), high manganese (0.70 to 1.60%), silicon (0.15 to 0.60%) steels produced to a ﬁne grained structure with uniform carbide dispersion achieved through careful composition control and heat treatment. The steels feature relatively high strength and toughness combinations with ductility transition temperatures as low as 130 °F. High Strength Low Alloy Steels These are low carbon, manganese steels containing alloying elements, such as chromium, columbium, copper, molybdenum, nickel, titanium, vanadium, and zirconium. Their mechanical properties and corrosion resistance are superior to the more widely used structural carbon steels. These steels usually are used without heat treatment although annealing, normalizing, and stress relieving is required. Carburizing Alloy Steels These steels, which include all the low carbon standard alloy steels, are widely used for carburized parts. Low alloy grades (e.g., AISI 4023, 4118, 5015) are used for parts requiring better core properties than are obtainable from carburizing grades of carbon steels. The higher alloy grades (e.g., AISI 3120, 4320, 5120, 8620) are used for better case and core properties. Steel selection among the higher alloy grades depends primarily on hardening ability required to obtain the desired set of properties for the speciﬁc conditions of size and heat requirements. Alloy steels, direct hardening grades quenched and tempered to speciﬁc strength and toughness levels, include most of the standard AISI and SAE alloy steels (13XX, 31XX, 40XX, 41XX, 43XX, 46XX, 50XX, 51XX, 5XXX, 61XX, 86XX, 87XX, 92XX, 98XX); they are by far the most widely used alloy steels. They are also classiﬁed by SAE by the carbon content and as low, medium, and high hardening ability grades. Other designations include water or oil hardening types, or simply quenched and tempered steels. H-Alloy Steels They are standard direct hardening alloy steels that meet speciﬁc hardening ability limits as determined by end quench tests. The steels carry the letter “H” following their conventional numerical designations (e.g., AISI 3120H, 4340H, 8740H).

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10 Injection Mold Design Boron Alloy Steels Boron steels are direct hardening alloy steels containing very small amounts of boron to intensify hardening ability. They are also H-alloy steels and carry the letter “B” after the ﬁrst two digits of their designations (e.g., AISI 10B18, 50B46, 94B17). Nitriding Steels These are low and medium carbon/chromium, chromium/molybdenum, or chromium/aluminum steels that are case hardened by nitriding. They are referred to as low carbon, quenched and tempered, constructional alloy steels to distinguish them from the higher carbon quenched and tempered constructional alloy steels. They are noted for high strength, toughness, and weldability and for greater corrosion resistance than structural carbon steels. Their carbon content ranges from 0.10 to 0.20%, the typical alloying elements are chromium, molybdenum, nickel, copper, vanadium, titanium, boron, and zirconium. Low Temperature Alloy Steels These are primarily low carbon, nickel steels of relatively high strength and very high toughness at temperatures of 75 to 320 °F. The three most common grades have a maximum carbon content of 0.13 or 0.20% and nominal nickel contents of 2.25, 3.50 and 9%. Tool Steels Tool steels are medium and high carbon alloy steels noted primarily for their high hardness, abrasion resistance, and resistance to softening at elevated temperatures. Low carbon and relatively low alloy grades also are available. Tool steels are further classiﬁed as prehardened, carburizing, water, oil, and air hardening grades, hot and cold work grades, high speed grades, shock resisting grades, and by other designations. Despite their name, their primary use is for injection mold components and for non-tooling applications. Electrical Steels Electrical steels are alloy steels containing from 0.50 to 5.0% silicon and featuring relatively high permeability, high electrical resistance, and low hysteresis loss. They also are classiﬁed as grain-oriented and non-oriented types with the latter type further classiﬁed as low, medium, and high silicon grades. AISI designates the steels by the preﬁx letter “M” (for magnetic material) followed by a number that originally was approximately equal to 10 times the core loss in watts/seconds (for 29 gage sheet at 15 kilogausses and 60 cycles). Stainless Steels These are alloy steels that have superior corrosion resistance than carbon and conventional alloy steels as a result of moderate to high additions of chromium. Most grades also are noted for heat and oxidation resistance and some stainless steels have very high strength. These materials can be grouped into six major types: austenitic, martensitic, ferritic, age-hardenable austenitic, age-hardenable semi-austenitic, and age-hardenable martensitic. • Austenitic stainless steels 200 and 300 series designations are non-magnetic, non-hardenable by heat treatment

10.6 Steels for Thermoplastic Injection Molds • Ferritic stainless steels 400 series are magnetic and non-hardenable by heat treatment. • Martensitic stainless steels 400 and 500 series designations are magnetic, can be hardened by quenching and tempering. • Age-hardenable austenitic stainless steels (e.g., A-286) are strengthened by heating at moderate temperatures. • Semi-austenitic stainless steels (e.g., 17-7PH) are austenitic in the annealed condition and martensitic in the hardened condition. • Age-hardenable martensitic stainless steels (e.g., 17-4PH) are strengthened by aging reactions during tempering. High Temperature Steels Steels for high temperature service can be put into two categories: those primarily used for heat or oxidation resistance and those used for structural requirements. The most common heat resistant steels are the austenitic stainless steels that can be used up to about 2,000 °F. However, numerous other steels, such as the martensitic stainless steels, aluminized, and chromized carbon steels also fall into this category. High temperature (700 to 1,200 °F) structural steels include low alloy martensitic steels, Cr-Mo-V medium alloy air hardening steels, martensitic stainless steels, precipitation hardening stainless steels, and Cr-Ni-Mo alloy steels. Both categories of high temperature steels are classiﬁed under the AISI 600 series of high temperature and high strength alloys. Ultra High Strength Steels These are the highest strength steels produced. The designation is somewhat arbitrary, but generally refers to steels having yield strengths above 160,000 psi. The major types of ultra high strength steels are: medium carbon, low alloy, quenched, and tempered steels such as AISI 4130 and AISI 4340, 5Cr-Mo-V medium alloy air hardening steels, martensitic stainless steels, cold rolled austenitic stainless steels, precipitation hardening stainless steels, 12 and 18% nickel maraging steels, and 9Ni-4Co quenched and tempered steels.

10.6

Steels for Thermoplastic Injection Molds

The selection of a thermoplastic injection mold material depends on a number of factors, including such considerations as product design (geometry, part dimensional control, and tolerances), application end use requirements, thermoplastic material, number of mold cavities, expected production life of the mold, injection molding process conditions, and mold design and construction. The particular usefulness of steels is based on the unique combination of properties achievable through alloying (chemical composition) and uniform heat treatment process. Speciﬁc properties of steels that can be modiﬁed to achieve desirable characteristics are machinability, polishing ability, surface hardness and depth hardening, wear resistance, corrosion resistance, dimensional stability, thermal conductivity, mechanical strength, toughness, and cost. No other class of mold material is as versatile as steel.

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10 Injection Mold Design The main factors to be considered in selecting the optimum steels for making a thermoplastic injection mold are: • Product design geometry, tolerances, end use requirements • Molded product manufacturing costs • Type of thermoplastic (properties and process characteristics) • Injection molding process conditions • Number of molded parts to be produced with the mold • Surface ﬁnish of the molded parts • Number of cavities and mold size • Mold cooling and venting design • Type and size of runner layout (cold, balanced, or hot runner) • Geometry, location, and number of gates (two plate, three plate mold, or hot runner drops) • Cavity forming methods (machined, hobbing, EDM) composed of several assembled inserts or a single unit • Heat treatment method.

10.6.1

General Steel Selection Procedures

The following recommended tools are used to assist in the selection of optimum steels for thermoplastic injection mold components: • Material selection guide publication by the American Society for Metals (ASM) Metals Handbook, Volume 1, “Materials for Plastic Molds” • Technical information provided in this section, “Types of Steels Required for Injection Molds” • Comparative tested properties provided by steel suppliers design handbooks of candidate steel compositions • Mold designers’ and mold manufacturers’ recommendations of the types of steels with proven performance in similar injection mold applications and the years of experience working with these materials. The principal kinds of steel used in making thermoplastic injection molds are selected by their chemical composition, nitriding, prehardened, carbonizing, oil hardening, air hardening, and stainless properties. The principal methods used for forming cavities in steel molds are conventional machining, hobbing, and electrical discharge machining (EDM). Heat treatment is part of the mold making process, unless a prehardened type of steel is selected. Finishing of the mold surface cavity is usually accomplished by grinding and polishing. It is expected that the information provided in this section will help in selecting a suitable mold steel and mold making process for economic production of a thermoplastic injection molded part. Such decisions are generally considered to be the responsibility of mold designers, although they may receive assistance from others such as mold makers, product designers, process engineers, tool engineers, and resin supplier technical engineers.

10.6 Steels for Thermoplastic Injection Molds

10.6.2

Properties and Characteristics of Tool Steels

Uniformity of mechanical properties in the annealed as well as heat treated condition is considered important for tool steels. This requires attainment of a uniform micro structure throughout the mold component section after annealing and after heat treating, or, in a case hardened zone, after carbonizing and heat treating. Good machinability results from a uniform micro structure together with a relatively low hardness level in the annealed condition. There is a direct correlation between annealed hardness and hobbability. Working with soft steels, hobbability improves dramatically as hardness decreases. A uniform micro structure also facilitates machining steels in the prehardened condition. Low distortion in heat treatment is dependent on uniformity of the annealed micro structure, chemical composition, slow heating, and a controlled quenching rate. Good polishing ability is dependent on steel quality, hardness, and the type and uniformity of micro structure of the hardened cavity surface. Abrasion resistance, including the ability to retain a satisfactory polished surface under service conditions, depends on hardness level and micro structure. The presence of ﬁne excess alloy carbides in a high carbon martensitic matrix results in high abrasion resistance and good polishing ability. Higher corrosion resistance is attained with stainless mold steels or by chrome or nickel plating of the nonstainless steels. Commercially available high quality tool steel, alloy steels, and stainless are used in the construction of thermoplastic injection molds. Extensive precautions are generally taken in melting, processing, and inspection to prevent the occurrence of defective conditions. For example, using ultrasonic quality control inspections to detect defects is routine practice in the steel making process.

10.6.3

Effects of Alloying Elements on Tool Steel Properties

The steels used for thermoplastic injection molds range in composition from the relatively low alloy oil hardening types containing less than 0.50% total alloying elements through the medium alloy air hardening types up to the most highly alloyed steels that contain up to 21.00% of alloying elements by weight. Alloying elements are added to steels to achieve certain desirable properties or characteristics that would otherwise be unattainable. Some alloying elements, either alone or in combination, enable heat treatments to be carried out which alter the micro structure providing the desirable tool steel properties. The addition of more than one alloying element to a steel often produces a synergistic effect. Thus, the combined effects of two or more alloying elements may be greater than the sum of the individual effects of each alloying element.

10.6.4

Chemical Composition of Steels Used for Molds

Typical chemical compositions of commonly used steels used to fabricate the mold components are given in Table 10-1.

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10 Injection Mold Design Table 10-1 Chemical Composition of Steels Used for Molds

Steel Type

C

Mn

Si

Cr

Mo

V

P

S

Ni

Al

SAE-1015

0.15

0.30

–

–

–

–

0.04

0.05

–

–

SAE-1018

0.18

0.60

–

–

–

–

0.04

0.05

–

–

SAE-1020

0.20

0.90

0.20

–

–

–

0.04

0.05

–

–

SAE-1030

0.30

0.75

0.25

–

–

–

0.04

0.05

–

–

AISI-4130

0.30

0.48

0.20

0.95

0.20

–

0.15

0.15

–

–

AISI-4140

0.43

0.75

0.15

0.80

0.15

–

0.15

0.15

–

–

AISI-4150

0.50

0.90

0.30

0.95

0.20

–

–

–

–

–

SAE-6145

0.45

0.70

0.20

0.80

–

0.15

0.04

0.04

–

–

SAE-A2

1.00

0.70

0.30

5.00

1.00

–

–

–

–

–

SAE-A4

1.00

2.00

0.35

1.00

1.00

–

–

–

–

–

SAE-A6

0.70

2.00

0.30

1.00

1.25

–

–

–

–

–

SAE-D2

1.50

0.30

0.30 12.00 1.00

1.00

–

–

–

–

SAE-H13

0.35

0.40

1.10

5.00

1.50

1.10

–

–

–

–

SAE-O2

0.90

1.60

0.25

–

–

–

–

–

–

–

SAE-P2

0.07

0.30

0.15

2.00

0.20

–

–

–

0.50

–

SAE-P6

0.10

0.50

0.25

1.50

–

–

–

–

3.50

–

SAE-P20

0.35

0.90

0.50

1.70

0.40

–

–

–

–

–

SAE-P21

0.20

0.30

0.30

0.25

–

0.20

–

–

0.40

1.20

SAE-S7

0.50

0.70

0.25

3.25

1.40

–

–

–

–

–

Stainless 410

0.15

1.00

1.00 12.25

–

–

–

–

–

–

Stainless 420

0.15

1.00

1.00 13.00

–

–

–

–

–

–

Stainless 440C

1.05

1.00

1.00 17.00 0.75

–

–

–

–

–

10.6.5

Effects of Alloying on Tool Steels

The contribution of alloying elements on the mechanical properties and the characteristics of tool steels can be summarized as follows: Carbon (C) Carbon is the most influential element in controlling hardness, depth of hardening, and strength. Raising the carbon content by different amounts will increase the hardness, depth of hardness, strength, and abrasion resistance achievable after heat treatment and reduce ductility and toughness. Carbon combines with the carbide forming elements (Fe, Mn, Si, Cr, Mo, W, V, P, S, Ni and Al) to produce hard carbide particles that contribute signiﬁcantly to wear resistance. The amount of carbon in tool steels is speciﬁed for attaining certain properties (such as in the water hardening category, where higher carbon content may be chosen to improve wear resistance, although to the detriment of toughness) or, in the alloyed types of tool steels, in conformance with the other constituents for producing well balanced metallurgical and performance properties.

10.6 Steels for Thermoplastic Injection Molds Manganese (Mn) The principal function of manganese is to combine with free sulfur to form discrete sulﬁde inclusions and thus to improve hot working ability. Manganese is also a deoxidizing agent. In smaller amounts (0.60%), manganese is added for reducing brittleness and to improve forging ability. Larger amounts of manganese improve hardening ability, permitting oil quenching for non-alloyed carbon steels, thus reducing deformation and toughness. Molybdenum (Mo) Molybdenum is a strong promoter of certain metallurgical properties of alloy steels, such as deep hardening and toughness and is used for this purpose in most of the thermoplastic injection mold steels. Molybdenum contributes to secondary hardening on tempering when added in amounts from 0.20% to 1.50% or higher. Secondary hardening enables exceptionally high tempering temperatures to be used to obtain a given hardness and strength level. The higher tempering temperatures increase elevated temperature stability and strength and result in more complete relief of residual stresses for greater dimensional stability. It is used often in larger amounts in certain high speed tool steels to replace tungsten, primarily for economic reasons, often with nearly equivalent results. Vanadium (V) This element contributes to the reﬁnement of the carbide structure, improving the forging ability of alloy tool steels. Vanadium has a very strong tendency to form a hard carbide, which improves both the hardness and the wear properties of tool steels; however, a large amount of vanadium carbide makes grinding very difﬁcult.Vanadium is a relatively expensive alloying element and a strong carbide former that is usually added to control grain size and to increase wear resistance. Vanadium combined with carbon produces one of the hardest alloy steels. Aluminum (Al) Aluminum combines with nickel to form an intermetallic alloy steel where its hardness is controlled by the cooling rate. Aluminum is used to produce tool steels such as SAE-P21 (thermoplastic injection mold steel). Aluminum and manganese are the principal elements contributing to hardening ability so that full heat treated hardness can be obtained by air cooling. Distortion during heat treatment is lessened as the required quenching rate is reduced. Silicon (Si) Silicon is used in most tool steels in quantities from 0.15% to 1.10%. The principal function of silicon is as a deoxidizing agent during melting and to improve the hot forming properties of the steel. In combination with certain alloying elements, the silicon content is raised for increasing the strength and toughness of steels used for tools which have to sustain shock loads. Silicon increases hardening ability slightly and, in higher quantities, retards tempering reactions, thus allowing the use of higher tempering and operating temperatures. Chromium (Cr) Chromium is a carbide forming element with a dual function. When present as carbides, all or some may dissolve on hardening and reprecipitate on tempering, it contributes strongly to hardening ability, abrasion resistance, and toughness. When additional chromium, more than what can be utilized by the carbon to

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10 Injection Mold Design form carbides, is added, this chromium remains in solution and contributes to corrosion resistance. However, high chromium levels raise the hardening temperature of the tool steel, causing hardening deformation problems. A high percentage of chromium also affects the grinding ability of the tool steel. Nickel (Ni) Nickel in combination with other alloying elements is used to improve the toughness and the wear resistance of tool steels. Nickel is added to increase the hardening ability of alloy steels. In SAE-P21 tool steel, nickel combines with aluminum to form an intermetallic compound on aging to increase hardness and strength. Large amounts of nickel are needed to ensure the formation of martensite without carbon. Because of its tendency to promote higher annealed hardness and lower machinability, nickel improves aging, providing strength and hardness. Tungsten (W) Tungsten is one of the important alloying elements of tool steels. Tungsten improves the hot hardness, or the resistance to the softening effect at elevated temperature; tungsten forms hard, abrasion resistant carbides, thus improving the wear resistant properties of the tool steels. Cobalt (Co) A cobalt alloying element is used in applications where an increase in hot hardness is needed for the tool steels. Substantial addition of cobalt, however, raises the critical quenching temperature of the tool steel with a tendency to increase the decarburization of the surface and the reduction in toughness.

10.6.6

Effects of Heat Treatment on Tool Steel Properties

Heat treatment of thermoplastic injection molds is carried out to achieve satisfactory combinations of abrasion resistance, strength, and toughness in both carburizing and deep hardening types of mold steels. When a mold is to be heat treated, precautions should be taken to prevent surface carburization or decarburization as well as distortion and cracking. The thermoplastic injection mold cavities need to be heat treated by methods that introduce residual stresses; such a condition may result in distortion during subsequent heat treatment. Therefore, a stress relief treatment at about 1,200 °F is recommended if considerable machining has been done. To allow for distortion on subsequent heat treatment, thermoplastic injection mold components are generally rough machined within 0.125 to 0.25 in of ﬁnal dimensions before stress relieving. By a “trueing up” machining operation after the stress relief treatment, more reliable allowances for distortion can be made to limit the amount of grinding necessary after heat treatment. The typical methods for heat treatment of thermoplastic injection mold components are the following: Flame Hardening This is probably the oldest method used to increase wear resistance of the steel materials. AISI 4140 steel is usually purchased by screw manufacturers in the heat treated condition. Normally, this condition is 28–32 Rc, which provides

10.6 Steels for Thermoplastic Injection Molds good mechanical strength of approx. 100,000 psi tensile yield. The steel is still readily machinable in this condition. The mold component can be further ﬂame hardened to approx. 48–55 Rc. Of course, this cannot be done with low carbon steels. Approximately 0.40% carbon content is needed to achieve this result on a practical basis. The process employs an open gas/oxygen ﬂame followed by rapid quenching. The usual depth of hardness is approx. 0.125 in, but the hardness tapers off as you go deeper from the outside. Induction Hardening This process gives the same result as ﬂame hardening, but uses induction heat created by magnetic ﬂux reversals rather than a ﬂame. Nitriding A hard outside case can be obtained by subjecting the mold components to a high nitrogen atmosphere (ammonia gas) at elevated temperatures of approx. 950 °F. This comparatively low temperature causes minimal distortion but provides a very high case hardness of 60–70 Rc. The depth of case ranges from 0.020 to 0.024 in. This thin case diminishes in hardness from the outside, causing the mold components to wear rapidly once any signiﬁcant wear has occurred. Allowance for 0.0005 to 0.001 in of growth must be considered when designing mold components before the nitriding process. The high hardness is caused by the formation of metallic nitrides. A proper nitriding steel, such as Crucible Nitriding 135 or Ryerson Nitralloy 135M should be used to develop maximum hardness from the process. These steels are similar to AISI 4140 in their chemistry, but they have 0.95% to 1.30% aluminum added to form very hard aluminum nitrides. Alloy steels, such as AISI 4140 can be nitrided, but they exhibit a slightly lower hardness with a little increased depth of case. Nitriding is usually done over the entire mold component. This improves wear resistance to abrasion by ﬁber glass or abrasive mineral reinforced thermoplastic resins. Ion nitriding is a process superior to gas nitriding, which is described above. Ion nitriding is more expensive, but causes less distortion because of lower processing temperatures. The hardness, depth of case, and wear properties are very similar in either process. Precipitation Hardening Precipitation hardening is a low temperature process used to harden certain grades of stainless steels. 17-4PH stainless steel is an example of this type, where the PH stands for precipitation hardening. These grades are usually supplied in condition “A” (solution treated), which is similar to being annealed. The stainless steel component is then machined and hard-surfaced before precipitation hardening. Heat Treatment Process Rapid heating of a thermoplastic injection mold component to the austenitizing temperature and inadequate support of the mold component in the furnace can result in unacceptable distortion. Bolting the mold component to a support can prevent sagging. A support may also be necessary to prevent distortion of a large mold component during its removal from a furnace before quenching. Since the greatest shape distortion generally occurs during quenching, the objective should be to cool all sections of the mold component at the same rate. A mold

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10 Injection Mold Design component with uniform wall thicknesses on all sides will cool more uniformly than a mold component with both thick and thin walled cross sections. It is desirable to use a quenching rate that is not much faster than necessary to achieve the desired degree of hardening. Lower distortion can be achieved by air hardening or precipitation hardening of the steel. In carrying out heat treating operations, such as stress relieving and austenitizing, the holding time at temperature is determined by the maximum cross section wall thickness of the mold component. Tempering temperatures should be selected to achieve a satisfactory combination of hardness and toughness. Shape distortions on tempering can be minimized by heating the mold component slowly to the tempering temperature starting with a cold furnace, tempering as required by the desired properties and cross section wall thickness, followed by furnace cooling. The use of a protective furnace atmosphere, a vacuum furnace, or salt bath is generally necessary to prevent excessive oxidation and carburization or decarburization during austenitizing that could adversely affect the surface ﬁnish and surface properties. Deterioration and nonuniformity may also result from failure to clean all the oil or grease from a mold component surface before hardening by using carburizing compound for case hardening or by not removing any copper that may be on the surface.

10.6.7

Prehardened Tool Steels

The most commonly used prehardened steels for making thermoplastic injection mold components are SAE P20, P21 and other medium carbon low alloy grades of suitable quality. Prehardened SAE H13 tool steel is available in rounds from a limited number of suppliers. Stainless steel 420 is also available in the prehardened condition. The use of prehardened steels eliminates the possibility of cracking, distortion, scaling, and pitting that can occur during heat treatment of a mold component that has been machined in the annealed condition. Prehardened steels with hardnesses in the range of 28 to 32 Rc are used mainly for injection mold plates. The upper limit of this hardness range corresponds to about the upper limit of adequate machinability using conventional metal cutting tools. A stress relief at 100 °F below the tempering temperature of these steels is usually carried out after rough machining the mold cavity to avoid excessive residual stress that might result in distortion or possible cracking during subsequent grinding. Prehardened SAE P20 Tool Steel Prehardened SAE P20 tool steel is a 0.35% carbon, 1.70% chromium, 0.40% molybdenum, 0.90% manganese, and 0.50% silicon tool steel available in the heat treated condition at 28–32 Rc as a result of prior hardening and tempering. This steel has excellent dimensional stability, good machinability, polishing ability and toughness, but fair wear resistance. Eliminating a second machining operation normally required by steels that must be heat treated is considered suitable for various sizes of molding machines and types of injection molds. To obtain greater surface strength, carburizing is followed by conventional hardening and tempering treatments to obtain adequate hardness and strength. Prehardened SAE P21 Tool Steel Prehardened SAE P21 is mainly used for injection molds; its chemical composition is 0.20% carbon, 0.40% nickel, 0.25% chromium, 0.02% vanadium, 0.30%

10.6 Steels for Thermoplastic Injection Molds manganese, 0.30% silicon, and 1.20% aluminum; available in the precipitation hardened condition at 32–39 Rc as a result of prior solution treating and aging. This steel exhibits high strength, excellent dimensional stability, good machinability, fair toughness and wear resistance. Cavities and cores are cut directly from this steel. Uniformity of hardness and strength can be attained in a relatively large section, provided cooling is at a sufﬁcient rate from the solution treating temperature to prevent softening as a result of aging. Prehardened SAE H13 Tool Steel Prehardened SAE H13 steel is a 0.35% carbon, 0.40% manganese, 5.00% chromium, 1.50% molybdenum, 1.10% silicon, and 1.10% vanadium tool steel available in the heat treated condition at 30–36 Rc as a result of prior hardening and tempering (65–74 Rc) This steel has optimum dimensional stability, excellent toughness and wear resistance. This tool steel is primarily used for protection from abrasion, for injection molds requiring high thermal shock resistance, and good polishing ability. A drawback to this tool steel is the difﬁculty of machining or rebuilding by hard-surfaced welding. It is also difﬁcult to straighten by conventional methods A free machining grade of SAE H13 tool steel is produced by use of carefully controlled and evenly dispersed sulﬁde additions in the melting operation. Such a steel can be prehardened to 50–52 Rc and still possess adequate machinability. This has the advantage of higher strength although the polishing ability may be adversely affected by the increased sulﬁde content. Prehardened AISI 4130 Alloy Steel Medium carbon alloy steel AISI 4130 prehardened is a 0.30% carbon, 0.95% chromium, 0.48% manganese, 0.20% silicon, and 0.20% molybdenum alloy steel available in the heat treated condition at 28–32 Rc. The hardening ability is low and it is quenched in water. This steel has optimum dimensional stability and machinability, excellent toughness and adequate wear resistance. This steel is advantageous for injection molds because of high toughness, moderate strength, and high thermal shock resistance. Prehardened AISI 4140 Alloy Steel Medium carbon alloy steel AISI 4140 prehardened is a 0.40% carbon, 0.80% chromium, 0.75% manganese, 0.15% silicon, and 0.15% molybdenum alloy steel available in the heat treated condition at 28–32 Rc and stress relieved conditions. It has good strength and can be ﬂame hardened or rebuild by hardsurfaced welding. The hardening ability is medium and it is quenched in oil. This steel has optimum dimensional stability and machinability, good toughness, strength and wear resistance. It is important to make sure that special, easy to machine grades are not selected. These grades usually contain either lead or are resulphurized. This makes the mold component impossible to rebuild by hard-surfaced welding due to extreme porosity and stress cracks. Prehardened AISI 4150 Alloy Steel Medium carbon alloy steel AISI 4150 prehardened is a 0.50% carbon, 0.95% chromium, 0.90% manganese, 0.30% silicon, and 0.20% molybdenum alloy steel available in the heat treated condition at 30–36 Rc. The hardening ability is high

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10 Injection Mold Design and it is quenched in oil. This steel has high strength, very good dimensional stability and machinability, good wear resistance and moderate toughness. AISI 4340 Alloy Steel This alloy steel is similar to AISI 4140, but includes nickel as an alloying element plus a greater percentage of molybdenum. This provides a slightly higher strength, but the major difference is the greater penetration of heat treatment providing superior mechanical properties at the core of the mold component. Nitralloy 135M Alloy Steel A number of steel manufacturers produce a product similar or identical to Nitralloy 135M, but this is the name best known. Nitralloy is similar to AISI 4140, but is slightly lower in physical properties. The major difference is the inclusion of a small portion of aluminum. CPM-10V Tool Steel This material is a totally new concept for tool steels based on particle metallurgy. CPM stands for “Crucible Particle Metallurgy”. This process enables the incorporation of very high amounts of carbon, chromium, and vanadium not possible by conventional methods used to manufacture steels. CPM-10V steel is intermediate between tool steels and cemented carbides, yet it has excellent strength. It also has much higher toughness than conventional tool steels, if heat treated properly. It provides superb wear resistance in ﬁber glass reinforced thermoplastic processing components such as injection screws, check valves, barrels, nozzles, mold cavity inserts, and gate inserts. CPM-10V large bar stock is very expensive, but in smaller sizes (2.50 in) is fairly competitive. The material is probably the best choice for mold components subjected to more than 10% glass ﬁbers or abrasive mineral reinforced thermoplastic resins, especially if these polymers have high melt ﬂow rates.

10.6.8

Carburizing Tool Steels

The two most commonly used carburizing steels for making thermoplastic injection molds are SAE P2 and SAE P6. Hobbed or machined injection molds made from these steels are case hardened by a carburizing treatment followed by hardening and tempering treatments. The aim of case hardening is to attain an adequate combination of abrasion resistance and toughness for a particular application. Carburizing steels are used for injection molds when a higher hardness is required than can be achieved with prehardened steels. Selection of the carburized case depth depends on the service conditions to which the mold component will be subjected. The deeper and harder the case hardened zone, the greater the resistance to abrasion. The stronger and tougher the core zone, the shallower is the required case depth. However, a large cavity usually requires a greater case depth along with a tougher core than a small cavity. A greater case depth is also required if a mold is subjected to a high injection molding pressure or if a cavity is shallow compared to other dimensions. Carburizing SAE P2 Tool Steel SAE P2 steel is a 0.07% carbon, 2.00% chromium, 0.50% nickel, 0.15% silicon, and 0.30% molybdenum tool steel available in the annealed condition. This steel has fair dimensional stability, good toughness, and very good wear resistance.

10.6 Steels for Thermoplastic Injection Molds After a mold cavity is formed from SAE P2 tool steel by hobbing, the mold cavity is carburized to a case depth of 0.05 to 0.06 in and heat treated to a case hardness of 60–64 Rc with an internal core hardness of 14–18 Rc. Due to the high strength and abrasion resistance at the surface, carburized SAE P2 tool steel’s main advantages are that it can be easily cold hobbed in the annealed condition and it can subsequently be polished to a high luster in the carburized condition. Carburizing SAE P6 Tool Steel SAE P6 steel is a 0.10% carbon, 3.50% nickel, 1.50% chromium, 0.25% silicon, and 0.50% molybdenum tool steel available in the annealed condition. This steel has fair dimensional stability, good toughness and machinability, and very good wear resistance. After a mold cavity is formed from SAE P6 tool steel by machine cutting, the mold cavity is carburized and heat treated to a case hardness of 58–61 Rc with an internal core hardness of 26–28 Rc. SAE P6 tool steel is comparable to SAE P2 with regard to surface strength, abrasion resistance, and polishing ability. Due to the higher internal core hardness, SAE P6 mold cavities are capable of withstanding higher injection pressures than SAE P2 mold cavities.

10.6.9

Oil and Air Hardening Tool Steels

The most commonly used oil and air hardened tool steels for making thermoplastic injection mold components are SAE O2, S7, H13, A2, A4, A6, and D2. These tool steels offer suitable combinations of high abrasion resistance, high strength, and high polishing ability required for thermoplastic injection mold components. To achieve a mirror ﬁnish on the cavity surfaces it is necessary to attain a hardness greater than 54 Rc. Because of its lower hardness in the annealed condition, machinability is superior to that of the prehardened steels. Oil Hardening SAE O2 Tool Steel SAE O2 tool steel is a high carbon oil hardened tool steel available in the spheroidized/annealed condition. SAE O2 chemical composition is a 0.90% carbon, 1.60% manganese, and 0.25% silicone tool steel. It is capable of attaining surface hardness up to 63 Rc by heat treatment. This steel has optimum machinability, adequate dimensional stability, toughness, and wear resistance. Air Hardening SAE S7 Tool Steel SAE S7 tool steel is an air hardening 0.50% carbon, 0.70% manganese, 0.25% silicon, 3.25% chromium, and 1.40% molybdenum shock resisting tool steel that is available in the spheroidized/annealed condition. The suggested hardness range of 52–58 Rc can be attained by air hardening and tempering. This steel has excellent toughness, very good machinability, dimensional stability, and wear resistance. SAE S7 tool steel is suitable for injection mold components when high strength, toughness and thermal shock resistance are required. Air Hardening SAE A2, A4 and A6 Tool Steels These are commercially medium alloy air hardened tool steels in the spheroidized/annealed condition. Hardness in the range of 58–60 Rc can be attained by heat treatment. This hardness range is suitable for injection molds. SAE A2 steel is a 1.00% carbon, 0.70% manganese, 0.30% silicon, 5.00% chromium, and 1.00% molybdenum tool steel that has the highest abrasion resistance of the group.

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10 Injection Mold Design SAE A4 steel is a 1.00% carbon, 2.00% manganese, 0.35% silicon, 1.00% chromium, and 1.00% molybdenum produced as a free machining grade tool steel. SAE A4 tool steel has the advantage of requiring a lower hardening temperature than SAE A2. SAE A6 is a 0.70% carbon, 2.00% manganese, 0.30% silicon, 1.00% chromium, and 1.25% molybdenum produced as a free machining grade tool steel. SAE A6 tool steel has the higher toughness of the group and the advantage of requiring a lower hardening temperature than SAE A2. Air Hardening SAE D2 Tool Steel SAE D2 tool steel is a 1.50% carbon, 0.30% manganese, 12.00% chromium, 1.00% molybdenum, 0.30% silicon, and 1.00% vanadium air hardened tool steel available in the spheroidized/annealed condition. A hardness of 58–60 Rc can be achieved by heat treatment. This tool steel has excellent dimensional stability and wear resistance, but poor toughness. SAE D2 tool steel is the most abrasion resistant of the deep hardening steels listed in Table 10-1. It has been primarily used for applications requiring abrasion and corrosion resistance. A drawback to this tool steel is the difﬁculty of machining or rebuilding by hard-surfaced welding; it is difﬁcult to straighten by conventional methods.

10.6.10 Stainless Steels Stainless steel types 410, 420 and 440C are martensitic stainless steels commonly used for thermoplastic injection molds. Their high corrosion resistance is of advantage in molding relatively corrosive thermoplastics or molding under relatively corrosive atmospheric conditions. Stainless Steel 410 Stainless steel 410 is a 0.15% carbon, 1.00% manganese, 1.00% silicon, and 12.25% chromium steel available in the annealed condition. Suitable heat treatment can produce a hardness of 38–41 Rc. This stainless steel has excellent polishing ability eliminating the need for chrome plating. The toughness is excellent, dimensional stability and wear resistance properties are good, but the machinability is fair. It has slightly better corrosion resistance than stainless steel 420. Stainless Steel 420 Stainless steel 420 is a 0.15% carbon, 1.00% manganese, 1.00% silicon, and 13.00% chromium steel available in the annealed condition. By heat treatment, a hardness of 50–54 Rc is attained, which results in higher strength and abrasion resistance than stainless steel 410. This stainless steel has excellent polishing ability, eliminating the need for chrome plating. The toughness, dimensional stability, and wear resistance properties are good, but the machinability is fair. Stainless Steel 440C Stainless steel 440C is a 1.05% carbon, 1.00% manganese, 1.00% silicon, 17.00% chromium, and 0.75% molybdenum steel available in the annealed condition. By heat treatment, a hardness of 56–61 Rc can be attained, which results in higher strength and abrasion resistance than stainless steel 410 and 420. This stainless steel has excellent polishing ability, eliminating the need for chrome plating. The toughness is poor, dimensional stability and wear resistance properties are very good, but the machinability is fair.

569

10.6 Steels for Thermoplastic Injection Molds

10.6.11 Steels Used in Thermoplastic Injection Mold Components Prehardened, carburizing, and deep hardening steels are all used for general purpose thermoplastic injection molds because of their service record, cost, and ready availability. Prehardened steels of moderate strength and wear resistance are widely used, but as operating conditions become more severe, high alloy tool steels become more common. Carburized tool steels are used for high ﬂow, ﬁber glass, and abrasive mineral reinforced injection molding thermoplastics. Heat resistant steels are used for injection molding thermoplastics requiring high mold temperatures. Stainless steels, as well as chrome and nickel plated prehardened or deep hardened steels, are used for injection molding corrosive thermoplastic resins. The types of steels used in the construction of the mold components can be divided into two groups, typical commercial mold base combinations of steels as shown in Figure 10-2 and the steels used for the remaining mold components. The hardness of a mold cavity is usually in the range between 50 and 60 Rc and it may be chromium- or nickel-plated with plating layers usually under 0.002 to 0.005 in thick. Care must be taken when chromium-plating. The plated components must be treated to prevent hydrogen embrittlement where the hydrogen gas generated in plating can cause loss of steel toughness.

Steel “A” Steel “A”

Steel “A” Steel “A”

Steel “A” Steel “A” Economical grade

Steel “B” Steel “B”

The steels listed in Table 10-2 can be used for making thermoplastic injection mold components. The selection of steels depends on the mold component’s operating functions, type of thermoplastic resin to be molded, number of parts to be produced, and the method of forming the mold cavity.

Steel “B” Steel “B”

There are other materials used in mold components, depending on the application:

Steel “A”

Aluminum is used for prototype molds. The types used are 6061 and 7075 series, with tensile strengths over 75,000 psi and relatively good scratch resistance. Copper, aluminum, or beryllium copper alloys are used in molds, they have high thermal conductivity and can be hardened to 45 Rc. Their problems are poor machinability; they are expensive and have low strength and stiffness, so core deﬂection can be a problem. Bronze is used in some cases for wear resistance in mold mechanisms and occasionally for cavity inserts.

Steel “A” General grade

Steel “B” Steel “C”

Steel “C” Steel “B”

Steel “A” Steel “A” Performance grade

Steel “A”: SAE 1015, 1018, 1020, 1030. Low carbon hot rolled steel with good tensile strength properties, machines easily, allowing faster and smoother cuts. Steel “B”: AISI 4130, 4140, 4150. Preheat treated medium carbon alloy steels (28–36 Rc), high strength, machinability, and toughness, with good wear resistance properties. Steel “C”: SAE P20, P21. Medium hardening ability preheat treated tool steels (28–39 Rc), excellent dimensional stability and surface ﬁnishing, high strength, good machinability and toughness, but low wear resistance properties. Steel “D”: SAE A4, A6, O2, S7, H13, P6, P20. Preheat treated medium carbon tool steels, chrome and nickle plated, excellent dimensional stability and surface ﬁnishing, high strength. Steel “E”: Stainless Steel 410, 420, 440C High corrosion resistance by heat treatment to 38–54 Rc, high strength, surface ﬁnishing, and wear resistance, but hard to machine

Steel “D” Steel “E”

Steel “E” Steel “D”

Steel “D” Steel “D” Corrosion resistance

Figure 10-2 Typical commercial mold base combinations of steels

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10 Injection Mold Design Table 10-2 Steels Used in Thermoplastic Injection Mold Components

Mold components

Types of steels used

Locating ring

SAE 6521; 6524 (28–32 Rc)

Extension nozzle bushing

AISI 4140 (28–32 Rc)

Sprue bushing

SAE 6145; Ampco 940/stainless steel

Sprue puller pin

Nitrided H13 (65–74 Rc)

Ejector pin

Nitrided H13 (65–74 Rc); CM 50 (60–65 Rc)

Ejector sleeve

Nitrided H13 (65–74 Rc)

Ejector plate

SAE 1020 (30–35 Rc)

Ejector retainer

SAE 1020 (30–35 Rc)

Leader pin

SAE 1018; 8020 (30–35 Rc)

Return pin

Nitrided H13 (65–74 Rc)

Ejector stripper plate

Prehardened P20; H13; S7; A2; A4; A6

Cavity block

Prehardened or precipitation hardening P20; P21; H13; S7; A2; A4; A6; P2; P6; Stainless steel 420 (28–50 Rc)

Cavity insert

Prehardened or precipitation hardening H13; D2; S7; A2; A4; A6; P2; P6; stainless steel 420; 440C (50–64 Rc)

Gate insert

D2; CPM 10V; CPM 9V (50–60 Rc)

Core block

Prehardened or precipitation hardening P20; P21; H13; S7; A2; A4; A6; P2; P6; stainless steel 420 (28–50 Rc)

Core insert

Prehardened or precipitation hardening H13; S7; P2; stainless steel 420, 440C; beryllium copper; Ampco 940

Slide

Carburized or nitrided P20; nitrided P21; O2; A2; A6; P2

Wear plate

SAE A2; A6; P2 bronze-plated

Angle pin

Nitrided H13 (65–74 Rc)

Knockout rod

SAE 1020 (30–35 Rc)

Support pillar

SAE 1040 (28–32 Rc)

Stop pin

SAE 1040 (28–32 Rc)

Parting line interlock (male)

SAE 8620 (50–55 Rc)

Parting line interlock (female)

SAE 8620 (55–60 Rc)

Insulator sheet

Asbestos-free glass reinforced polymer composite

Zinc alloys are too soft and not strong enough for production molds, unless the injection pressure and melt temperature are very low, and the molded product does not require good dimensional tolerances. “Kirksite” is used for prototype molds, but “Kirksite” has higher thermal conductivity than steel and will produce molded parts with different shrinkage than parts from a production steel mold. Electro-deposited cavities are used for prototype molds and for small specialized cavities used in production steel molds.

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10.7 Mold Cavity Surface Finishing

10.7

Mold Cavity Surface Finishing

The cavity surface ﬁnishing, runner and gating system, mold temperature, the type of thermoplastic resin, and the process molding conditions will determine the quality surface ﬁnish that will be obtained on the molded thermoplastic parts. The type of tool steel, its quality, structure, and heat treatment process can also affect the polishing ability of the cavity and the surface ﬁnish obtained. Several new technical innovations or improvements by various leading steel mills are producing better quality tool steels as a result of alloying compositions, modern manufacturing processes, and metallurgical quality control. The product design surface ﬁnish speciﬁcations may require the use of a speciﬁc type of thermoplastic resin. The product surface ﬁnish must be studied in great detail before proceeding with the design and manufacture of the mold. The difference between a number three cavity surface ﬁnishing (good mold ﬁnish) and a number one surface ﬁnishing of optical quality (mirror ﬁnish) can double the price of the mold. When the mold design is more complicated, the tool steel for the mold cavity is more expensive, and the heat treating method is more difﬁcult; machining, grinding, and polishing represent the effort of many man hours reﬂecting the price increase. Processing with a mirror ﬁnish mold requires a better quality thermoplastic resin that is more expensive. The material handling efﬁciency for the resin is low, the resin must be dry before molding and the use of reground material (runners and rejected parts) is not allowed. The molding process efﬁciency is also low with long molding cycles and high rejection rates, which require extra quality control support. High maintenance costs, spare parts for inventory, and mold cavity insert replacements, requiring special equipment for repairs, which are also time consuming. The life expectancy of such a mold is typically low because of complications and a new spare mold will probably be needed for the molding process. The surface ﬁnish characteristics of the thermoplastic resins are important parameters for the injection molding process. Unreinforced amorphous resins have better surface ﬁnish characteristics than glass- or mineral-reinforced semicrystalline resins. For example, injection molded products (lenses, tail lights, or dial faces) from transparent materials such as acrylic and PC will require the highest quality ﬁnish. Other materials, such as low density polyethylene or high impact polystyrene, which have a natural shine, do not reproduce the high level ﬁnish that is generated with a diamond polished surface (No.1 surface ﬁnish). The Society of the Plastics Industry and the Society of Plastics Engineers developed mold steel surface ﬁnishing standards, which provide an excellent visual comparison. The standards consist of a group of six pieces of steel, each ﬁnished to a different level and used as a comparison guide in mold ﬁnishing. Finish

Process parameters

Number 1

8,000 Grit (0.0 to 3.0 diamond compound micro range)

Number 2

1,200 Grit (up to 15.0 diamond compound micro range)

Number 3

320 Grit, abrasive cloth

Number 4

280 Grit, abrasive stone

Number 5

240 Grit, dry blast (5.0 in distance at 100.0 psi pressure)

Number 6

24 Grit, dry blast (3.0 in distance at 100.0 psi pressure)

Number 4

Number 5

Number 6

Number 3 Number 1

Number 2

Figure 10-3 SPI-SPE mold steel surface ﬁnishing comparison kit

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10 Injection Mold Design The grades range from number six, which is a 24 grit dry blast ﬁnish, to number one, which reﬂects a diamond polish with an 8,000 grit compound. The ofﬁcial SPI-SPE mold steel surface ﬁnish standard comparison kit and the descriptions are shown in Figure 10-3. There are other, more sophisticated, methods of measuring surface characteristics. One such method is the use of an instrument called a proﬁlometer. It measures the surface “roughness” of the metal. The stylus of this instrument is moved over the surface of the steel; the peaks and valleys of the surface ﬁnish are fed into the analyzer and the surface roughness is charted in terms of micro inches. By deﬁnition, a micro inch is a millionth of an inch (0.000001 in). The surface of any material is not truly “smooth” when examined under the microscope, but rather generally covered with ridges and scratches. The proﬁlometer measures the depth of the scratches and the height of the ridges and reports them as maximum peak-to-valley height, along with the average deviation from the mean surface. The average deviation from the mean surface is called the “Root Mean Square” and is known as RMS. A micro inch is the most often used value for specifying surface ﬁnishes. Mold steel surface ﬁnishing comparison standards have established the mold steel surface ﬁnish assigned number versus the RMS relative value in micro inches. Table 10-3 shows the equivalent RMS values for the SPI-SPE mold steel surface ﬁnishes.

Table 10-3 Equivalent RMS Values of Standard Mold Steel Surface Finish

Mold ﬁnish

RMS value range

Mold ﬁnish

RMS value range

Number 1

0.50 to 1.00

Number 4

12.00 to 15.00

Number 2

1.00 to 2.00

Number 5

26.00 to 32.00

Number 3

7.00 to 7.50

Number 6

160.00 to 190.00

Table 10.4 provides a reference for the ﬁnishing process required to produce the desired RMS surface ﬁnish classiﬁcation.

Table 10-4 RMS Surface Classiﬁcations Based on the Finishing Process Used

RMS surface range

Finishing process parameters

0.00 to 10.00 RMS

Fine abrasive hand polish, coarse to ﬁne diamond polish, ultra-ﬁne hand-rubbed diamond ﬁnish.

10.00 to 20.00 RMS

Fine cylindrical grind, smooth ream, ﬁne surface grind, medium grind hand polish.

20.00 to 40.00 RMS

Fine machine grind, medium surface grind, rough abrasive polish, ream.

40.00 to 100.00 RMS

High grade machine ﬁnish, coarse surface grind, smooth cylindrical grind, smooth hand ﬁle.

100.00 to 250.00 RMS

Medium machine cuts, coarse surface grind, smooth disc ﬁnish grind, medium ﬁle.

250.00 to 500.00 RMS

Heavy machine cuts, rough ﬁling, rough disc grinding, sand cast surfaces.

10.7 Mold Cavity Surface Finishing

10.7.1

Mold Surface Finishing Process Procedures

The various surface textures for mold cavity and core surface ﬁnishing levels required in the thermoplastics injection molding industry are more demanding than the conventional molds built for other plastic processes, such as blow molding or thermoforming. The mold cavity surfaces must be prepared to accept the ﬁnal polish, whether it is a 320 grit cloth ﬁnish or a number one (1) diamond polish mirror ﬁnish. The polished surface must be free of waves, ripples, distortions, pits, scratches, or orange peel. Machining Machining is the ﬁrst step in the manufacturing process of a mold component. It determines the tools to be used, the surface condition obtained during the machining operation, and the methods of polishing and ﬁnishing. When the mold components are machined according to the size and geometry of the mold design that has already accounted for the mold shrinkage of the thermoplastic part, the cavity dimensions are left with about 0.001 to 0.003 in of stock for subsequent ﬁnishing operations. When the mold cavity and core inserts are to be heat treated, the details are machined to within 0.0003 to 0.0005 in of the ﬁnal surface ﬁnish. Mold components that come off the duplicating machine will have the characteristic ridges and valleys resulting from this particular type of machining. The ridges on either side of the valley must be leveled to expose an even surface for the ﬁnal ﬁnishing. The metal removal operation can be aided by the use of a metal dye painted over the surface to serve as a guide in metal removal. It is often advisable to machine cavities between 0.002 and 0.003 in deeper than necessary and remove the proper amount from the top surface, after the mold has been tested in the injection molding machine. Electrical Discharge Machining (EDM) Electrical erosion is a metal removal process using a master electrode that is electrically conductive. Copper alloys are generally used to make the master electrode because hardness is not a requirement. Cast zinc and machined graphite are also used in some instances. The principle of spark erosion is used by the mold making industry. The gap between the master and the cavity insert is quite uniform and small. As the master descends, small intense sparks are generated wherever the gap is reduced. Erosion occurs on both master and cavity inserts, the master’s negative polarity erodes only at 1/4 to 1/10 of the speed at which the cavity insert erodes at positive polarity. The cavity insert may be hardened before the EDM process begins so that distortion due to heat treatment is eliminated. The dielectric ﬂuid must circulate continuously to remove the minute particles that are formed between the master and the cavity insert. Electrical erosion is slow compared to mechanical cutting of soft steels, but for certain conditions, such as narrow deep slots, it offers great advantages. To utilize the EDM process, the mold maker roughs out stock by machining wherever possible and then sends the cavity insert to the EDM processor. Grinding To remove ridges and rough cut marks, a hand or ﬂexible shaft grinder is ﬁtted with a grinding wheel or abrasive disc. Extreme care must be exercised to prevent the grinder from following the ridges and removing more material than necessary.

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10 Injection Mold Design Grinding strokes should follow the crest of the ridges until the mean surface is reasonably constant. A more recently developed tool, the portable belt sander, can also be used in this “roughing” phase of the operation. Discs or abrasive belts have the advantage of being able to span a greater area, but care must be taken because they can also cut very quickly. Grits in the range between 60 and 320 are used. Filing provides ample control and keeps the surface quite even. The grade of ﬁle used will be determined by the amount of metal to be removed. Rough grinding, or coarse ﬁling, can leave the metal surface full of tears and waves. Finer wheels, drums, and ﬁles may be used to improve the surface texture and bring the cavity or core closer to print dimensions. After all the metal has been removed to make the part to print (plus the allowance for ﬁnishing stock), the surface should be examined to determine if it is ready for stoning. The surfaces of heat treated thermoplastic injection molds are ground before polishing. Improper grinding can generate very high temperatures and precautions should be taken to avoid the following defective surface conditions: • Softening from over tempering • Embrittlement due to rehardening • Distortion and unfavorable residual stress • Cracking Avoiding these problems requires grinding with a proper wheel, at its recommended rotational speed, at an adequately low rate of metal removal, and with an adequate coolant. If residual stresses imposed by grinding are not relieved, pitting during subsequent polishing may result. Adequate relief of residual grinding stresses can usually be accomplished by heating to about 100 °F below the tempering temperature. Stoning Choosing the initial grit of stone depends on the degree of ﬁnish left by the machining, grinding, or ﬁling operation. Machining or duplicating usually results in a coarser ﬁnish than grinding; therefore, a coarser grit stone would normally be used. Preliminary stoning may be done with a 240 grit stone to remove ﬁnal dips, depressions, waves, or other imperfections to obtain a ﬂat or properly contoured surface. If defects are not too great, a 320 grit stone will be sufﬁcient. The stone should be moved back and forth (medium pressure applied), over the surface in a 90° direction from the direction made by the last operation. Before they are used, the stones should be soaked in a contaminantfree oil base lubricant. The recommended stoning procedure for a cavity surface is as follows: • Stone with 240 grit • Stone at 90° to previous scratches with 320 grit until the previous scratches are removed • Stone at 90° to previous scratches with 400 grit until these scratches are removed • Stone at 90° to previous scratches with 900 grit

10.7 Mold Cavity Surface Finishing Polishing is the process of producing a series of overlapping “scratches” that get progressively ﬁner. To accomplish this, it is important that for each ﬁner grade stone used, the angle (direction) is changed relative to the marks made by the preceding coarser stone. After each grit ﬁnish is completed, the entire workplace must be thoroughly washed with clean stoning oil and wiped with a clean tissue to remove all particles of the grit remaining on the surface. This is necessary to ensure that none of the particles of the coarser grit will be picked up at a later time by a ﬁner grit stone, causing deeper scratches. Here are a few hints pertaining to the use of polishing stones: • Do not use a stone that is too coarse • Always dress the polishing stone with a grinding wheel or coarse paper to provide the maximum contact with the work surface • Use care when dressing the polishing stone • Use sufﬁcient stoning oil to prevent the stone from loading • Hold the polishing stone ﬁrmly for directional control, but press only hard enough to make the stone cut • Make sure the stone marks from previous grit size are all crossed out • Change stoning direction with each successive grit • Clean the work area thoroughly between each change of grit • Keep each grit of polishing stone in a separate stoning oil can • Exercise utmost care when stoning at an edge (parting line). If the mold cavities and cores need heat treatment, the question often arises as to how far the polishing should proceed before the part is heat treated. After the heat treating process, the knock out holes (or any other holes in the work area) must be covered or ﬁlled to avoid “dishing” (rounding off sharp edges) as the ﬁnishing continues. A 240 or 320 grit stone will remove the scale developed during the heat treatment. Make sure to “cross” stone the last marks until they are removed. Then proceed to ﬁner stones, such as 500, 600, or 900 grit. Surface ﬁnishes from stoning are often sufﬁcient after using a 600 to 900 grit stone. While the mold cavity surface exhibits a soft, “matte” ﬁnish, it will be smooth and ﬂat enough to allow easy ejection of the thermoplastic molded parts. Luster or shine can be developed on this type of surface with some 500 to 600 grit paper that will remove the ﬁne stone marks. Other methods for developing the ﬁnal “shine” on the surface include the use of diamond compound with brushes and felt bobs. Orange peel, the lightly dimpled ﬁnish that often appears during the ﬁnal bufﬁng operations of polishing, can be caused by several conditions. If the cavity insert steel has been overheated during heat treatment, the steel structure is changed resulting in an inconsistent surface hardness. This permits the softer surfaces to be deformed or abraded away more quickly than the harder surfaces. Consequently, care should be taken during heat treatment to ensure uniform surface hardness. If polishing pressure is too high, this softer area can be torn from the surface,

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10 Injection Mold Design leaving pits. By looking at the pits through a magnifying glass, a little “tail” at the edge of the pit following the direction of polishing can be seen. The following factors may contribute to the formation of orange peel and pits in polished mold cavity inserts: • Low mold cavity surface hardness due either to an incorrect tool steel composition or improper heat treatment • Presence of excessive nonmetallic particles • Presence of internal porosity, particularly at the center regions of mold cavity inserts of relatively large cross section • Retained austenite on the mold cavity surface usually caused by overheating during hardening or carburizing. Retained austenite has lower strength than martensite; therefore, it may cause deformation and breakage • Clustering of carbides causes overheating, nonuniform hardness, and low strength areas and the removal of the carbide particles can result in scratches during the carburizing process. If orange peel or pitting occurs, it may be possible to repair the mold cavity surface by ﬁrst removing the defective condition by hand polishing with a ﬁne stone, stress relieving at 100 °F below the tempering temperature, and hand polishing with diamond paste using light pressure and appropriate passes. If orange peel or pitting is caused by excessive retained austenite, this condition can usually be alleviated by transforming the retained austenite using additional tempering treatments. Using a second additional tempering treatment allows tempering of the brittle martensite formed during the ﬁrst tempering treatment. Another method of reducing the retained austenite content is to cool the mold cavity insert to a temperature between 100° and 150 °F. A subsequent tempering treatment of the brittle martensite formed during this cold treatment is advisable. Over-carburized or pack-hardened steel can also produce similar defects. In these instances, the orange peel is caused by a similar inconsistent hardness of the surface, but the pits are the result of very hard particles of iron carbide being pulled out of the surface. Most of these pits will have no tails. Orange peel can also appear on the surface of properly heat treated steel that has been ﬁnished with powered, mechanical devices. In this case, the surface defect is caused by excessive pressure of the polishing implement (brush, felt, etc.) against the steel, or by the over-polishing of the mold cavity surface. The harder the tool steel, the better the “polish” that can be achieved. Since orange peel occurs when the tool steel is stressed above its yield point (the yield point of steel can be increased by hardening or nitrating), it follows that the softer tool steels are more prone to orange peel. Remember, over polishing and polishing with too much pressure are the causes of pitting and orange peel. Diamond Polishing Polishing of thermoplastic injection mold cavity surfaces with abrasives is carried out either by some type of rotating machine or by hand. Machine polishing is more economical, but avoiding over-polishing caused by excessive pressure and/ or speed is more difﬁcult. Hand polishing usually involves preliminary work with a series of silicon carbide stones (240 to 900 grit) and a ﬁnishing operation with

10.7 Mold Cavity Surface Finishing diamond pastes from 15.0 to 2.0 micron. There is less chance of over-polishing if light pressures are used. Severe bufﬁng or the use of loaded stones may result in severe residual stress and pitting. The ﬁnal diamond polish is the last step in the process of polishing; if one of the previous stoning steps has not been done properly, the ﬁnal smooth luster surface will not be satisfactory. If mistakes have been made earlier during the ﬁnishing process, they will certainly show up on the ﬁnal surfaces. Stone ﬁnishing the mold cavity surfaces involves the preparation of the tool steel surface by producing a series of crisscrossing “scratches” that will get ﬁner and ﬁner. Coarse grits remove more metal and cut deeper. As the surface gets smoother and the surface begins to develop, it means that the peaks are being lowered to the depths of the valleys. In other words, the RMS measurement of the surface irregularity is becoming more minute. Since there is a practical limit to the size of the “scratch” that can be achieved with conventional abrasives, very small grit sizes of diamond must be used in order to obtain an optical quality polish or a Number One (1) ﬁnish. There is also another fundamental difference between the ﬁnishing process that has been done up to now and the ﬁnal diamond polish. The stoning process prepares the surface for the ﬁnal polish or the honing operation, in which the stone has the abrasive ability to cut the metal. On the other hand, diamond polishing is more a lapping operation. Fine diamond compounds remove very little metal while developing the ﬁnal luster of the ﬁnish. There are more than a dozen grades of diamond compounds ranging from 120 for fast cutting to 14,000 for super ﬁnishing, which are used to polish the mold cavity surfaces. The starting point of diamond polishing will depend, to some degree, on the sequence of stones that have been used to prepare the mold cavity surface. Beginning with a coarser diamond grade, a small amount of the compound is applied directly to the surface being worked. Then, by means of a bristle, brass, or steel brush, the compound is swirled over the surface using a rotary tool at slow speed. A speed of 500 rpm for roughing and 5,000 to 10,000 rpm maximum for ﬁnal polishing is a good general rule. Using light to moderate pressure, care must be taken to keep the brush ﬂat to the mold cavity surface to avoid cutting deep swirl marks. A “crisscrossing” action should be employed when using diamond compounds. The compound in use will become darker, indicating that metal is being removed and mixed with the compound. The surface should be brushed until all that is visible are ﬁne swirly marks left by the brush’s rotary action. There should be no stoning marks visible at all. The next step, the removal of the swirly marks left by the bristle brush, is accomplished with a felt product. Felt bobs are available in various degrees of hardness, preassembled in a shanked nylon holder. Mounted in a rotary tool and using light to moderate pressure, the surface is polished with diamond compound until all that is visible are felt swirls. The cavity surface should be thoroughly cleaned to remove all residual particles of the previous grade, before applying a ﬁner diamond compound. The ﬁnal step in polishing with diamond compound is a hand operation. Depending on the various conﬁgurations of the mold cavity surface, ﬁne tissue paper, felt sticks or cotton swabs may be used with an ultra ﬁne grade of compound to arrive at the ﬁnal high gloss luster.

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10 Injection Mold Design A better polishing surface is found in the ﬁber direction than across the direction of the ﬁber. If the ﬁber direction can be made to correspond to the largest cavity surface, less difﬁculty from pitting during polishing will be encountered. For ﬁnishing a stainless steel cavity surface, it is important to use polishing and bufﬁng compounds without iron oxide. If iron oxide particles are impregnated into the surface of a stainless steel, rust spots and surface pits will develop. The following diamond polishing guidelines should be used: • Apply a small amount of diamond compound at ﬁrst, then add more if required • Do not mix the diamond compound grades • If the diamond compound gets dry or hard, add a clean diamond thinner or lubricant • If the ﬁrst grade of diamond compound used does not remove marks from the last stoning operation, STOP! These marks must be removed with a coarser grade of compound or you will end up with highly polished, shiny scratches • Clean mold cavity surfaces thoroughly before progressing to a ﬁner grade of diamond compound • Do not use more than one grade of diamond compound on the same brush, felt, or lap • Be sure that each step completely removes marks left from the previous step • Never use an abrasive stick or a piece of abrasive cloth because the grit could become embedded in the lap and cause serious scratches on the mold cavity surface.

10.8

Thermoplastic Injection Mold Bases

The majority of the thermoplastic injection mold designs falls within the twoplate mold categories. However, several other mold designs are also used by the plastics industry. These molds are known as two-plate, interchangeable cavity inserts, vertical encapsulation, lost core, three-plate, hot runner, insulated runnerless, and stacked types of molds.

10.8.1

Standard Mold Base Components

Because of the similarities in the two-plate mold construction, it is desirable to have some standard mold base available to permit the thermoplastic injection molds to be produced in quantity, with a short delivery time, thereby reducing manufacturing costs. Logically, it is advantageous for the mold makers to purchase a mold base quickly at reasonable cost, rather than expending engineering mold design time, steel selection, construction, and assembly of the mold base components. The mold base and component manufacturers have developed tool technologies for the new engineering thermoplastic polymers, producing a range of standard high quality products that are available, ready to use in the designs and manufacture of thermoplastic injection molds.

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10.8 Thermoplastic Injection Mold Bases

Top clamping plate

Locating ring Sprue bushing

"A" cavity plates

"B" cavity plates

Leader pin Shoulder bushing

Support plate Ejector retainer plate

Ejector plate

Ejector pin

Sprue puller pin Ejector housing

Figure 10-4 Standard two-plate mold base components

A thermoplastic injection mold base may be deﬁned as an assembly of mold components that conforms to an accepted structural shape and size. The proper type of mold base is speciﬁed by the number of plates, premachining of the plates (if desired), steel selection, dimensions, ejection travel distance, size and type of locating ring and sprue bushing required by the injection molding machine. The remaining mold base components are suitably attached together and a guidance system incorporated. Of course, the mold base does not include runners, cavities, cooling, venting, hardness, ﬁnishing etc.; these aspects of mold manufacturing must be left to a specialized mold maker. The thermoplastic injection two-plate mold has been adopted as the standard mold base by mold component manufacturers because this particular mold construction is the most widely used design in industrial practice. Mold bases in a wide range of sizes made to suit a variety of purposes are produced by a number of manufacturers. These standard mold bases need only to be machined to make the mold components for a particular part. It is necessary to know the terminology and function of the components that make up the mold base. Figure 10-4 shows the location and nomenclature of the basic thermoplastic injection two-plate mold base.

10.8.2

Functions of the Mold Base Components

The most important features of standard thermoplastic injection mold bases are reviewed here to provide the reader with a working knowledge of the basic components and their functions.

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10 Injection Mold Design Top Clamping Plate The top clamping plate supports the “A” cavity plate, locating ring, and sprue bushing. In some mold designs, these two plates may be combined into a single, thicker plate, which serves both functions. It holds the stationary half of the mold to the stationary platen of the injection molding machine. Locating Ring The locating ring is that portion of the mold that is ﬁtted into the stationary platen of the injection machine. Its intended purpose is to properly situate the mold in relation to the injection nozzle of the machine. Most standard injection machines use different diameters with a variety of design features to accommodate the platen entry for the nozzle. The correct sizes of locating rings are provided in the molds to ﬁt a particular type of injection machine. Where relatively thin-walled parts are being molded and injection pressures may be high, the locating ring may be required to retain the sprue bushing within the mold, so the nozzle and sprue bushing are aligned. Sprue Bushing The sprue bushing seals off the melt from the injection nozzle conveying the melt through a conical shaped internal channel and forcing it into the sprue puller, runner, gate, and cavity conﬁnes of the mold itself. The sprue bushing has a tapered internal round channel that can vary in size as needed. In proper mold design, the sprue is made as short as possible, consistent with a given part design, thereby reducing the injection pressure drop in the runner system. Both, the locating ring and the sprue bushing, are usually supported by the top clamping plate, which is used to support the stationary half of the mold. “A” Cavity Plate The “A” cavity plate contains and supports the cavity or cavities or the core insert, sprue bushing, and the runners for the parts to be molded. In some cases, the cavity may be cut directly into the solid steel plate, while in others the cavities can be constructed separately and inserted into pockets within the cavity plate. The “A” cavity plate is part of the stationary section of the mold half; this plate is where the leader pins are mounted. “B” Cavity Plate The “B” cavity plate contains or supports the other half of the cavity or a core section of the molded part and also contains the leader pin bushings. The plane between these two plates is the normal parting line of the mold, which separates the two halves of the tool. The “B” cavity plate is the top plate of the movable section of the mold half. It is used to hold the sprue puller and ejector pins as well as the core inserts, or the cavity inserts. Support Plate The “B” cavity plate is mounted on top of the support plate. The support plate is used to provide strength to the cavities to avoid deﬂection during melt injection inside the cavities. Ejector Housing The ejector housing parallel blocks are added to provide the height required for the movement of the ejector system. The base plate of the ejector housing is used

10.8 Thermoplastic Injection Mold Bases for clamping the moving half of the mold to the moving platen of the machine. The ejector housing is a single unit for the ejection system. The injection mold base is manufactured with the parallels (vertical supports) welded to the bottom clamping plate. Bottom Clamping Plate A bottom clamping plate secures the movable half of the mold to the movable platen of the injection molding machine. If the mold is exceptionally large, the ejector system may require additional support, provided by the insertion of support pillars that bear the load between the bottom support plate and bottom clamping plate. Ejector Retainer Plate Mounted on top of the ejector plate, this plate retains the ejector head pins, ejector return pins, and sprue puller pin through counter bored holes. Ejector Plate The ejector plate is bolted together with the ejector retainer plate to form a unit. It acts as a back support plate for the ejector pins, return pins, and the knockout bar. These pins pass through drilled holes in the “B” cavity plate, insert cavity, and support plate. Stop Pins The stop pins are mounted on top of the bottom clamping plate; they are used as stops for the ejector housing when the ejector system returns as the mold closes. Support Pillars The support pillars are round bars placed between the support plate and the bottom clamping plate; they have the same height as the parallels. Bolted to the bottom clamping plate, they are used as additional support to avoid deﬂection of the “B” cavity plate. Sprue Puller Pin Pin located below the main runner, directly under the large diameter of the sprue channel. It is used to pull the solid sprue out of the bushing automatically when the mold opens and the molded parts and runner system are ejected. Ejector Pins The ejector pins enter the cavity to make contact with the molded part. Return Pins The return pins contact the stationary cavity plate and prompt the movement of the ejector plates back to the normal position prior to the next injection shot (not identiﬁed in Figure 10-4). Leader Pins The leader pins, used to align the plates on the closing of the mold, are hardened and ground steel pins mounted into one of the mold halves. One of the leader pins is offset so that the mold halves can only be closed when the leader pins are in the correct relative position.

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10 Injection Mold Design Shoulder Bushings Hardened and ground steel bushings are mounted into the other half of the mold, in-line with the leader pins. They serve as bearing surfaces for the leader pins.

10.8.3

Types of Standard Mold Bases

Figure 10-5 shows two basic styles of mold bases; one has a rectangular cross section (this is the most common form), the other is round. The rectangular cross section mold bases are commercially available in various standard conﬁgurations as shown in Figure 10-6. It is important for the mold designer to appreciate that the majority of mold designs can be broadly classiﬁed within one of these mold base types. For example, a slide-type of mold can be described as a two-plate mold with the addition of a slide assembly. The mold designer may have to accept a compromise in mold size; mold base types, and the range of sizes; for certain molds, some modiﬁcation to the mold base is necessary to accommodate a speciﬁc feature. Considered overall, however, the advantages of using standard mold bases outweigh the disadvantages and for a large number of mold designs, the standard mold base can be beneﬁcially used by the mold designer, mold maker, and end user. Nevertheless, on occasion it is necessary to have a “made-to-measure mold” so a special feature may be incorporated or a speciﬁc mold cooling system be adopted, or simply because a standard mold base of a suitable shape or size is not available.

Figure 10-5 Rectangular and round standard mold bases

Economical series standard mold base

Most frequently used standard mold base

Five plate reverse mold base with stripper plate

Six plate reverse mold base with stripper plate

Standard mold base with one floating plate

Standard mold base with two floating plates

Figure 10-6 Types of standard commercial mold bases (rectangular)

10.9 Types of Thermoplastic Injection Molds Advantages of Standard Mold Bases • Design drawing for individual unit sizes, reduces mold design time • Less steel needs to be carried in stock, investment is reduced • Buying and stock controls are simpliﬁed • Mold base price is known, estimating the cost of the mold is easier • Waiting time for steel blanks, etc., is avoided • Shaping, planing, and drilling of steel plates and blocks is avoided • Machining, grinding, hardening, ﬁtting of pins and sleeves is avoided • The ejector plate is already in place • The individual mold plates are screwed and doweled together • Machining time and labor costs are reduced • Design work on the insert cavities can usually begin immediately • Mold base components are standard; if a component is damaged during manufacture or in production, the part can be quickly replaced • Mold maintenance and down time is reduced • Mold delivery time is reduced • Efﬁcient mold maker’s engineering and use of production labor, provide mold cost reductions, satisﬁed customers and higher proﬁts. Disadvantages of Standard Mold Bases • Number of mold base sizes available is limited • Large size mold bases are limited • Maximum depth of mold base plates is also relatively small • The ejector system travel may be larger than is actually required • Cooling channels in a desired mold location are difﬁcult to obtain because of conﬂict with other mold base components • Extra number of support pillars positioned relatively far apart are needed to avoid deﬂection of the mold “B” cavity plate • The support plate cannot be unscrewed independently to expose the ejector assembly; the mold’s moving half must be disassembled • Limited number of special tool steel plates for the standard mold bases.

10.9

Types of Thermoplastic Injection Molds

Designs for injection molds differ depending on the product design geometry, end use applications, product tolerances (type of gate and location), product life, allowable product cost, molding production requirements, efﬁciency (maximum economy), amount of mechanization, factors that will dictate the size of the mold and the type of thermoplastic material being molded. The most common types of injection mold designs are the following:

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10 Injection Mold Design

10.9.1

Two-Plate Molds

Figure 10-7 shows a so called two-plate mold that certainly has more than two plates. It is the common name for a thermoplastic injection mold with a single parting line. The parting line of a mold can best be deﬁned as that surface where both halves of the mold separate to permit the injection molded parts and runners to be ejected from the mold. One side of the cavity is mounted in the “A” cavity plate with the locating ring and sprue bushing assembled into the stationary half of the mold. The “B” cavity plate is part of the moving half of the mold and contains the cores, the runner systems, and the ejector system, that are operated by the knockout bar that is attached to the machine actuator and the ejector plate assembly, moving the ejector pins and sprue puller pin forward to eject the molded parts and the runner system when the mold opens. The molten thermoplastic is fed directly from the machine nozzle through the sprue bushing that is connected to the runner and edge gates, ﬁlling the cavities with melt. Then the mold cavities cool off, forming the ﬁnal solid product (small containers).

Locating ring Sprue bushing Top clamping plate "A" cavity plate Cavity insert Shoulder bushing

tin Par

Support plate

Leader pin

Ejector retainer plate Ejector pin Ejector plate Sprue puller pin

Ejector housing

Knock-out bar

Figure 10-7 Two-plate mold and components (cross section)

alf

gh

vin Mo

ne

g li

tin Par

"B" cavity plate Core insert

ne

g li

f hal ary n tio Sta

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10.9 Types of Thermoplastic Injection Molds

10.9.2

Round Mate® Interchangeable Insert Molds

The Round Mate® System consists of a circular master frame with easily removable interchangeable round inserts, as shown in Figure 10-8. For high injection molding production operations, or to produce different parts at the same time, the use of multi-cavity inserts and a hot runner system is available in all sizes. The Round Mate® System provides the ﬂexibility and ease of operation to change inserts in the injection molding machine in just ten minutes. This Round Mate® System saves needless hours of mold making time. Using the universal Round Mate® Frame, standard machining and chase work are done once.

Fixed frame half

Fixed insert

For new injection molded parts, only the individual, interchangeable inserts are needed at a reduced cost of a complete, traditional mold. The Armoloy plated, AISI 4140 tool steel master frame is completely rust and corrosion resistant. The design of the inserts allows the cooling system to be wiped free of all mineral deposits and corrosion. Maintenance is reduced to a minimum. Cavities and runner systems can be cut directly into the solid insert. Inserts are available in a choice of tool steels and can be heat treated or plated. SAE P20 (prehardened) and SAE H13 (heat treatable) are in stock for immediate shipment. Other tool steels, such as stainless steel 420, SAE S7, SAE A6, are available as special order. Master frames and inserts are available in 2.50, 4.00, 6.00 and 8.00 in diameters. The Round Mate® system has several advantages:

Moving frame half

Moving insert

Figure 10-8 Round Mate® frame and interchangeable cavity inserts

• Fast standard mold delivery time • Reduced mold costs. Design and labor time are reduced to a minimum because the master frame and mold insert blanks are standard. Normal chase work is already completed. The Round Mate® system has a unique water cooling jacket, cavity pockets, guided ejection system, and many other features already incorporated. • Fast insert changeover. Simple hex-key wrenches are used to change the mold insert in the injection molding machine in less than ten minutes. • Improved part quality. The beneﬁts of a round shape include more uniform clamping pressure in the mold, balanced heat transfer and cooling, completely self-centering, more efﬁcient runner layout, better part quality, and faster molding cycles.

10.9.3

Master Unit Die Interchangeable Insert Molds

MUD frame

With the advent of the standard mold base, Master Unit Die Products, Inc. has developed a system for the quick replacement of inserts, which are easily interchangeable within a single frame for all injection molding machine applications. This approach was certainly a major breakthrough in reducing mold costs for the plastics industry. By combining a standard mold frame with interchangeable mold inserts, this concept not only cuts mold building costs and reduces mold delivery time, it also increases productivity by reducing the injection molding machine down time signiﬁcantly. Master Unit Die Products, Inc. offers frames in four basic styles and companion inserts in two basic styles. These frames and their companion inserts are in use throughout the world. A typical MUD mold is shown in Figure 10-9.

Interchangeable companion insert

Figure 10-9 MUD mold frame and interchangeable cavity inserts

586

10 Injection Mold Design MUD systems ﬁnd applications in different areas: • Prototype parts. The MUD system is ideal for prototype parts. Fast mold design is achieved by not using the entire mold assembly. Prototype part mold costs can be reduced to an absolute minimum. • Short production runs. The almost instant interchangeability of MUD inserts is a big advantage when a variety of parts are scheduled for production. Production time can be gained during changeovers. MUD systems offer several advantages: • Lower mold costs. The concept makes the inserts easily interchangeable within a single frame. Mold costs are reduced because only the insert is replaced, not an entire standard mold base • Quicker delivery. The simplicity of the system means less time is required for mold fabrication. Standard blank inserts are usually available from stock within ten days • Faster setup. The Master Unit Die frame remains in the injection molding machine with the cooling lines operational when inserts are interchanged. Initial production setup is simply a matter of sliding and clamping the insert in position (leader pins and bushings in the core and cavity plates assure alignment) • Instant changeover. Inserts are easily removed and interchanged by loosening four clamps, disconnecting the cooling lines, removing one insert and sliding in another, then simply reversing the process, the complete change taking normally less than ten minutes • Minimum purging. The injection molding machine’s plastifying unit is less likely to overheat, requiring purging, because of instant changeover of the MUD inserts (material and production savings). • Easier maintenance and repair. The MUD inserts are easier to remove and reinstall, they are lighter, smaller, easier to handle and store, which is a big plus when maintenance or repair is required • Greater ﬂexibility. Most Master Unit Die frames will accommodate two or more inserts. When single molding applications are scheduled, a blank insert can be installed in the other section of the frame. • Maximum versatility. Inserts are available in “T” and standard styles. Mold design latitude is almost unlimited because inserts can be engineered for parts requiring stripper plates, sleeve ejection, single or double cam action, hydraulic, mechanical or pneumatic powered cylinders, and any feature desired including three or four plate molds.

10.9.4

Three-Plate Mold Cold Runner System

The three-plate mold design is used where a center pin point gate is required on a multi-cavity mold or when the molded part geometry requires a stripper plate ejection. This three-plate mold system operates as follows: The mold is opened at the parting line and the sprue is pulled immediately by the sprue puller pin. The entire three-plate runner system thereby moves back with the moving cavity

587

10.9 Types of Thermoplastic Injection Molds

Figure 10-10 Three-plate mold cold runner system

plate, the runner is withdrawn from the sucker pin by the stripper plate. When the ﬂoating distance of the sucker pin has been taken up, the three-plate runner system with the pin point gate breaks off from the cavity as a result of the pull-off force produced by the sucker pin. The continuous movement of the “A” cavity plate causes the sucker pin to shear off from the three-plate runner and frees itself from the cavities ready for ejection. When the mold is closed, the plates are progressively returned to their original positions. An illustration of a three-plate mold system, cold runner, and a pin point gate is shown in Figure 10-10.

10.9.5

Vertical Insert Mold for Thermoplastic Encapsulations

The vertical mold clamp and horizontal injection system for encapsulating inserts was developed to produce electrical and electronic thermoplastic products such as electric cord plugs, transformers, motor housings, automotive speed and temperature sensors, etc. Figure 10-11 shows a thermoplastic encapsulated automotive ABS sensor and a fuel injector produced by this molding process. The properties of special thermoplastic injection molding polymers are ideal for encapsulating various types of inserts by using vertical clamp and horizontal injection molds. These polymers have excellent electrical properties, high end use temperatures, and good chemical resistance. Some of these thermoplastic resins are also “wire friendly” for encapsulating delicate electrical wires and leads. Figure 10-12 shows a typical vertical clamp and horizontal injection mold for encapsulating wire wound bobbins and magnet inserts for the production of electronic ABS sensors. This type of mold may look similar to a two-plate mold, but it operates completely differently. The upper half of the mold (top clamping and “A” cavity plates) is mounted to the moving vertical platen of the machine, but it does not have a sprue bushing and a locating ring. The lower half of the mold is mounted to the ﬁxed platen, shuttle platen, or rotating table of the machine, depending on

Figure 10-11 Thermoplastic encapsulated ABS sensor and fuel injector

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10 Injection Mold Design

Figure 10-12 Vertical clamp, horizontal injection unit

the application. The ejection system is located in this lower half and connected to the knock-out bar that is operated by a hydraulic actuator. The injection unit is in-line with the mold parting line; the melt is injected into the split sprue bushing (mounted in each half of the mold) that is connected directly to the main cold runner and sub runners, distributing the melt through the gates to the cavities. The inserts are loaded horizontally in the lower cavities, on top of the support pins. When the mold is closed, the spring loaded upper support pins lock the inserts in position, preventing the inserts from moving during the melt injection in the cavities. The melt shrinks around the inserts forming a good mechanical bond between the inserts and the thermoplastic during the mold cavity cooling. The molded parts and the runner system remain in the lower half of the mold, when the mold opens, the ejector pins push the molded parts out from the lower cavities and the operator removes the parts and loads the inserts to start a new molding cycle.

10.9.6

Hot Runner Molding Systems

In a hot runner system the melt directly ﬁlls the cavity from which the solidiﬁed thermoplastic part is ejected when the mold opens. The melt is conducted through heated channels through the runnerless manifold where the melt is proportionally subdivided, depending on the number of cavities and injected to the cavities through the temperature controlled gates or drops. Hot runner molding can result in as much as a 50% cost savings, part quality is improved, cycle time is faster, and since parts are ﬁnished when they leave the mold, it is not necessary to remove sprues or runners. Hot runner molding systems eliminate the use of cold runners, the melt is maintained at a constant temperature until it reaches the cavities, providing a uniform melt viscosity and better injection pressure control reducing the melt shot capacity. The hot runner molding systems are more expensive, they require an additional temperature controller, higher maintenance, and operational training. Figure 10-13 shows a typical hot runner mold cross section illustrating the operational sequences of this molding process.

10.9 Types of Thermoplastic Injection Molds

Figure 10-13 Typical hot runner mold system (cross section)

10.9.7

Hot Runner Mold Temperature Control Systems

Second only to gating, the most common source of problems in hot runner molds is temperature control throughout the system. A high degree of thermal uniformity in the manifold and from drop to drop is essential to avoid degradation of resin and to obtain uniform ﬁlling, packing, shrinkage, appearance, etc. in all cavities. Each drop has a different temperature proﬁle. The temperature setting for drops may be substantially higher than in the manifold. Manifold temperature should not exceed the process melt temperature. However, lower manifold temperatures will minimize minor streamlining ﬂaws with thermally sensitive resins. If residence time in the manifold is short, a low manifold temperature will have little effect on the temperature of the ﬂowing melt. Some of the devices used to maintain uniform temperature control are the following: Cartridge Heaters The runnerless manifold temperature is controlled with cartridge heaters. The temperature is dependent on the cartridge heater design, wattage, location, support, and installation tolerances of heaters, thermocouples, thermal conductivity of metals used and independent zone controls. Serious problems may result when a single cartridge heater fails, causing the controller to increase heat from other cartridges in the same zone. Temperature uniformity will be upset by the development of hot or cold areas, increasing the overall manifold temperature. Loose ﬁtting cartridges also tend to burn out more frequently. Cartridge heaters are also used as internal heaters inside the manifold runner channels or as torpedoes for drops. Different installation tolerances will cause temperature variations because the thermocouple in the cartridge heater is insensitive to ﬁt problems. Tubular Heaters Tubular heaters, similar to those used to heat the oven of an electric range, are also used to heat hot runner manifolds. They are pressed into grooves on both

589

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10 Injection Mold Design sides of the manifold plate. They may be shaped and conﬁgured to follow the melt ﬂow paths around the connecting area of each drop. They also provide a more uniform heat input along the length of the heater and are more resistant to burnout because fewer individual heaters are used. Cast-in Heaters Thermally conductive metals such as copper and copper alloys are often used in hot sprue bushings, portions of manifolds, and drops to improve heat distribution and uniformity. Permanent heater elements are often cast into these metals to improve durability. Although failure is much less frequent, they are more expensive. Heat Pipes Kona Corporation has patents on the use of heat pipes to obtain uniform temperatures in manifolds and drops. These devices not only conduct heat rapidly from one area to another, but they tend to equalize temperature along their path by releasing more heat to colder areas. This is a natural result of conducting heat by vaporizing and condensing a liquid in the heat pipe. The colder the area, the more liquid will condense and the more heat will be released. Heat pipes require external heater bands located only at the end of the bushing. They also make it possible to reduce manifold temperature below that needed to keep gates operating, which is an important consideration for heat sensitive resins. Bands and Coil Heaters These are most often used in sprue bushing, pipe connected manifolds and drops. Coil heaters lend themselves to the necessary wattage distribution to avoid overheating the center section of a bushing in an effort to provide enough heat to keep the gate open. Placement of band heaters close to the gate area and use of conductive metal in drops or sleeves can also control the temperature distribution. Torpedo Heaters They are used in sprue bushings, drops, and internally heated manifolds. These torpedo heaters are helpful in distributing wattage effectively in drops and internally heated manifolds. Manufacturers of torpedo heaters have made signiﬁcant progress in overcoming the inherent lack of temperature uniformity of heated manifolds or torpedoes supported at one or both ends in cold steel mold plates.

10.9.8

Hot Runner Mold Gates (Drops)

The most critical hot runner mold design parameter is the gate. The thermoplastic melt must be kept in a ﬂuid condition up to the point of separation. The gate area must freeze rapidly and without serious ﬂaws. Control of gate freezing along with gate vestige, part quality, and appearance are not only functions of temperature control, but of the speciﬁc gate design details needed to avoid serious molding problems. A small insulating ring of titanium (very low heat conducting metal) between the bushing and cavity is used to maintain the temperature differential at the gate. The types of gates used with the hot runner molds are the following:

591

10.9 Types of Thermoplastic Injection Molds Sprue

Thermocouple

Hot bushing Hot bushing

Hot bushing

Thermocouple

Coil heater Coil heater

Air gap

Top insert

Insert Insert Molten insulation

Straight sprue gate

Reverse taper sprue gate

Insulated insert

Straight sprue insulated gate

Figure 10-14 Two sprue gates (drops), straight and reverse taper Hotbushing

Straight Sprue Gate

Thermocouple

Coil heater Top insert

An open gate is a simple, restricted opening to the cavity. It is usually difﬁcult to keep this gate from freezing with semi-crystalline and high melt temperature resins, which leave a long vestige. A straight sprue gate is shown in Figure 10-14, left illustration.

Pin hole Molten insulation

Insulated pin point gate

Reverse Taper Sprue Gate The reverse taper gate is a modiﬁcation of the straight sprue gate. A reverse taper sprue gate land between 0.25 and 0.75 in moves the separation point away from the cold cavity and into the bushing where it can be controlled better, making this gate more suitable for engineering polymers. The resulting long gate vestige on the part is a serious disadvantage. A reverse taper sprue gate is shown in Figure 10-14, right illustration.

Insulated insert

Figure 10-15 Two insulated gates, straight sprue and pin point

Straight Sprue Molten Insulated Gate Figure 10-15 shows these types of gates; they are straight sprue gates with molten insulated traps for melt surrounding the gate close to the cavity. These traps are intended to provide an impediment to heat transfer away from the gate. Much of the thermoplastic trapped in this insulated melt may freeze, but portions closest to the gate may be hot, which constitutes a serious hold-up spot for heat sensitive resins.

Main sprue

Cavity Air gap Sub sprue

Tit Edge Gate The concept of a hot bushing drop feeding two or four side tit edge gates is an attractive way to simplify runnerless molds for small parts. Current designs do not appear to be well suited to control gate freezing and separation for semicrystalline and high temperature resins. Tit edge gates tend to shear ﬂush with the gate opening and leave a frozen stub in the gate, which either freezes solidly preventing subsequent shots or is injected into the next shot as a cold slug. Figure 10-16 shows these types of gates.

Sprue and tw o horizontal tit edge gate s

Hot bushing Coil heater

Hot Tip Spreader Insert Gate The hot tip spreader is a small, conical insert in the gate end of a hot bushing drop made of copper alloys to aid heat transfer. The spreader’s internal tip contacts the inside diameter of the hot bushing and draws heat from the melt and bushing heater. Several streamlined passages around the cone allow melt to pass into the cavity. The purpose of this spreader is to deliver heat to the gate separation point (prevent freezing) and to establish the exact location of the separation point.

Gate inserts

Sprue and angled tit edge gates Figure 10-16 Hot runner drops, main and sub sprues tit edge gates

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10 Injection Mold Design

Figure 10-17 Hot runner drop, hot tip spreader gate

The external spreader tip is very close to the cavity surface and it tends to limit gate vestige. Figure 10-17 shows this type of gate. Reciprocating Pin Valve Gate A valve gate is an open gate with a reciprocating valve pin. The valve is opened and closed during each cycle by means of a hydraulic, pneumatic, or spring mechanism. Valve pin gates provide maximum control of gate vestige and cosmetics, but they add complexity and introduce serious wear problems for glass or mineral reinforced thermoplastic resins. Valve gates also tend to push cold slugs into the gate area of the part, causing lower impact resistance or appearance problems. Figure 10-18 shows a pin valve gate (top illustration) and a guided pin valve gate (bottom illustration).

Figure 10-18 Hot runner drops, reciprocating pin valve gates (Courtesy: Mold Master)

Hot Tip Fixed and Valve Torpedo Gates These types of gate are established, when an internally heated hot tip torpedo is used in the drop. The hot tip torpedo approaches or enters the gate area. The reciprocating hot tip torpedo works as a valve pin controlling melt ﬂow at the gate. The torpedo is forced to retract from the gate by the injection force of the melt on the torpedo tip (gate open), just before the melt is injected. When injection is complete, Belleville springs above the torpedo ﬂange support, which were compressed during injection, cause the torpedo to close the gate. Figure 10-19 shows three types of hot tip torpedo gates. Melt

Insert

Torpedo down

Torpedo up

Melt Air gap

Hot bushing

Water

Water

"O" ring

Gate close

Fixed hot tip torpedo and water heated insert

Gate open

Hot tip valve torpedo and water heated insert

Figure 10-19 Hot runner drops, hot tip ﬁxed and valve torpedo gates

Fixed hot tip torpedo and hot bushing

593

10.9 Types of Thermoplastic Injection Molds Spear tip "OFF" (cold) gate close and stop melt flow

Water cooling

Stop melt flow

Coil heater "ON" "O" ring

Gate insert

Fixed Spear Tip "OFF" and Water Heated Insert Spear tip cold

Cold slug

Spear tip "ON" (hot) gate open and melt flow into cavity Water cooling Coil heater "ON" "O" ring Melt flow

Spear tip hot

Gate insert

Figure 10-20 Hot spear tip thermo-valve gate (A) spear tip “OFF” (cold) gate closed and melt ﬂow stops; (B) spear tip thermo-valve gate “ON” (hot) gate opens and melt ﬂows into cavity

Hot Spear Tip Thermo-Valve Gate This type of thermo-valve gate is a ﬁxed hot tip torpedo with dual temperature control probe heaters. The main heater in the probe keeps the thermoplastic melt ﬂowing through the whole length of the drop and a very small tip heater raises the tip temperature very rapidly. The tip heater cycles during every injection molding cycle, permitting the melt to ﬂow (high tip temperature) during injection and stopping the melt ﬂow or freezing (low tip temperature) during the remainder of the cycle. This thermo-valve gate is a patent of the Spear Systems, Inc. Figure 10-20A shows the spear tip “OFF” or gate closed and Figure 10-20B shows the spear tip “ON” or gate open.

10.9.9

Types of Hot Runner Molding Systems

There are three different methods used to control the molten thermoplastic temperature inside the hot runner molding systems: • Internally heated hot runner molding systems use torpedo cartridge heaters with thermocouples that are mounted inside the runners, leaving a clearance area between the outside diameter of the torpedo and the inside diameter of the runner to convey the melt • Externally heated hot runner molding systems use different types of heat sources to control the manifold melt temperature surrounding the runner channels

594

10 Injection Mold Design • Insulated hot runner molding systems take advantage of the heat insulation properties of the thermoplastic melt by allowing relatively thick layers of thermally insulating thermoplastic to freeze around a large outside diameter of an insulated runner, allowing the hot thermoplastic melt to ﬂow through the center of the insulated runner. The balancing of hot runners is an important mold design requirement, as well as for cold runners. A balanced runner distributes the melt to the cavities at the same time, packing all the cavities with the same injection pressure, to obtain uniform dimensional control in all the molded parts. Where there is a substantial difference in temperature between the hot manifold and the cold mold plates, provisions must be made to adjust for the differential coefﬁcient of linear thermal expansion between these two components. Supports, spacers, and some bushings may be surface mounted on plates so that they can slide during heating and cooling, although this can lead to problems in sealing the manifold and drop melt pressure in the channels. The mold design can accommodate movement during expansion or contraction, but the length must be adequate to allow this motion. Drops containing torpedoes can accommodate expansion, if the torpedo is heated ﬁrst and cooled last so that the torpedo is not bound in thermoplastic, and can ﬂoat during expansion and contraction. This, however, does provide an opportunity for error by the operator and possible damage to the mold. Insulated hot runner molding systems will not tolerate signiﬁcant molding cycle interruptions without freezing of the melt ﬂow path or insulated runner, which then requires removal of the insulated runner and restarting the mold. A delay time of one to two minutes can cause freezing of the melt path depending on resin melt temperature, mold temperature, melt throughput, and other factors. Although runner removal and restarting are relatively simple, it takes between ﬁve to ten minutes. Recurring interruptions will affect productivity and become a problem. The removal of the insulated runner can be a process advantage; it eliminates the need for the difﬁcult purging process required during start-up or color change; in addition, insulated runners are always clean and empty for startup. Small particles of dirt or trapped metal that plug small gates are also easily removed by changing the insulated runner. In a hot runner mold, with externally or internally heated manifolds, the restarting and cleaning process requires a lot of maintenance work and operational down time. It requires removal and disassembly of the hot runner molds, requiring heaters, seal replacements, and cleaning using ﬂuidized bed equipment, or sending the unit to the vendor for service. 10.9.9.1

Internally Heated Hot Runner Systems

The thermoplastic melt is injected through a sprue bushing heated with an internal torpedo. The system consists of bores drilled into the internally heated manifold. A torpedo is inserted in each runner and the drops are centered with end caps. The melt ﬂowing in the runners, around the torpedo’s surface, is kept molten by a thermocouple and cartridge heater inside the torpedo. Heating the material “from the inside out” is very efﬁcient, since it allows heater loads to be much less than externally heated hot runner manifolds. The outer layer of the melt stream solidiﬁes to form an insulating layer of solid thermoplastic that further

595

10.9 Types of Thermoplastic Injection Molds Locating ring

Runner (trapezoidal)

Resistance element

Cartridge heater

Runner torpedo Insulation

Gate torpedo

Sprue torpedo

Melt flow (round) Cavity

Figure 10-21 “H” manifold, internally heated hot runner

reduces energy requirements and permits the manifold surface temperatures to be maintained between 100 and 110 °F lower than the torpedo. Often, internally heated hot runner designs have a single main runner that intersects a gate torpedo to direct the melt ﬂow to the gates. The main and secondary runners use a common “H” manifold layout pattern providing a balanced melt ﬂow between the main and the intersecting secondary runners. A variety of internally heated manifold runner layouts are possible for carrying molten thermoplastic to the cavities. Either round or trapezoidal runner cross sections are used to deﬁne the manifold runner parting line. The use of integral cartridge heaters in all torpedoes also improves heat transfer throughout the system. The hot runner internally heated system temperature control inside the melt channels (sprue, runners, and drops) is inherently nonuniform, because the torpedo must be adequately supported in the colder mold plates at one or both ends. The torpedo cartridge heaters tend to run hotter as the distance from the supported end increases. Streamlining the internal runner layout is not possible, when the large main runner diameters and internal torpedoes intercept the sprue, sub runners and drops that are located perpendicularly, the interception from the path is offset, causing hold-up spots in the conveying channels and the necessity for the thermoplastic melt to ﬂow around these obstacles. Cleaning the internally heated hot runner system requires disassembly of the mold, removing the runner with the torpedoes, and reassembling the mold using new torpedoes and pressure seals, because they are difﬁcult to remove from the runner without damage. Figure 10-21 shows a cross section view of an internally heated hot runner system. High temperature engineering thermoplastic polymers and heat sensitive resins are not recommended for use with the internally heated hot runner systems. However, for commodity thermoplastic resins processed at low temperatures that are not heat sensitive (PE and PP), this system is ideal for high production applications. The cost of the internally heated hot runner systems is lower than the externally heated hot runner systems. 10.9.9.2

Externally Heated Hot Runner Systems

Each externally heated hot runner manifold design has a balanced runner conﬁguration, free of hold-up spots and smooth surface ﬁnish to maximize part

596

10 Injection Mold Design production. The hot sprue bushing satisﬁes various melt ﬂow requirements, having a replaceable seal ring that provides a leak proof seal between the hot sprue bushing and the manifold. The externally heated hot runner manifolds use several temperature control systems, such as cartridge heaters, cast-in heaters, heat pipes, heater bands, coil heaters, and torpedo heaters to ensure full heat transfer and faster start-up. They are designed to support high injection pressures (20,000 psi) and temperatures, with adjustments for the differential coefﬁcient of linear thermal expansion between the hot manifold and the cold cavity plate that houses the drops. Supports are used to avoid deﬂection with spacers for thermal insulation. A structural joint seal design is used for assembling the hot sprue bushing, main and sub manifolds, and drops. When selecting a nozzle, it is important to take into consideration the part design complexity, the wall thickness, ﬂow length, and the shear rate of the resin. The nozzle should satisfy a range of shot weights based on the internal channel length and diameter, and the viscosity of the material to be molded. Several types of gates (drops) are used for the externally heated hot runner systems, depending on the gate vestige, color changes, resin, shot size, and overall performance required by the application. The drops are mounted in the cavity plate clearance pocket, using an air gap for thermal insulation of the drops and the thermal expansion of the hot manifold. The externally heated hot runner systems require that the heated components, manifolds, and drops be insolated from the cooler parts of the mold. Some type of separation between the manifold and the mold plate pocket is needed to prevent complete freezing or to maintain adequate ﬂow through the melt path diameter. Hot tip insert torpedo gates

“A” cavity plate

Tubular heater

Hot sprue bushing

Locating ring

Externally heated “H-H” manifold

Top clamping plate Figure 10-22 Externally heated,“H-H” hot manifold with mold plates

Hot runner clearance

597

10.9 Types of Thermoplastic Injection Molds Main & "X" sub manifold valve gated

Hot sprue bushing

Main & "X" sub manifold hot torpedo gated Locating ring Main manifold

Valve actuator

"X" sub manifold

Valve pin

Pin valve gate drop

Hot tip insert torpedo drop

Figure 10-23 Two integrated cast-in externally heated hot manifolds

The externally heated hot runner manifolds, sprue bushing, and drops must be designed carefully to provide a streamlined ﬂow path. All these components should be manufactured from corrosion resistant, high heat conductive, high temperature resistant, and high strength metal alloys. The externally heated hot runner systems are more expensive, require more training and maintenance, are difﬁcult to clean, and need more time to heat the polymer inside the ﬂow channels for restarting the molding process after a long interruption. Figure 10-22 shows a typical externally heated hot runner tubular system using an “H-H” manifold with hot tip insert torpedo gates. Figure 10-23 shows two integrated cast-in, externally heated hot runner manifolds (main and “X” sub manifold). The left illustration has reciprocated pin valve drops and the right illustration has hot tip insert torpedo drops. 10.9.9.3

Insulated Hot Runner Molding Systems

The insulated hot runner molding system was developed at Du Pont’s Plastics Technical Center in Geneva, Switzerland more than 25 years ago. The insulated hot runner mold is not really a commercial system. Du Pont has arrangements with special suppliers for the three hot tip torpedo designs and the additional system components. However, Du Pont technical assistance is required for the selection of the insulated hot runner molding system, mold design recommendations, detailed drawings of the speciﬁc system components and speciﬁcations. Information about the closest supplier of system components, mold start-up and troubleshooting assistance, insulated hot runner molding system technical training and other services are offered only to Du Pont customers. The insulated hot runner mold system is a particularly good choice when there will be frequent color changes or for other reasons such as high quality requirements that make thorough purging of a hot manifold very troublesome. Insulated hot runner molds are more difﬁcult to start and operate than threeplate molds, but are much easier than other hot runner molds. An insulated hot runner mold is a cross between a hot runner mold and a threeplate mold. The insulated runnerless mold system is relatively inexpensive and ideal for heat sensitive thermoplastic polymers. Most of the higher melting point engineering polymers are a little more difﬁcult to process because of the greater differential temperatures between melt and mold.

598

10 Injection Mold Design The thermoplastic melt remains ﬂuid enough in the center of the insulated runner to permit repeated injection, even if the runner inserts and the mold plates are colder. The solidiﬁed melt on the surface of the insulated runner provides exceptional thermal insulation for the inner ﬂuid core. The insulated runner is nearly round and has either approx. 0.90 in diameter for the standard systems or 0.50 in diameter for the mini-insulated hot runner molds. The runner half in each insulated runner insert mounted in each mold plate is tapered and cut a little less deep than the radius to ease runner ejection. The insulated hot runner mold systems beneﬁt from thermal insulation properties of the thermoplastic polymers. All runner systems require some thermal isolation to aid in start-up, to prolong freeze-off during brief cycle interruptions, and especially to keep ﬂow paths sufﬁciently large to avoid an excessive injection pressure drop through the insulated runner. Thermal isolation is achieved by using two separate insert halves to form the large insulated runner and by reducing the metal to metal contact with adjacent cavity plates. Reduction of contact is achieved by “wafﬂing” the contacting insert surfaces and mold plate cavities or by inserting a thin (0.005 in thick) sheet of Nomex® aramid paper between these two base surfaces. In addition, coolant lines around the sprue bushing and cartridge heaters without thermocouples are temperature controlled by a common unit. The cartridge heaters should be mounted at each side, in the same direction and close to the half runner. Both runner insert halves should have the same amount of heaters for a thermal balance of the insulated runners. The runner inserts and sprue bushing operating temperatures should be less than the melt temperature and higher than the mold cavity temperature. Figure 10-24 shows a cross section view of the Du Pont reciprocating insulated hot runner molding system. The other Du Pont systems operate the same way, except that the hot tip torpedoes have different conﬁgurations and are ﬁxed, requiring the torpedo gate height to be adjusted by moving the spacers on top of the hot tip torpedo, to satisfy the application requirements.

Figure 10-24 Du Pont insulated hot runner reciprocating system

10.9 Types of Thermoplastic Injection Molds The temperature of the gate area is controlled by circulating mold coolant around the lower torpedo insert bushing using “O” rings to avoid freezing the hot gates. This bushing is also useful for replacing worn or damaged gates without major reconstruction of cavities. If unusually low mold coolant temperatures are to be used, the requirements for maintaining an adequate ﬂow path through the insulated runner may become more stringent. Higher throughput and/or better heating and isolation of the insulated runner may be necessary. To maintain an adequate open ﬂow path through the insulated runner, a minimum of 8 g of the melt must ﬂow through each insulated runner and drops every minute. This can be accomplished by a combination of short molding cycles and small parts or longer cycles and larger parts. Many insulated runnerless mold systems are not directly gated into the part but instead into small sub runners. This is done for a variety of reasons: • To cluster small parts and increase melt throughput • To hold large numbers of cavities with fewer drops • To use multiple gates in a single part to control roundness or to achieve adequate ﬂow in long, thin, or complicated parts • To gate through sub runners with tunnel gates to eliminate gate vestiges or stringing. Long insulated runners are not recommended because the melt ﬂow decreases with distance and causes bending of hot tip torpedoes during ejection of the insulated runner because of shrinkage. The maximum ﬂow distance from the machine nozzle to the farthest drop should not exceed 15.00 in. During molding cycle interruptions or at the end of a shift, the entire cross section of the insulated runner will solidify in time and provisions must be made to ensure easy removal of the solidiﬁed insulated runner system. This is accomplished most simply by closing the mold, removing the four latches from the insulated runner parting line of the mold that normally remains closed and replacing the four latches to lock the main cavity parting line. Then, the mold is opened at the insulated runner parting line, the runner is pulled off the hot tip torpedoes by the stripping washer force activated by stripper bolts under the extended runner lips, dropping the insulated runner automatically. Without this automatic feature, damage to hot tip torpedoes, parting line, etc. during manual insulated runner removal is almost certain. Four parting line latches (two at each side) are usually sufﬁcient, because these latches hold either parting line closed, during injection of the molded part or the insulated runner. The molding machine clamping force is needed to resist injection force. Because the two parting lines of an insulated hot runner mold are stacked, the clamping force required is determined by the projected area of the cavities or of the insulated runner, whichever is greater, not by the sum. One must never attempt to purge through an insulated hot runner mold with these external latches, because the mold will be damaged and personnel can be injured. There is no need for purging through an insulated hot runner mold; runner removal is easier and more effective. Figure 10-25 shows the Du Pont insulated hot runner operational sequence; the top two illustrations show the normal injection molding process, where the mold cavity is injected and the molded part is ejected. The bottom two illustrations show the automatic ejection of the insulated runner.

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Figure 10-25 Du Pont insulated runner mold operational sequence

The insulated hot runner molding systems employ three hot tip torpedo design shapes. The three hot tip torpedoes are built to meet Du Pont speciﬁcations. They contain a short special design cartridge heater at the bottom with a ceramic insulator above this to protect the wires and a hold-down spring above the insulator. The thermocouple is in the cartridge heater and both are pressed inside the torpedo tip insert to control the heat toward the tip. The torpedo tips inserts/heaters are mounted (hot interference ﬁt) with the torpedo shells, the insert tips are diamond polished, wear resistant and thermally conductive. Standard Fixed Hot Tip Torpedo “DU-085” The standard ﬁxed hot tip torpedo is for general purpose applications and is most commonly used for engineering polymers. The hot tip torpedo is ﬁxed in position during the molding process. The control of gate vestiges and freezing is dependent on the tip temperature and the tip location in the gate area. The exact tip position can be adjusted for different resins, molding cycles, and to correct gate vestiges or freezing problems. This is done by interchanging the four spacers of slightly different wall thicknesses; one spacer is above the hot tip torpedo top support ﬂange and three other spacers below. This adjustment will be made during original development and start-up of a new mold and will

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10.9 Types of Thermoplastic Injection Molds rarely have to be changed afterwards. Figure 10-26 shows the standard ﬁxed torpedo “DU-085”. Reciprocating Standard Hot Tip Torpedo “DU-005” The difference between the ﬁxed and the reciprocating hot tip torpedoes is their geometry. The reciprocating hot tip torpedo works as a mechanical valve to control the melt ﬂow at the gate. The reciprocating hot tip torpedo is forced to retract from the gate seat by the injection force of the melt on the tip of the torpedo (gate open), just before the melt is injected. When injection is complete, Belleville spring washers above the support ﬂange, which were compressed during injection, push the hot tip torpedo down against the gate seat, closing the gate. The reciprocating hot tip torpedo stroke, although it is only 0.06 in, may cause the cartridge heater or even more delicate thermocouple wire breakage. Figure 10-27 shows the standard reciprocating hot tip torpedo “DU-005”.

Figure 10-26 Standard ﬁxed hot tip torpedo “DU-085”

Figure 10-27 Standard reciprocating hot tip torpedo “DU-005”

Fixed Mini Hot Tip Torpedo “DU-200” The ﬁxed mini insulated hot runner molding system satisﬁes the need for a runnerless mold to ﬁt in small injection molding machines used in the production of very small parts and to permit closer mold layouts of these small cavities. The small molding machines are often incapable of holding larger molds with the size and stack height required by the standard large hot tip torpedo systems. Since the ﬁxed mini insulated hot runner molding system depends on the same limitations imposed by the insulated runner concept, the insulated runner diameter for this system is smaller. The test results of the ﬁxed mini insulated hot runner molding system have shown that much lower melt throughput than 8.00 g/min is viable. The ﬁxed mini hot tip torpedo has a larger diameter upper portion and a tapered lower portion design to improve the durability of the unit. The ﬁxed mini hot tip torpedo operates with a 120 V cartridge heater and a thermocouple for closed loop control. Figure 10-28 shows the ﬁxed mini hot tip torpedo “DU-200”.

10.9.10 Thermoplastic Stack Injection Molds Thermoplastic stack injection molds are similar to standard molds, but they use cavities on more than one level or on both sides of the mold parting lines. These types of molds require higher clamp forces, usually 15 to 20% more, but they can increase molding output by 80 to 100%. Applications of stack molding include shallow containers and lids, threaded and tamper proof bottle caps, disposable tableware, and deep draw containers. The stack mold was originally developed in the plastic compression molding industry. Admiral Distributors, Inc. of Chicago, Illinois in 1955 employed two identical sets of compression molds, one above the other, used for the production of phenolic radio phonograph lids. This allowed two pieces to be molded from a conventional one-cavity press, with the same molding cycle and pressure as required for molding single parts. Thermoplastic stack injection molds have been used since the mid 1960s. However, it was not until the late 1970s and the early 1980s that the stack molds

Figure 10-28 Fixed mini hot tip torpedo “DU-200”

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10 Injection Mold Design gained in popularity. Stork Plastics Machinery of Hengelo, Holland and Husky Corp. of Canada developed a unique thermoplastic stack injection molding machine and companion stack molds, used to mold double layers of the parts automatically, using the clamp force available.

Figure 10-29 Stack mold, four level or tandem mold (Courtesy: Husky Corp.)

The new thermoplastic stack injection molding process incorporates the specialized tooling required into the machine. This includes hot runner systems, valve gating, and rack and pinion gears for mold segment separation during clamp opening and closing. These thermoplastic stack injection molding process innovations are shown in Figure 10-29. The stack mold and this special machine is a turn-key molding system built by Husky Corp. in Canada.

10.9.11 Lost Core Thermoplastic Injection Molds These molds are used in the manufacturing of large, multifunctional, and hollow thermoplastic injection molded components as shown in Figure 10-30. This new technology represents a breakthrough for design engineers who have long searched for a method to produce such parts, meeting quality, dimensional, and repeatability demands in a high volume production environment. “Lost Core Molding System”

Figure 10-30 Lost core thermoplastic V-6 intake manifold (Courtesy: Ford)

A metallic core is the latest advancement in the metal alloy casting process. This technology is a proven foundry process and it has found acceptance in the plastic industry as a method for injection molding hollow thermoplastic components. The basic principal is to use a metal alloy to form a core insert that provides the internal shape (see Figure 10-31, left) while the mold surface cavity forms the external geometry of the thermoplastic part. Once the encapsulated component is formed (see Figure 10-31, middle), the inner core insert is melted-out, leaving the ﬁnished hollow component (see Figure 10-31, right). The lost core molding system combines the use of low melting point metal alloys to produce the metallic alloy cores, with the thermoplastic injection molding process using the metal cores as an insert. Once the thermoplastic component has been encapsulated around the metal core, the metal is melted-out via the fully integrated oil bath with a near total retrieval of the low melting point alloy for future reuse.

Figure 10-31 Lost core insert and thermoplastic intake manifold (Courtesy: Electrovert)

10.9 Types of Thermoplastic Injection Molds The metallic core process uses several metal alloys suited for the type of thermoplastic resin selected and application requirements. For engineering thermoplastic resins, a 42% tin and 52% bismuth alloy is used, having a melting temperature of 281 °F. The metallic cores can withstand the injection speeds, pressures, and temperatures of the thermoplastic injection molding process. The encapsulated assembly (thermoplastic component and metallic core) is placed in an inclined position inside a hot oil bath for two to three hours (meltout stage). The core is efﬁciently extracted from inside the part without harming the polymer’s physical properties, or causing warpage of the thermoplastic part. Many components have been commercialized using the lost core molding process. These include intake manifolds, water pump housings, power steering ﬂuid reservoirs, thermostat housings, fuel rail lines, and other automotive components. All these applications are physically viable and economical alternatives to vibration welded assemblies of two thermoplastic injection molded shells or metal systems (sand casting molds and machining). Lost Core Molding Design The lost core manufacturing procedure consists of three main processes: • Metallic alloy core forming • Injection molding to encapsulate the metallic core insert • Metallic alloy core melt-out The design of the lost core injection molding system consists of metal alloy core forming equipment and a vertical clamp type of injection molding machine as shown in Figure 10-32 (higher equipment cost) or a standard horizontal (lower equipment cost) type injection molding machine. The metal core is inserted horizontally from the back side using a loading robot. The encapsulated molded part with the metal core is horizontally removed from the working side of the molding machine using a second robot. Both parts are hung by the robot on a monorail to be conveyed to the hot oil tank where the metallic core is meltedout.

Figure 10-32 Lost core intake manifold injection molding machine

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10 Injection Mold Design Two weight scales on both sides of the injection molding machine (for the metal core and the encapsulated molded part) are used to detect and discharge faulty products automatically. Molding Machine Clamping System Besides the high strength, high cycle double-toggle mold clamping mechanisms, the metal core insertion and part removal operation is incorporated during mold opening and closing to shorten the molding cycle. The mold clamping force overcomes the polymer melt injection pressure; however, the mold size is much bigger than the mold cavity projected area, because of the complex design. Sometimes the mold weight may reach as much as 15,000 pounds. Rigid supports for the machine’s moving platen are needed to avoid damaging the mold parting line surfaces, because the clamping force plus the weight of the mold could over stress, deforming the machine moving platen. Therefore, the toggle clamp type system is more advantageous than a direct hydraulic or a hydro-mechanical clamp system. An injection molding machine with a large tie-bar clearance is needed. A tie bar puller is used as an effective countermeasure, allowing the use of smaller size molding machines. The most important function of the clamping system is to deliver stable and precise mold opening/stopping accuracy even at maximum speed, without affecting the hydraulic oil temperature or overloading the clamping system. Injection Unit For intake manifold applications, using a large shot size (product weight about 4.00 to 6.00 lbs), the injection capacity should be 1.5 to 2.0 times larger than the product. A precise injection process control with a stable high speed ﬁlling and instantaneous switch-over from ﬁll to packing conditions is important. Barrel, plastifying screw, free ﬂow check valve, and nozzle of wear resistant steels are recommended when processing nylon 6/6 resin reinforced with 30% to 35% ﬁber glass. In addition, a high compression injection screw with a compression ratio between 3.50 and 4.00 should be used. A screw length-to-outside diameter ratio between 20 and 24 is needed. Lost Core Thermoplastic Injection Molds A typical lost core thermoplastic injection mold used for the production of nylon intake manifolds (automotive four cylinder engines), is shown in Figure 10-33. These types of molds are very complicated to design, build, and operate because of the product design complexity and the critical and important functions required in the lost core thermoplastic injection molding process. The following mold design considerations are required in this process:

Figure 10-33 Lost core thermoplastic intake manifold mold (Courtesy: Ford)

• The low melting temperature metallic alloy cores have less strength than common steel materials; they are easily moved or bent by the hot melt of the polymer. The metal core is ﬁxed inside the mold with a locking pin, but if the pin reinforcement force is less than the thrust force generated by the melt injection pressure, the metal core will shift and the over molded part wall thickness will be nonuniform. Both ends of the metal core need to be

10.9 Types of Thermoplastic Injection Molds supported, but the thrust force generated by the injection pressure should be a compressive stress. • Nonuniform mold cavity ﬁlling during injection causes metal core deformation and nonuniform wall thickness of the parts due to the injection pressure. Type, dimensions, number, and location of the gates, melt and mold temperatures, injection speed and pressure distributions should be estimated by using a “Melt Flow Computer Analysis Program”, making corrections if needed, such as adding gates on the opposite side or increasing the number of gates to balance the melt injection forces that can deform the metal core insert. • The hydraulic actuators for the mold slides should be bigger with longer slide strokes, because this causes the core operating time to be increased, the simultaneous movement of the mold closing/opening operation is effective in shortening the cycle time. • Provisions should be made in the mold, so that the space needed for the pick-up can be secured without interfering with the robots and the mold during core insertion and part removal. • After insertion, the metal core is ﬁxed by a locking pin through the mold hydraulic slides, so when the stopping accuracy varies, the locking pin cannot be inserted. • The hydraulic ejector does not need a long stroke to push out a normally inserted metal core, but if the ejector pin does not properly return, the subsequent insertion cannot be done. The lost core thermoplastic molding process provides several advantages, among them the ﬂexibility in designing and manufacturing hollow thermoplastic components of varied internal geometries in one piece. The lost core molding process provides several design advantages, for example, the intake manifold increases engine horsepower by increasing air inlet speed through the channels (smooth surface ﬁnish), lower air inlet temperature (thermoplastic insulation), and multi-functional design The lost core molding process is applicable for several engineering thermoplastic resins. It is able to produce a textured surface in the internal walls of the thermoplastic components via a metallic core insert. It also produces good dimensional control of the thermoplastic components and metallic core inserts. It has proven to be economically viable in the production of one-piece thermoplastic components. This is especially true in medium to large volume production runs. The lost core molding process allows for the manufacturing of components that were not previously considered for injection molding as one piece thermoplastic parts, especially where the “value-added factor” of the component is enhanced. The process produces uniformly structured metallic cores with good surface ﬁnish, free from voids. Post machining of the core, before the injection molding process, is not normally required. The metal alloy used to manufacture the metallic cores is recycled nearly 100% using an integrated metal recovery system. The metal alloy cores are certiﬁed as having no health hazards. Since the alloy is reusable, there are no environmental disposal problems.

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10.10

Number of Mold Cavities

The mold cost increases with the number of cavities, while the molding process costs or machine use time decreases. The optimum number of cavities is the interception between the mold cost and the molding process as a function of the number of cavities. When the number of cavities is higher than the optimum number, the mold cost and the molding process costs affect the efﬁciency of the injection molding process. The cost of manufacturing of an injection molded product consists of three major elements: the cost of the thermoplastic materials, the cost of the molding process, and the cost of the mold. The thermoplastic material costs for a molded product are generally independent of processing variables. The remaining two costs, molding process and mold costs, are directly affected by many parameters, including the number of cavities per mold. Since the number of cavities used affects the cost of a molded product, considering the economics resulting from the number of cavities chosen is important. Too often, the cost of manufacturing is an uneconomical incentive when the mold cost is too high, resulting from too many cavities, or high molding process costs, resulting from too few cavities.

10.10.1 Cavity Number Limitations Often, using the economic optimum number of cavities may be impossible to implement. In these cases, the maximum number of cavities should be considered. Judgment resulting from mold fabrication experience can never be replaced by theoretical calculations. Incorporating the optimum number of cavities into a mold may be impractical due to such things as: • Complexity of runner, gating, or venting systems • Overall dimensions of mold base or tie-bar spacing • Projected area of cavities for clamping capacities • Core pulling or cam action mechanisms • Tolerances required on ﬁnished parts

10.10.2 Number of Mold Cavities Equation The number of cavities equation is designed to balance the mold investment against the molded product cost. The number of suggested cavities is based on an empirical rule. The yearly molding production rate of parts is divided into an ordering frequency of about every 50 days. The 50 days molding production requirements should be fabricated in 200 hours using a three shift molding operation.

Example 10-1 For example, if the yearly molding production requirement is 500,000 molded parts, the 50 days frequency requirement will be 100,000 parts. The estimated cycle is 60 shots per hour or 60 molded parts per cavity per hour.

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10.11 Mold Parting Line Parting surface

The number of cavities is: Number of cavities =

Molded parts needed to produce in 50 days (200 hours) × (molded parts /cavity /hour)

Molded product

100,000 molded parts needed in 50 days 200 hours × 60 parts per cavity per hour = 8.33 or eight (8) cavities =

This equation has been checked and has proven to be an accurate method to calculate the number of mold cavities. The important criteria for specifying a suitable tool steel, surface hardness, and ﬁnishing for the cavities is the severity of usage and the duration of product activity. Severe usage is deﬁned as a mold running for about half of the available time, whereas running for 10% of the available time is viewed as mild usage.

10.11

Cavity Seal-off surface

Mold moving half

Flat contact surfaces

Mold Parting Line

The parting lines of a mold are those portions of both mold plates, next to the cavities, which butt together to form a seal and prevent the loss of thermoplastic material from the mold cavities. We can classify the parting line as either ﬂat or non-ﬂat. The non-ﬂat types include stepped, proﬁled, and angled parting lines.

Molded part ejected Moving Half

Fix half

Figure 10-34 Flat mold parting line and cavity location Flat parting line

The ﬂat parting line is the simplest to manufacture and maintain. It can be surface ground and is easily bedded down in the mold cavity surface. To bed down a pair of mold plates is the process of marrying the two mold surfaces together. This is accomplished by blue marking one surface, momentarily bringing the two plates together and subsequently removing any high spots that will be apparent on the other surface. The plates are said to be bedded down when an even ﬁlm of blue marking is transferred from one plate to the other.

10.11.1 Flat Mold Parting Line

Cavity Undercuts cause ejection problems Poor mold design Flat parting line

The parting line depends entirely on the shape of the component. The cavity for the rectangular molded part in Figure 10-34 shows that the mold cavity can be die-sunk into one mold plate. The position of the parting surface will therefore be at the top of the molded part, and the parting line is a ﬂat surface. For appearance, this is the ideal arrangement as the parting line is not noticeable, unless ﬂashing molding problems develop in the injection molding process. The ﬂat parting line must be selected so that the molded part can be removed from the mold. Figure 10-35 shows a ﬂat rectangular cavity that incorporates a double beveled edge. The parting line for this component cannot be on its top surface, because this will create an undercut in the mold. The only suitable choice for the parting line is at the center of the double bevel. This parting line selection allows for half of the required form to be die-sunk into each of the two mold halves.

Cavity

Molded product Good mold design

Figure 10-35 Flat parting line selection affects mold performance

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Parting line

Molded part

Figure 10-36 Typical ﬂat parting line for different molded shapes Parting line

Figure 10-37 Proﬁled mold parting line

A number of molded shapes that permit a ﬂat parting surface to be adopted are shown in Figure 10-36. The arrows show the parting line on the molded parts.

10.11.2 Non-Flat Mold Parting Line Many molded parts require a parting line that lies on an angular or curved surface. In these cases, the parting line must be stepped, proﬁled, or angled. Proﬁled Mold Parting Line Molded part

Impractical ejection

A proﬁled parting line illustration is shown in Figure 10-37. The molded part is shown in the top illustration. It should be noted that, while in cross section the molding form is constant, the general form (side view) incorporates curves. As the edge of the component is square with the face (apart from cavity draft), the entire cavity geometry can be sunk into one mold cavity plate. Thus, the general form of the parting surface will follow the inside surface of the mold cavity. Angled Mold Parting Line The designer is frequently confronted with a component that, while regular in form, cannot be ejected from the mold cavity if a ﬂat parting line is adopted.

Poor mold design

Figure 10-38 shows the molded part on the left, the ejection problem in the middle, and the recommended angled parting line in the bottom illustration. Complex Mold Parting Line The complex parting lines consider other component’s edge forms (that is, either square, double-beveled, or incorporating a radius) where the edge form is not constant. This often leads to quite complex parting line surfaces. Figure 10-39 shows a simple product design geometry using a complex parting line.

Good mold design

Figure 10-38 Angled mold parting line

To determine the parting line, a number of cross sections are drawn through the part to be molded and the maximum dimension of each is deﬁned when viewed in the draw direction. The parting line will pass around all these points

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10.11 Mold Parting Line

Parting line

Parting line

Figure 10-39 Complex mold parting line

of maximum dimension. Once the parting line has been determined, the mold’s parting line surface can be drawn. This parting surface does not vary in the transverse direction and, if wanted, it can be machined completely across the mold’s surface. This surface is conﬁned to the area next to the cavity to reduce the time spent in bedding-down on the undulating surface to a minimum.

Parting line

Molded part

Parting line

Local Stepped Mold Parting Line On certain components it may be found that the change from one point of maximum dimension to another is quite abrupt. In these cases, the parting line is stepping from one plane to the other, preferably at an angle. Permitting the mold maker certain latitude in blending the cavity form in the step region to ensure that local undercuts do not occur is usually necessary. Incorporating a step or proﬁled parting surface to compensate for one or two small irregularities in an otherwise regular form is frequently necessary. Normally, this is best achieved by localizing the change in parting surface to permit most of the surfaces to be kept ﬂat. Consider a more complex example, as shown in a box shaped component with a lug on one of the side walls in Figure 10-40. This component can be molded in two ways: either by locally stepping the parting line surface or by designing a mold of the side cavity type. The simpler of the two methods is to step the parting surface locally and should be the method of choice when applicable. The molded part, cavity top view, core bottom view, and the part cross section through the mold are shown in Figure 10-40. The lug is above the general parting surface, which requires a projection from the core side to raise the level locally to the core parting surface. A complementary recess is die-sunk into the cavity plate to accommodate this projection. The projection is bedded into the recess on a slight taper. An angle (φ) of between 2° and 5° per side is suitable. While straight sides are simpler to produce and to bed down, ﬂashing problems can develop quickly because of wear occurring between the two sliding members. The recess in the cavity plate is locally deepened to form the lower part of the cavity for the lug. The top part is formed by the top surface of the core parting line.

Parting line Parting line

Cross section view

Parting line

Cavity top view

Parting line

Core bottom view Figure 10-40 Local stepped mold parting line

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10.11.3 Balancing of Mold Parting Line Surfaces Force

Arrows force

Force

When the parting surface is not ﬂat, there is the question of unbalanced forces to consider in certain instances. This is best shown in a mold with stepped parting line surfaces as in Figure 10-41. The thermoplastic melt, when under pressure within the cavities, will exert a force that will tend to open the mold in the lateral direction. If this happens, some ﬂashing may occur on the angled face. The movement between the two mold halves will be resisted by the support pillars; but even so, because of the large forces involved, balancing the mold by reversing the step is desirable so that the parting surface continues across the mold as a mirror image of the section that includes the cavity. Specifying an even number of cavities is often convenient (2, 4, 8, 16, 32 and 64) when considering this type of mold, as cavities positioned on opposite sides of the mold’s center line serve to balance the mold. When balancing is not practical because of the size, very sturdy support pillars must be incorporated behind the support plate of the mold.

Unbalanced parting line

10.12

Mold Ejection Systems

The injection molded parts are subjected to size reduction during the cooling cycle, caused by the mold shrinkage characteristics of the thermoplastic material. If parts with accurate dimensions are needed, allowances must be made for this shrinkage when establishing the dimensions for the cavities. The molded parts are ejected from the mold cavities after they have cooled down. Special mold manufacturing procedures are required for the ejection systems automatically to remove the runner system and the molded parts from the mold.

Balanced parting line

Figure 10-41 Balancing of mold cavity parting line surface forces

All mold surfaces that come in contact with the thermoplastic melt and the draft walls in the direction of the draw should be carefully polished. Lubrication of the mold surfaces that contact the thermoplastic melt is not recommended. There is a danger that scratches and cracks will develop when the mold is opened or the parts ejected, particularly when molding ﬁber glass reinforced resins. The mold cavities should have a mirror ﬁnish by grinding and polishing the areas in the direction of ejection to eliminate any scratches and indentations. Wherever possible, the surface should be lapped, since even microscopic scratches and indentations are ﬁlled with thermoplastic melt under high injection pressure, preventing smooth ejection. Ejection equipment is provided on the injection machine behind the moving platen for automatic actuation of an ejector system. Because of this, the mold’s ejector system will be most effectively operated if it is placed in the moving half of the mold. Only ﬂat parts with a pronounced draft toward the direction of ejection can be ejected without special ejector pins on the parts themselves. However, this requires the runner system, the gate, the sprue, and the molded part to have sufﬁcient strength to eject the molded parts from the center by means of the sprue puller ejector pin.

10.12 Mold Ejection Systems

10.12.1 Ejector Plate Assembly The ejector plate assembly is the part of the mold to which the ejector components are attached. The ejector assembly is contained in a pocket, directly behind the support plate. The assembly consists of an ejector plate, a retaining plate, and a knock-out bar. One end of the knock-out bar is threaded and then screwed into the ejector plate. In this design, the knock-out bar functions not only as an actuating member but also for guiding the assembly. The parallel portion of the knock-out bar passes through a bushing ﬁtted in the bottom holding plate of the mold.

10.12.2 Ejector Plate The purpose of this member is to transmit the ejector force from the actuating system of the machine to the molded part via an ejector component. The force required to eject the molded parts is high, particularly with those molded parts that are deep and incorporate little draft. Most ejector plates should be thick enough to resist deﬂecting to any signiﬁcant extent. Deﬂection tends to occur at the beginning of the ejector stroke when there is maximum adhesion between the molded part and the core. The deﬂection of any beam is inversely proportional to the cube of its depth and, therefore, a relatively small increase in plate thickness will decrease deﬂection of the plate. If an ejector plate does deﬂect to any extent, side forces are applied to the ejector components that result in increased wear in the mold plate holes or ejector sleeves and bent ejector pins. During the injection part of the cycle, with certain pin and sleeve type ejector systems, the melt pressure acts directly on the ejector component. H-13 tool steel (65–74 Rc hardened) is suitable for this application.

10.12.3 Retaining Plate The retaining plate is securely attached to the ejector plate by screws. Its purpose is to retain the ejector pins. The thickness of the plate is governed by the depth of the heads of the ejector pins it retains. For small molds the retaining plate is made to the same general dimensions as the ejector plate. For larger molds, however, holding one or most of the ejector pins is convenient from the mold making viewpoint to incorporate local retaining plates.

10.12.4 Ejector Sleeves Ejector sleeves are used in cases where containers and tubular parts cannot be ejected by means of a stripper plate. To prevent the ejector sleeves from making scratches on the polished mold surfaces, which would interfere with smooth ejection, the external and internal diameters are made 0.004 to 0.007 in smaller or larger than the respective mold surfaces. During the ejection process, they slide by without making contact with the interior and exterior. Since, for the most part, the ejector sleeves are thin-walled tubes, which are difﬁcult to harden without deforming them, it is advisable to obtain ejector sleeves made of high strength tool steel (to 65–74 Rc). They should be hardened and tempered before ﬁnishing.

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10.12.5 Types of Mold Ejection Systems When the melt cools off inside the mold cavities, it contracts by an amount depending on the material being processed. For a molded part that has no internal form, for example a rectangular shape, the molded part will shrink away from the cavity external walls, around the internal cores, thereby permitting a simple ejection technique to be adopted (for example, a valve jet of air). However, when the mold cavity has an internal form, the thermoplastic melt, as it cools, will shrink onto the core and a positive type of ejection system is necessary. The mold designer has several ejection systems from which to choose, but, usually, the choice will be restricted depending upon the shape of the molded part. The basic mold ejection systems are the following: • Ejection pin • Ejection sleeve • Ejection blade • Air poppet valve ejection • Mechanic valve ejection • Stripper plate or ring ejection Mold Ejector Pin System Molded part Ejector pin Ejector sleeve

Ejector plate

Figure 10-42 Mold ejector pin system

The mold ejector pin system is the most common type of ejection because it is generally the simplest to incorporate in a mold design. With this mold ejection system, the molded product is ejected by the application of a force by a circular steel rod called an ejector pin. Figure 10-42 shows the principle of operation. The ejector pin is in the moving half of the mold and it is held back by the return pins (not shown). In operation, the ejector plate assembly, to which the ejector pin is attached, is moved forward relative to the mold cavity plate, causing the ejector pin to push out (eject) the molded part from the mold cavity. The working diameter of the ejector pin must be a good slide ﬁt in its mating hole in the mold plate. If it is not, the thermoplastic melt will creep through the clearance and a mass of material will progressively build up behind the mold plate. The rear part (head) of the ejector pin is ﬁtted into a suitable hole that is bored and counter bored in the ejector retaining plate. The rear face of the ejector pin head is backed up by the ejector plate. This arrangement must allow the ejector pin to ﬂoat. Mold Ejection Blade System The main purpose of the mold ejection blade system is the ejection of very slender parts, such as ribs and other projections, which cannot satisfactorily be ejected by the standard mold ejector pin system. A blade is basically a rectangular ejector pin. While the blade ejector can be machined from a solid rod, it is usually fabricated by inserting a blade of steel into a slot machined into a standard type of ejector pin. The blade may be pinned or it may be brazed. The advantage of the two part construction is that the blade can easily be replaced should be it become damaged.

613

10.12 Mold Ejection Systems The blade ejector element is ﬁtted to the ejector assembly in an identical manner to the one described for the standard ejector pin. Figure 10-43 shows a cross sectional view through the mold that incorporates the ejection blade system. The rectangular ejection blades are ﬁtted in shaped holes in the mold “B” cavity plate. If this slot is to be machined into a solid mold plate, it presents a costly endmilling operation. Easing the machining of this aperture by using a commercial ejector blade subassembly is desirable. Therefore, a blade ejector can be used on the edge of a molded part. A slot is machined completely through the shoulder portion. Then the core insert is ﬁtted into a suitable recess in the mold cavity plate forming a rectangular aperture, into which the ejector blade slides.

Molded part Ejector blade

Ejector plate

Figure 10-43 Mold ejection blade system

Mechanical Ejection Valve System The valve ejector system is basically a large diameter poppet ejector pin. Valve type ejection is normally used for the ejection of relatively large components in situations where using standard parting surface pins is impractical. It is also used as an alternative to stripper plate ejection in certain situations. The valve type ejector system is designed to apply the ejector force to the inside surface of the cavities as shown in Figure 10-44. We have previously noted (for pin ejection) that this is an undesirable practice because of the possibility of distorting the molded parts during ejection. However, with the mechanical valve ejection, the risk of the molded parts being distorted is reduced, because the poppet has a large effective ejection area. The ejector element incorporates a valve type head (poppet) that is seated in a nest machined into the cores as shown in Figure 10-44. To ensure a leak free joint, a small parallel portion is provided at the major diameter of the valve. This is a sliding ﬁt in a complementary parallel recess in the cores. Because the ejector element forms a relatively large part of the cavity, a large force will be applied to the element by the melt tending to move it rearwards. The head must therefore be carefully bedded into its nest in the core insert block. In operation, when the mold opens, the shank of the valve ejector element is moved by the knock-out bar of the injection machine and the molded parts are ejected. Immediately the mold closing stroke commences the return pins causing the valve to be returned to its seating. For fully automatic operation it is normal practice to use a jet of air to blow the molded part off the valve as there is a tendency in certain instances for the molded parts to stick to the ejection valves. For a multi-cavity mold, an ejector assembly and an ejector grid must be used. The valve element is attached directly to the ejector plate by screws. A gap of 0.125 in minimum should be incorporated behind the ejector plate to ensure that both valves are seated in their respective nests. The return pins, not shown, are incorporated to return the ejector plate and valve elements to their rear positions as the mold is closed. Air Poppet Valve Ejection System With this ejection method the ejector force is provided by compressed air, which is introduced directly onto the lower cavity face via a small air ejector valve. For this method to operate efﬁciently, the adhesion between the cavity wall and the core must be broken locally to permit the compressed air to be introduced. This is achieved by causing the valve ejector to move forward slightly (by air pressure).

Figure 10-44 Multi-cavity mechanical ejection valve system

614

10 Injection Mold Design Air poppet valve open Molded part

Compressed air inlet

Figure 10-45 Air poppet valve ejection system

The effective ejector force depends on the pressure of the compressed air and the area on which it acts. Therefore, the larger the area of the component to be ejected, the greater the ejector force. There is, however, one drawback in that, while the compressed air ﬁnds an escape route, the effective ejector force is rapidly diminished. This method of ejection is particularly suitable for box-type components, where the side walls act as a seal during a major part of the ejection stroke, preventing the escape of the compressed air. Figure 10-45 shows an air poppet valve ejector element. The head nestles in a complementary shaped recess in the core insert, while the shank of the element is a sliding ﬁt in a hole next to the nest. To provide a passage for the compressed air, small slots are machined longitudinally on the periphery of the hole. The element is held onto its seating by means of a compression spring. This spring is held in position by a nut. To prevent the loss of compressed air from the rear side of the core insert, a sealing plate and associated “O” ring are incorporated. A port connects the air ejector valve to the compressed air supply. The compressed air can be introduced either manually or automatically. With manual control, the machine operator controls the two way valve at any suitable point in the cycle. With automatic control, a cam is ﬁtted onto the machine bed so that at the required point in the opening stroke, the valve spool is “ON” and compressed air ﬂows into the mold. Immediately, the roller (attached to the spool) runs off the cam at either end and the compressed air supply is cut off. The operational sequence of the air poppet ejection valve system is as follows: when the control valve is opened, compressed air passes into a chamber. A force is applied to the air valve ejector that overcomes the force exerted by the compression spring and the valve ejector leaves its seating. At this point, the adhesion between the molded part and the core joint surfaces is broken in the area next to the air valve ejector. Compressed air is now free to ﬂow from the chamber through the slots, through the open valve and into the space between the molded part and the core, ejecting the molded part. Air Poppet Valve Ejection System Advantages • No ejector grid or ejector assembly is required; lower mold costs • Can be ﬁtted in either mold half. It is a suitable method of ejecting box-type components that are to be center gated on the inside • The ejection can be performed anytime during the mold opening • The air valve ejector element acts as a vacuum breaker between the molded part and the core surfaces. This system should work in coordination with other ejection systems to complete the ejection. Stripper Plate Ejection System In very deep, hollow parts, which shrink very ﬁrmly onto the cores, individual ejector pins are no longer adequate, because the individual ejector pins are likely to pierce the bottom of a deep container. Such cases require ejection by means of a stripper plate. The stripper plate covers the entire rim of the container and pushes the molded part off the core as the ejector rod actuates the ejection mechanism. Locating this stripper plate on the stationary half of the mold to help in loosening certain parts when the core has to be located on the stationary half of the mold is also possible.

615

10.13 Injection Mold Cooling Figure 10-46 shows the principle of operation used in this stripper plate ejection technique. The stripper plate is mounted between the cavity plate and the core cavity plate. The aperture in the stripper plate is a sliding ﬁt on the core. When the mold starts to open, the stripper plate moves back with the core plate. The molded part sticks to the core because of the shrinkage forced during the cooling cycle. Once the molded part is clear of the cavity, the movement of the stripper plate is initiated, while the core plate continues the rearward movement. The molded part is then withdrawn from the core by the stripper plate and the molded part ejected. The mold is then closed in preparation for the next molding cycle. The stripper plate depth must be relatively thick; the depth dimension must be sufﬁcient to resist the bending forces applied to the stripper plate as the molded part is ejected. The lengths of the support pillars are generally longer than what is speciﬁed for a corresponding two-plate mold. The reason is that the support pillars must support the stripper plate over the complete ejection stroke and also the projected mold moving half assembly when the mold is opened. The leader pin’s dimensional tolerances must be precise enough to enter the mold’s ﬁxed half shoulder bushings during mold closing time, and to ensure precision alignment between the cavity and the core.

10.13

Injection Mold Cooling

One fundamental principle of the thermoplastic injection molding process is that hot melt enters the mold cavity, where it cools rapidly to a temperature at which it solidiﬁes sufﬁciently to retain the shape of the cavity. While the melt ﬂows more freely in a mold cavity, a longer cooling period is required before the solidiﬁed molded parts can be ejected. On the other hand, while the melt solidiﬁes quickly in a cold mold, it may not reach the extremities of the cavity. A compromise between the two extremes must be made to obtain the optimum molding cycle. The thermoplastic injection molding process embodies the technological ability, innovation, and efﬁciency required for injection molding a product, while maximizing the amount of proﬁt. The thermoplastic injection molding process efﬁciency is affected by the mold cooling design. The difference in productivity between a correct and incorrect mold cooling design can represent an increase of 20 to 40% in the molding process costs. The term mold cooling means lowering the temperature of the thermoplastic melt in the cavity to form a molded product. When molds require heat for proper operation they are being practically cooled as the temperature of the mold is lower than the thermoplastic melt temperature. The heat transfer ﬂows from a high temperature source to a contacting element of lower temperature. The main source of heat removal or mold cooling is obtained by an adequate circulation temperature control and volume control of the cooling ﬂuid. If an adequate amount of properly treated water is available at any temperature and volume required by the molding process, the mold will be properly cooled.

Figure 10-46 Stripper plate ejector system

616

10 Injection Mold Design Heat removal depends on temperature, pressure, viscosity, thermal diffusivity, and thermal conductivity. The heat transfer calculations are based on steady state or equilibrium conditions. In the thermoplastic injection molding process, the temperature, pressure, and viscosity are constantly changing as the melt ﬂows and cools in the mold cavity. The complexities of the shapes of the molded products are beyond analytical determination. The molecular weight, molecular structure, and distribution of the thermoplastic melt are not constant. In spite of all these variables, several mold cooling computer analysis programs have been developed by making many heat transfer assumptions to simulate the mold cooling process and the thermal behavior of the thermoplastic melt. These mold cooling programs are approximations and have some technical value in the development of new products and for a novice engineer who is learning thermoplastic injection mold cooling technology. During the cavity ﬁlling stage, the hottest material will be near the entry point, i.e., the gate, and the coolest material will be at the point farthest from the entry. The temperature of the coolant ﬂuid, however, increases as it passes through the mold. Therefore, to achieve an even cooling rate over the molding surface, it is necessary to place the incoming coolant ﬂuid next to the mold cavity surfaces. Ultimately, adopting the idealized approach is not always practical and the mold designer must use a fair amount of common sense when laying out coolant circuits to avoid unnecessarily expensive molds. The layout of a circuit is often complicated by the fact that the cooling channels must not be drilled too close to any other hole in the same mold plate. The mold plate has several holes or recesses, to ﬁt ejector pins, support pillars, guide bushings, sprue bushing, cavity, and core inserts, etc. To obtain the best possible position for a mold cooling circuit, it is good practice to lay the circuit in at the earliest opportunity in the mold design. The other mold components such as ejector pins, bushings, vents, etc., can then be positioned accordingly.

10.13.1 Mold Temperature Control The thermoplastic injection molding process requires the rapid removal of heat from the mold cavities, so that the molded parts can be removed from the mold in the shortest time and in a condition that the parts meet the quality control requirements. The mold temperature control system includes the mold, the cooling channels, the different mold cooling systems used in the mold design, the type of ﬂuid with adequate capacity for its circulation, and the method of temperature control. Predicting the best temperature conditions for a given mold and thermoplastic material is neither possible nor necessary. In many applications there will be several different temperatures maintained for different mold components to meet the quality control speciﬁcations. These process settings should be evaluated to select the best balance between economics and part quality. The mold cavity surface temperature is measured at the beginning of the molding cycle, for example 140 °F. When the hot thermoplastic melt is injected into the mold cavity, the mold temperature increases to 160 °F. At the end of each molding cycle, the mold temperature decreases to 140 °F. A constant mold temperature means that the amount of heat removed per shot by the total cooling system is the same as the amount of heat provided by the hot thermoplastic melt. Mold

10.13 Injection Mold Cooling temperature ﬂuctuations will affect the dimensional control of the molded part, warpage, ﬂashing, surface ﬁnishing, and reduce physical properties. The mold cavity reaches its maximum temperature very quickly and the mold cooling stage of the cycle is used to reduce this temperature to the base operating temperature. This rate depends on the temperature differences, the area of the cooling surfaces in the mold and in the cooling channels, the transfer rates of the heat from the thermoplastic melt to the thermoplastic/ metal interface through the metal and through the metal/water interface. The heat removed by the mold and radiated into the operating molding area is relatively constant. It is important that a mold cooling system produces a uniform temperature over the entire surface of the mold cavity at the same level as the cooling channels. It will remove heat from the thermoplastic melt at the highest possible rate consistent with the quality and properties required in the molded part.

10.13.2 Factors Affecting Mold Cooling The mold temperature is affected by several thermoplastic injection molding process factors related to mold cooling: • Thermoplastic material (process melt temperature, crystallization rates, modulus of elasticity) • Part wall thickness, size, complexity, dimensional control, and ﬁnish • Shot weight, process automation, cooling time • Mold base, cavity, and core material • Size and shape of the mold, cavity, and core • Efﬁciency of mold cooling systems • Size and location of cooling channels • Velocity, capacity, pressure drop, and temperature of cooling ﬂuid • Operating ambient conditions (temperature, moisture, and air ﬂow)

10.13.3 Effects Caused by Elevated Mold Temperature If the mold temperature is too high, the molding cycle time increases, reducing the efﬁciency of the injection molding process. The thermoplastic melt remains in the ﬂuid stage longer with higher mold temperatures, causing the melt to ﬂash through the parting line vents. Higher thermoplastic melt ﬂow rates cause burn marks on the molded parts and steel corrosion of the cavities. Even if the melt does not burn, the inadequate venting may seriously reduce the physical properties of the molded part at the weld lines. With a hot mold, the gate will remain open longer, permitting more melt to be packed into the cavity. The excessive ejection forces may lead to moldedin stresses and failure of the molded part during ejection or in service. This extra injection pressure results in forcing the thermoplastic melt more tightly against the cavity surface, requiring higher ejection forces that may lead to ejection problems. When ejecting a hot molded part, the ejector pins may cause ejection marks, increasing the stress or causing perforation of the part

617

618

10 Injection Mold Design surface. Over-packing the cavities will increase the deformation (deﬂection) of the mold steel plate that may cause shifting, dimensional control, and other ejection problems. When hot thermoplastic melt hits the cold mold cavity surface, the melt freezes against the cavity wall, orienting the molecular structure in the direction of the ﬂow. Orientation leads to molded-in stress of the molded parts; they have a higher tensile strength in the direction of the ﬂow, because they have more carbon-tocarbon molecular linkages than in the direction perpendicular to the melt ﬂow. The carbon-to-carbon linkage forces are stronger than the weaker electrostatic bonds that are predominantly perpendicular to the direction of the melt ﬂow. The sprue performance is also affected by the temperature of the mold (too hot). The sprue requires a longer cooling time to become solid, a hot sprue will break or stick, causing cycle interruptions. Too high a mold temperature causes excessive shrinkage, sink marks, and voids. If one side of the mold overheats, the thermal expansion of that side may cause the leader pins to seize in the bushings; this condition can pull the mold off from the molding machine platen.

10.13.4 Effects Caused by Too Low a Mold Temperature Cold molds cause processing problems, such as incomplete molded parts, poor surface ﬁnishing, oversize parts, uneven density distribution causing severe molded-in stresses that may lead to immediate or premature failure. A cold mold causes the melt to freeze before it ﬁlls the cavity giving an incomplete shot. If it does ﬁll, there may not be enough melt in the cavity causing voids, sink marks, splay marks, poor weld lines, poor mold shrinkage control, and reduction of the physical properties of the part causing brittleness, poor temperature resistance, and poor dimensional stability at elevated temperatures. These are just some major problems caused by the mold temperature conditions being too cold. These are all compounded if the cooling system lacks the capacity or instrumentation to maintain a consistent operation.

10.13.5 Mold Heat Transfer Methods There are three methods for transferring heat in the injection molding process: radiation, convection, and conduction. The hot thermoplastic melt from the plastifying unit nozzle is transferred into the mold sprue bushing. The heat from the thermoplastic melt moves by convection through the polymer until it reaches the cavity surface of the mold. The heat is then conducted through the mold cavities to the mold cooling system and the cooling channels. The heat is then transferred to the cooling ﬂuid. A substantial amount of the heat reaches outside the mold surface, where the heat is lost by radiation. Of the heat added by the plastifying unit to the thermoplastic melt in the injection molding process, approximately 60% is removed by the mold; however, some heat is left in the molded parts after they are removed from the mold cavities after ejection. Approximately 35% of the heat is removed by radiation from the mold and 25% is removed by the cooling ﬂuid that controls the mold cavity surface temperature. The percentages vary, depending on the thermoplastic material, melt process temperature, mold temperature control efﬁciency, molding cycle, molded product wall thickness, application requirements, size and complexity of the product design.

619

10.13 Injection Mold Cooling 10.13.5.1 Cooling Channel Diameter Injection molds with a cooling channel circuit require seals in the connections to prevent leakage of the cooling ﬂuid. The presence of cooling ﬂuid on the surface of the mold cavities will cause defects in the surface and reduction of the physical properties of the molded part. Pipe threads are used to connect the cooling channels with the other components and this is an effective and economical type of seal. The cooling channel diameter is related to the size of the pipe ﬁtting inside diameter; the point of a change in cross section is largely problem free. With inserted cores having drilled cooling channels, such sealing problems are avoided with threaded pipes. Figure 10-47 shows a cross section assembly of a pipe ﬁtting/mold plate and a table for the cooling channel diameters per nominal pipe sizes. Nominal pipe size (in)

Tap drill used (in)

Channel dia. “D” (in)

Surface area “A” (ft2)

1/8

5/16

0.3125

0.982

1/4

7/16

0.4375

1.374

3/8

9/16

0.5625

1.767

1/2

11/16

0.6875

2.160

10.13.5.2 Mold Cooling Channel Location The rate of heat removal from the mold will vary directly with the thermal conductivity of the material used in the construction of the mold components. Table 10-5 states the values of various materials used for mold construction. From this table, we can see that beryllium copper will remove heat approximately six times faster than tool steel and eight times faster than stainless steel. Table 10-5 Thermal Conductivity of Mold Materials

Mold material

Description

Thermal conductivity (BTU/h/ft2/°F/in)

1020 Alloy steel

Hot rolled steel

20.00

4135 Alloy steel

Medium carbon steel

24.70

P-20 Steel

Tool steel

20.00

S-7 Steel

Alloy tool steel

21.00

H-13 Steel

Alloy tool steel

16.30

420 Alloy steel

Stainless steel

14.40

440 Alloy steel

Stainless steel

14.00

Kirksite

Cast low Carbon steel

62.00

Brass

Brass casting

70.00

Moldmax® copper alloy

Beryllium copper

60.00

Ampco 940 copper alloy

Ampco copper

125.00

Protherm® copper alloy

Beryllium copper

145.00

2024 Aluminum alloy

Aluminum

70.00

6061 Aluminum alloy

Aluminum

96.00

7010 Aluminum alloy

Aluminum

92.00

7075 Aluminum alloy

Aluminum

75.00

Cooling channel diameter "D"

Mold plate

Pipe fitting

Figure 10-47 Cooling channel related to nominal pipe size

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10 Injection Mold Design Cavity surface temperature 140° F. 140° F. 139° F.

B

D = 0.687 in. Spacing = B x 3 = 2.00 inch

Cavity surface temperature 122° F. 122° F. 113° F.

B

D = 0.312 in. Spacing = B x 4 = 1.25 inch

Figure 10-48 Surface temperature proﬁle per channel diameter, location, spacing, and cooling ﬂuid temperature

10.13.5.3 Cavity Surface Temperature Proﬁles Figure 10-48 shows the surface temperature proﬁles, based on the cooling channel diameters and spacing. The minimum cooling channel location in reference to the cavity surface for maximum cooling to maintain the integrity of the mold varies with the modulus of elasticity of the cavity insert material. For example, if the cavity insert is made from tool steel, the edge distance from the cavity surface to the tangent of the cooling channel diameter (B) should be equivalent to the cooling channel diameter (B). If the cavity plate is made from beryllium copper, this distance (B) should be 1.50 diameters. The reason is that the modulus of elasticity of tool steel is 30,000,000 psi, while that of beryllium copper is only 20,000,000 psi. Therefore, the distance from the cooling channel to the cavity surface is 50% more for beryllium copper than for tool steel. This reduces the effective cooling advantage of beryllium copper over tool steel to a factor of two rather than three. The closer the cooling channel is to the thermoplastic melt the higher the rate of heat removal. Drilling long cooling channels requires accurate machining. If the hole wanders too close to the surface edge of the cavity, the plate may collapse under molding pressures. If the spacing between the cooling channels and the cavity surface varies, so will the cooling rate. The very low thermal conductivity of the thermoplastic materials prevents the rapid transfer of heat from inside the molten thermoplastic wall to the external surface where it can move by conduction through the metal of the mold into the cooling ﬂuid. 10.13.5.4 Spacing Between Cooling Channels The smaller the spacing, the more uniform the temperature in the mold cavity and the higher the temperature increases at the surface of the cavity during injection. It is recommended that the spacing between the cooling channels be between three and four cooling channel diameters (D). If, however, the design requires a larger spacing between cooling channels, the spacing (B) from the cavity surface and the channel diameter (D) must also increase. With increasing wall thickness, more heat must be removed from the mold, requiring an increase in the cooling channel diameters (D).

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10.13 Injection Mold Cooling 10.13.5.5 Mold Heat Removal Analysis Some factors that affect the rate and amount of heat transfer are the thermoplastic material, size, geometry, and wall thickness of the molded parts, the melt ﬂow rate, melt viscosity, speciﬁc heat, thermal conductivity, density, molding process, and mold temperature operating conditions. How much heat is transferred in the mold is calculated by using a complex set of equations to obtain the values for the Reynolds number, Prandtl number, Nusselt laminar ﬂow number, Nusselt turbulent ﬂow number, heat transfer coefﬁcient, and heat transfer rate. Reynolds Number Osborne Reynolds discovered some basic principles of ﬂow in the 18th century. He took a container with a source of inlet water that kept it always full. A pot of dye was fed externally into a tube. Water could be bled from this container at a controlled rate through an outlet valve. When the ﬂow rate in the tube was low, the dye assumed a straight path parallel to the tube. Reynolds called this ﬂow rate condition laminar ﬂow. As the ﬂow increased, mixing of the dye began to occur. When ﬂow became very rapid there was a complete mixing of the dye in the tube. This he called turbulent ﬂow. He derived a formula to describe this phenomenon for a circular tube using a dimensionless number now known as the Reynolds number. In a laminar ﬂow, the ﬂuid will glide along the walls, unlikely to collide with the walls or other molecules or to distract the ﬂow from the direct straight line path. In a turbulent ﬂow, the ﬂuid molecules are moving with much higher speed and slam into the walls, giving better heat transfer characteristics. It is desirable to have a Reynolds number higher than 3,500 as the minimum starting point to control the temperature of the mold. For example, the turbulent ﬂow rate of cooling water at 50 °F has similar cooling characteristics to the laminar ﬂow rate of water at 32 °F. Reynolds Number Equation Re =

v×D× ρ 12 × μF

(10-1)

Prandtl Number Equation Pr =

3,600 × c ρ × μF

(10-2)

K

Nusselt Laminar Flow Number Equation D   Nu = 1.86  Re × Pr ×   12 × L 

0.33

 μC   μ  F

0.14

(10-3)

Nusselt Turbulent Flow Number Equation Nu = 0.116 × Pr

0.33

(Re

0.66

 − 125) 1 + 

 D    12 × L 

 μC      μF 

0.66 

0.14

(10-4)

622

10 Injection Mold Design Heat Transfer Coefﬁcient Equation U =

Nu × k × 12 D

(10-5)

The heat transfer rate is equal to the heat transfer coefﬁcient, multiplied by the exposed surface cooling area, multiplied by the difference in temperature between the mold cavity and the coolant ﬂuid. Heat Flow Rate Equation Q = U × A × L × (TC − TF )

(10-6)

Re = Reynolds number Pr = Prandtl number Nu = Nusselt ﬂow number L = Cooling channel length (ft) v = Cooling ﬂuid velocity (ft/s) D = Cooling channel diameter (in) ρ = Cooling ﬂuid density (lb/ft3) μF = Cooling ﬂuid viscosity (lb-s/in2) μC = Cooling ﬂuid viscosity at cavity temperature ((lb-s/in2) k = Cooling ﬂuid thermal conductivity (BTU/h/ft2/°F/in) cρ = Cooling ﬂuid speciﬁc heat (BTU-lb-°F) TC = Surface cavity temperature (°F) TF = Cooling ﬂuid temperature (°F) A = Exposed surface cooling area (ft2) = (π × D) / 12 U = Heat transfer coefﬁcient (BTU/ft2-h-°F) Q = Heat ﬂow rate (BTU/h)

0.00554 0.000878 0.000515 0.000456

Density ρ (lb/ft3)

66.45999 62.3799 62.1677

62.000

Thermal conductivity K (BTU/h/ft2/°F/in)

0.240

0.333

0.357

0.363

Speciﬁc heat Cρ (BTU/ft2-h-°F)

0.760

1.0490

0.9980

0.9980

Prandlt number Pr

63.16

9.97

5.18

4.15

EG = Inhibited ethylene glycol ﬂuid

Water 150 °F

Viscosity μ (lb-s/in2)

50% EG 50% Water 150 °F

Water 100 °F

Water 85 °F

Water 50 °F

50% EG 50% Water 32 °F

Table 10-6 Properties of Circulating and Cleaning Cooling Fluid

0.001

0.000287

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10.13 Injection Mold Cooling From the heat transfer rate (Eq. 10-6) the conditions to increase the rate can be deducted: • Increase the difference in temperature between the mold cavity surface and the mean cooling ﬂuid • Improve the efﬁciency of the heat transfer coefﬁcient • Increase the exposed surface cooling area. There are two ways to increase the difference in temperature between the thermoplastic melt and the cooling ﬂuid. One way would be to add more heat to the thermoplastic melt, this approach is self defeating, increasing the molding cycle time. Lowering the mold cooling ﬂuid temperature may cause molding problems. Increasing the exposed surface cooling area channels and making them as large as possible, is very beneﬁcial. The limitations are that large-diameter channels will reduce the cooling ﬂuid velocity and decrease the Reynolds number. Using a greater number of small diameter channels will increase the pressure drop in the cooling ﬂuid system or lead to a reduction of ﬂuid that can be passed through the cooling system.

Example 10-2 Determine the Reynolds number and the heat ﬂow rate of an injection mold required to operate with a mold cavity temperature of 150 °F using water at 50 °F as cooling ﬂuid at a ﬂow velocity of 0.50 (ft/s), through a cooling channel 1 ft long with a diameter of 0.4375 in behind the cavity insert. Reynolds Number v×D× ρ 12 × μF 0.5 × 0.4375 × 62.3799 = = 1,295.14 (Laminar Flow) 12 × 0.000878

Re =

(10-1)

Prandlt Number Pr =

3,600 × c ρ × μF

k 3,600 × 1.049 × 0.000878 = = 9.956 0.333

(10-2)

Nusselt Laminar Flow Number D   Nu = 1.86  Re × Pr ×   12 × L 

0.33

 μC   μ  F

0.4375   = 1.86 1295 × 9.956 ×   12 × 1.0 

0.33

0.14

 0.000278    0.000878 

(10-3)

0.14

= 16.920

624

10 Injection Mold Design

Heat Transfer Coefﬁcient Nu × K × 12 D 16.920 × 0.333 × 12 = = 145.54 (BTU/ft 2 -h-°F) 0.4375

U =

(10-5)

Heat Flow Rate Q = U × A × L × (TC − TF )

(10-6)

π × D 3.1416 × 0.4375 = = 0.114 (ft 2 ) 12 12

A=

Q = 145.54 × 0.114 × 1.0 × (150 − 50) = 1,659.156 (BTU/h)

Example 10-3 Determine the Reynolds number and the heat ﬂow rate of an injection mold required to operate at a mold cavity temperature of 150 °F using water at 50 °F as cooling ﬂuid at a ﬂow velocity of 2.00 (ft/s), through a cooling channel 1 ft long with a diameter of 0.4375 in behind the cavity insert. Reynolds Number v×D× ρ 12 × μF 2.0 × 0.4375 × 62.3799 = = 5,180.56 (Turbulent Flow) 12 × 0.000878

Re =

(10-7)

Prandlt Number Pr =

3,600 × c ρ × μF

k 3,600 × 1.049 × 0.000878 = = 9.956 0.333

(10-8)

Nusselt Turbulent Flow Number Nu = 0.116 × Pr

0.33

(Re

0.66

 − 125) 1 + 

 D    12 × L 

= 0.116 × 9.9560.33 (5,180.560.66 − 125) 0.66 0.14   0.4375    0.000278  × 1 +      12 × 1.0    0.000878    = 56.487

 μC      μF 

0.66 

0.14

(10-9)

625

10.13 Injection Mold Cooling

Heat Transfer Coefﬁcient Nu × k × 12 D 56.487 × 0.333 × 12 = = 515.936 (BTU/ft 2 -h-°F) 0.4375

U =

(10-10)

Heat Flow Rate Q = U × A × L × (TC − TF )

(10-11)

π × D 3.1416 × 0.4375 = = 0.114 (ft 2 ) 12 12

A=

Q = 515.936 × 0.114 × 1.0 × (150 − 50) = 5,881.67 (BTU/h)

Example 10-4 Determine the Reynolds number and the heat ﬂow rate of an injection mold required to operate at a mold cavity temperature of 150 °F using 50% additive and 50% water at 32 °F as cooling ﬂuid at a ﬂow velocity of 2.00 (ft/s), through a cooling channel 1 ft long with a diameter of 0.4375 in behind the cavity. Reynolds Number v×D× ρ 12 × μF 2.0 × 0.4375 × 66.4599 = = 874.73 (Laminar Flow) 12 × 0.005540

Re =

(10-12)

Prandlt Number Pr =

3,600 × c ρ × μF

k 3,600 × 0.760 × 0.005540 = = 63.156 0.240

(10-13)

Nusselt Laminar Flow Number D   Nu = 1.86  Re × Pr ×   12 × L 

0.33

 μC   μ  F

0.14

0.4375   = 1.86  874.73 × 63.156 ×   12 × 1.0  = 28.860

0.33

 0.001    0.00554 

0.14

(10-14)

626

10 Injection Mold Design

Heat Transfer Coefﬁcient Nu × K × 12 D 28.860 × 0.240 × 12 = = 189.981 (BTU/ft 2 -h-°F) 0.4375

U =

(10-15)

Heat Flow Rate Q = U × A × L × (TC − TF )

(10-16)

π × D 3.1416 × 0.4375 = = 0.114 (ft 2 ) 12 12

A=

Q = 189.981 × 0.114 × 1.0 × (150 − 32) = 2,555.62 (BTU/h)

Example 10-5 Determine the Reynolds number and the heat ﬂow rate of an injection mold required to operate at a mold cavity temperature of 150 °F using water at 50 °F as cooling ﬂuid at a ﬂow velocity of 2.00 (ft/s), through a 4.00 ft long cooling channel with a diameter of 0.4375 in and behind the cavity. Reynolds Number v×D× ρ 12 × μF 2.0 × 0.4375 × 62.3799 = = 5,180.56 (Turbulent Flow) 12 × 0.000878

Re =

(10-17)

Prandlt Number Pr =

3,600 × c ρ × μF

k 3,600 × 1.049 × 0.000878 = = 9.956 0.333

(10-18)

Nusselt Turbulent Flow Number  Nu = 0.116 × Pr 0.33 (Re0.66 − 125) 1 + 

 D    12 × L 

= 0.116 × 9.9560.33 (5,180.560.66 − 125) 0.66 0.14   0.4375    0.000278  ⋅ 1 +    12 × 4.0    0.000878   

= 53.160

 μC      μF 

0.66 

0.14

(10-19)

627

10.13 Injection Mold Cooling

Heat Transfer Coefﬁcient Nu × K × 12 D 53.160 × 0.333 × 12 = = 485.548 (BTU/ft 2 -h-°F) 0.4375

U =

(10-20)

Heat Flow Rate Q = U × A × L × (TC − TF ) A=

(10-21)

π × D 3.1416 × 0.4375 = = 0.114 (ft 2 ) 12 12

Q = 485.548 × 0.114 × 1.0 × (150 − 50) = 22,140.98 (BTU/h)

10.13.5.6 Heat Transfer Barriers Clean circulating ﬂuid, mold cooling system, and cooling channel conditions are very important for mold temperature control. To avoid internal corrosion in the mold cooling channels, it is highly recommended to coat the channels with electroless nickel or an equivalent coating before the mold goes into production. Being able to clean the mold cooling channels is also essential. Blind holes should be eliminated. When possible, the mold cooling channels should be drilled to the end of the mold plates and plugged. One way of cleaning them is with copper gauze or a wire brush on a wooden mandrel attached to a hand drill. Another method is using circulating mold cooling channel cleaning equipment with a cleaning agent designed to remove degraded circulating ﬂuid. High temperature heat transfer ﬂuids and ethylene glycol that contain deposits of carbon, scale, sludge, and most foreign matter found in cooling and heating circuits can be removed during maintenance of the mold. Air gaps between the assembled components of the mold cause a tremendous loss in heat transfer. No matter how well a cavity or core insert is ﬁtted into a mold cavity plate, there will be an air gap. Connecting the cooling channels into the cavities and core inserts wherever possible is essential. The use of insulation sheets on the external surfaces of the top and bottom clamping plates of the mold improve the mold cooling efﬁciency by preventing heat dissipation through the injection molding machine platens. 10.13.5.7 Injection Mold Temperature Controllers Closed loop injection mold temperature controllers are used to control the velocity of the heat exchange ﬂuid. These closed circulation systems are available in various models, depending on the application (temperature, capacity, number of cooling units, type of cooling ﬂuid, etc.). Heat is supplied when required by electric heaters. Cooling is supplied by the heat exchange of refrigerated or tower water through copper coils in the controller. This system has two advantages. It provides fast coolant circulating rates and a high Reynolds number (3,500 or higher), equivalent to a turbulent ﬂow. It also

628

10 Injection Mold Design permits easy treatment of distilled water to prevent scale from forming in the cooling channels of the mold, keeping the cooling channels and the cooling system clean. However, since because cooling tanks are made of stainless steel they should have manifolds to permit different ﬂow rates to different parts of the mold. Each manifold should have its own pressure gauge and thermometer, eliminating a major cause for loss of efﬁciency in the cooling system. Connecting water lines is time consuming. It may account for as much as half of the mold set-up time. This often leads to shortcuts by the set-up crew who do not always understand the necessity for the best possible mold cooling. Determining the best mold temperature conditions means switching water lines and seeing the results. To ease set-up and experimenting with the mold, quick disconnect ﬁttings and high temperature/pressure hoses are used to connect the mold to the cooling system. The male quick disconnect plugs are installed on the mold either protruding or countersunk and can be left in the mold. The quick disconnect female or socket has a ball check in it so that when it is disconnected it does not leak. There should be no restriction in the ﬂow because the inside diameter of the hoses or of the quick disconnect ﬁttings are smaller than the inside diameter of the system. It is preferable that the hose diameters be slightly larger. Manifolds at the source of water for intake and return should be available. The larger the temperature differences between the thermoplastic melt to be cooled and its cooling ﬂuid, the quicker the rate of heat removal. 10.13.5.8 Mold Cooling Fluid The most important properties of the cooling fluid are the heat transfer coefﬁcient and the ﬂow viscosity. Salt water is one of the best cooling ﬂuids, but causes steel corrosion inside the cooling channels. A common and practical closed loop cooling ﬂuid is a blend of distilled water with undiluted ethylene glycol additive with corrosion inhibitors, speciﬁcally formulated for use in injection mold cooling. This special additive has a synergistic package of industrial inhibitors that protect metal components by forming an extremely thin, bonded chemical coating. The inhibitors also keep the PH slightly on the base (non-acid) side, offering further corrosion protection. A blend of 25% additive and 75% distilled water for mold cavity temperatures between 28 °F and 197 °F and 50% additive and 50% distilled water for mold cavity temperatures between 0 °F and 204 °F is recommended. Do not use more than 65% of the circulating additive, because it reduces the heat transfer properties, making the cooling ﬂuid more viscous, thus reducing the ﬂow velocity. This circulating ﬂuid additive is a better additive than pure ethylene glycol, which is as corrosive as water and almost neutral in PH. When exposed to air in a circulator tank, it quickly becomes acidic and even more corrosive. It is also better than an automotive antifreeze additive, which also becomes acidic very quickly. While based on ethylene glycol, antifreeze has silica-based inhibitors that depend on agitation to scrub away corrosion. After two years, the inhibitors form a silica gel that prevents heat transfer and the entire cooling ﬂuid with antifreeze must be replaced.

629

10.13 Injection Mold Cooling External hose connection

Cores

Male pipe quick disconnect

Fluid flow

“IN" fluid ng Cooli flow lu F id " “OUT luid f g n Cooli

Mold core plate

Figure 10-49 Mold core plate parallel cooling with hose connections

10.13.5.9 Mold Plate Cooling Systems The temperature of a mold plate is controlled by circulating cooling ﬂuid through cooling channels drilled in the mold plate. The cooling channels are normally interconnected to form a circuit. The cooling circuit may be formed by one or more levels, depending on the depth of the mold cavity plate. The stationary cavity inserts and the movable core plates should always be cooled separately, using two separate mold cooling controllers to adjust each cooling ﬂuid ﬂow rate and temperature. Parallel Flow Mold Cooling The method of drilling parallel cooling channels across the mold plate is not recommended. The cooling ﬂuid temperature increases proportionally to the traveling distance through the mold cooling channels. This leads to temperature variations between the mold cavities, affecting the mold shrinkage of the molded parts, increasing the risk of signiﬁcant dimensional differences between the cavities. To improve the mold cooling efﬁciency, a uniform turbulent ﬂow around the cavities and the gate areas is recommended, where spiral or curved cooling channels may be located close to the gate. Mold Plate Parallel Circuit Cooling Parallel circuit cooling using external hose connections is an economical method for cooling mold plates. Figure 10-49 shows this cooling method. The temperature across the mold plate is nonuniform, but the cooling ﬂuid entrance area is the coldest section of the plate. The cooling ﬂuid removes heat from the hot cores, increasing its temperature as it travels through the parallel cooling channels. The cores placed close to the cooling channel entrance have lower temperatures than the cores located at the cooling circuit discharge end. 10.13.5.10 Mold Plate Cooling Channel Connections The simplest approach for cooling a mold plate having a small, shallow cavity is to drill two cooling channels, one on either side parallel to the cavity, connected at one end by means of a mechanical ﬁxture, using an internal cross drilling or a milled slot in conjunction with a connecting mold plate. Figure 10-50 shows the two design concepts for connecting the parallel cooling channels, using a mechanical ﬁxture with an internal channel milled in the base wall (mold plate), providing a continuous ﬂow path for the coolant. These mechanical connecting ﬁxtures are mounted onto the side of the mold plate and

Mold plate Screw Inlet Fixture “O” ring

Cooling channels Outlet

Surface mount cooling channel connecting fixture Mold plate Screw Inlet Fixture “O” ring

Cooling channels Outlet

Sunk mount cooling channel connecting fixture Figure 10-50 Mold plate cooling channels connecting ﬁxture designs

630

10 Injection Mold Design secured by screws. A gasket is incorporated to prevent leakage of the cooling ﬂuid. The top illustration shows the mechanical connecting ﬁxture externally mounted onto the mold plate. This design is cheaper, but it suffers from the disadvantage that the connecting ﬁxture may be disturbed during the mold setting operation and cooling ﬂuid leakage may occur. The bottom illustration shows the mechanical connecting ﬁxture mounted in a pocket, ﬂush with the side surface of the mold plate. This alternative design is recommended for connecting the parallel cooling channels without interferences. 10.13.5.11 Mold Cavity Plate Cooling Systems Figure 10-51 shows a simple mold cavity layout using three methods for cooling the mold cavity plate. The top illustration shows the mold cavity plate connected with three external hoses, starting with a top level going to the bottom and coming back to the top again. This cooling method is not recommended, because the cavity temperature is nonuniform, causing uneven cooling. The lower left illustration shows the mold cavity plate being cooled by the cooling ﬂuid. The drilled channels are interconnected and plugged, going in and around three sides of the mold cavity plate. This method will also give uneven cooling as the left hand side of the cavity will be warmer because there is no cooling channel to remove the heat. The lower right illustration shows the proper mold cavity plate cooling method. The cooling ﬂuid goes in, ﬂows around all sides of the cavity and out. Pipe plugs that can be removed are used for easy cleaning of the cooling channels. Figure 10-52 shows a mold cavity plate with four cavities and a single cooling circuit formed by drilling cooling channels, using plugs to seal the end of the channels and placing a diverting plug to control the cooling ﬂow direction.

Cavity plate Outlet Hose

Cavity

Inlet Diverting plug Inlet Outlet

Cavity (4)

Cavity (1)

Cavity (3)

Cavity (2)

Hoses

Low level

Cavity plate

Pipe plugs

Pipe plug

Cavity

Top level

Diverting plug Outlet Inlet

Outlet Cavity Inlet

Figure 10-52 Four-cavity mold plate cooling channel pattern design

Cooling channel

Figure 10-51 Mold cavity plate cooling channel connection systems

Cavity Pipe plug

631

10.13 Injection Mold Cooling This mold cooling system does not provide an even temperature control for the cavities. The sides of the cavities that are next to each other are hotter than the other sides, because the heat cannot be removed from these areas. Additionally, the average temperature in each cavity is different. Cavity (4) is the coldest, while cavity (1) has the highest temperature, because of the effects of a single cooling circuit and the mold layout. This mold plate cooling method is not recommended for critical molding applications.

10.13.6 Mold Cavity Insert Cooling Mold cavity insert cooling consists of spiral grooves cut to cool the outside surface of the cavity insert. Figure 10-53 shows a cavity insert cross section assembly and the grooves cut into the cooling plate. This cooling method provides the best performance and uniform cooling temperature of the cavity surface. This method is used when drilling inside the cavity is difﬁcult or when added cooling is needed. The cooling ﬂuid circulates around the cooling plate spiral groove, starting from the middle and exits at the end of the spiral groove or cavity outside diameter. The area exposed to the cooling ﬂuid (formed by the grooves) is in direct contact with the cavity base wall to control the uniformity of the cavity surface temperature. It is best to use an “O” ring rather than other types of gaskets, because the “O” ring allows a metal-to-metal seal. This makes holding the cavity insert in place easier.

“O" Ring "O" ring Cooling spiral groove

Circular groove

Cooling base plate

Cooling channel

Inlet

Cavity

“O" Ring

Outlet Preferred "O" rings mounting

Cavity insert housing

“O" Ring

et

Inl

Figure 10-53 Round cavity insert and spiral groove cooling base plate

Circular groove

Cooling channel

Inlet

10.13.6.1 Circular Cavity Insert Cooling Figure 10-54 shows cavity insert cooling by circulating cooling ﬂuid around the insert. The cooling ﬂuid is prevented from leaking by using “O” rings. The bottom illustration shows an incorrect method of applying the “O” rings. When the cavity insert is mounted into the mold cavity plate, the chances of tearing or damaging the “O” rings are high. This installation problem does not show until the mold is in the machine. The top illustration shows the preferred method of placing the “O” rings in compression; this method makes it very difﬁcult to damage the “O” rings.

“O" Ring

Poor "O" rings mounting design

Figure 10-54 Circular cavity insert cooling, “O” ring placement

632

10 Injection Mold Design 10.13.6.2 Round Deep Cavity Insert Cooling

Schematic pattern cooling flow

Figure 10-55 shows a large round and deep cavity insert cooling system using cooling channel ﬂow patterns along the path of the deep cavity. This method offers the maximum amount of cooling and a uniform temperature for the cavity. 10.13.6.3 Rectangular Deep Cavity Insert Cooling Figure 10-56 shows a large rectangular and deep cavity insert cooling system using a ﬂow pattern with three levels of cooling channels along the path of the deep cavity. This cooling method also provides the maximum amount of cooling with a uniform temperature control for the cavity.

t

Ou

In

10.13.6.4 Core Insert Cooling Systems

Pipe plugs Down

Up

Cavity

Up

Down

Diverting plug Down

Up

Outlet

Inlet

Figure 10-55 Round, deep cavity insert cooling channel system

During the thermoplastic injection molding process, when the hot melt is injected into the mold cavity, it shrinks onto the cores, forming a vacuum between the melt and core surfaces, providing for direct heat transfer. The elimination of air between these joints causes the melt to transfer heat at an accelerated rate, controlling the temperature of the core. Meanwhile, the melt shrinks away from the cavity surfaces, forming an air gap or thermal insulation, in which the heat is conducted poorly, reducing heat transfer to the cavity. Accordingly, the greatest amount of heat is transmitted to the core. To offset this, the core temperature should be set lower than the cavity to compensate for the heat transfer differences between these components. For shallow depth cores (less than 1.00 in), parallel cooling channels drilled into the core plate (single level) can be used; the cooling channels being situated beneath the core as shown in Figure 10-49. For deeper cores, however, the single level circuit is not sufﬁcient to permit the coolant to transfer heat away from the core surface fast enough. Special arrangement must be made to permit the circulation of coolant inside the core insert. If not enough heat is transmitted to the body of the mold and if the core is not cooled adequately by an appropriately designed channel, the temperature in the core increases as the molding production continues, causing a thermal overload. The consequences are increase in cycle time until production is stopped, and for cores with a square or rectangular cross section, the danger of side wall distortion.

t Inle

Inlet

let Out

Down level from 2 to 3

Outlet

Down level from 1 to 2

Inlet Level 1

Cavity Level 2

Level 3

Outlet

Schematic pattern cooling flow Figure 10-56 Deep cavity insert with three levels of channel cooling

Top view

Side view

633

10.13 Injection Mold Cooling Efﬁcient core insert cooling systems normally involve passing a coolant ﬂuid directly through channels or holes incorporated within the body of the cavity inserts. The cooling design is usually dependent on the size and shape of the insert. Only by circulating the coolant ﬂuid deeply inside the core insert can efﬁcient transfer of heat from the core surface be achieved. There are many alternative arrangements for cooling deep core inserts and most of these designs will be covered in the following section. Cooling System for Small Cores (0.125 in Diameter) Cooling of long cores with small diameters or width that are used to form blind holes, through holes, grooves, and slots is sometimes very problematical and requires appropriate design measurements of the mold. Air may be used to cool cores with diameters or widths of 0.125 in and above. The air cools the core either from the outside during the ejection time, e.g., through appropriately placed openings in a stripper ring, or ﬂows through a channel in the core. The disadvantage is the poor cooling efﬁciency. Cores provided with a central channel are better because of the efﬁcient cooling they provide. The air may ﬂow during the entire time the mold is open (time for opening and closing the machine and for ejecting the part). If, however, this time is not adequate to remove the heat, air cooling may be used during the entire cycle time with an appropriate mold design (discharge of the air through holes in the stationary half of the mold).

Air jet supply "INLET" Cores smaller than 0.125 inches diameter

This cooling system requires through holes or slots in the part to be molded, depending on the air pressure used (high pressure tends to produce whistling). Signiﬁcant temperature ﬂuctuations of these core cooling systems occur at the core during a molding cycle. The ideal situation for precision injection molding with largely constant speciﬁed mold temperatures cannot be realized here. It should, however, be pointed out very clearly that these cooling systems do provide opportunities for removing the heat from thin cores reliably. Figure 10-57 shows an air jet mold core cooling system. Cooling Systems for Cores Greater than 0.250 in Diameter In this case, the good thermal conductivity of copper or its alloys are used. A copper pin is press-ﬁtted into a blind hole centrally positioned in the core. This copper pin protrudes into a cooling channel and allows the conduction of heat from the core to the cooling medium. An increase in diameter of the copper pin ﬁt inside the core improves the heat conduction and heat transfer to the cooling ﬂuid because of the copper pin’s greater surface area. Figure 10-58 shows a solid copper rod mold core cooling system. Figure 10-59 shows a 0.250 in diameter mold core bubbler cooling system. The core is channeled down the center (blind hole) and a straight bubbler with an outside diameter smaller than that of the channel is introduced. The temperature control medium (cooling ﬂuid) is fed to the tip of the core through the tube and controls the core temperature during the return ﬂow. In single cavity molds, this is likely to be the hottest part of the core insert as it is directly opposite the sprue. While being the cheapest of the deep cooling methods to incorporate, the deep bubbler cooling design suffers from two major disadvantages:

Molded part

Core insert

Cooling air discharge

Figure 10-57 Air jet pressure core insert cooling system

634

10 Injection Mold Design

Figure 10-58 Solid copper rod core insert cooling system

Figure 10-59 Thermal pin core insert cooling system

• The ﬂow rate of the cooling ﬂuid drops markedly as it enters the core insert. This means that the required turbulent ﬂow is not achieved in this region and that the transfer of heat to the coolant is less effective. • It is possible for an air pocket to be formed at the top of the bubbler. An uneven temperature proﬁle, with associated molding problems, will result. It is essential that this port is always situated at the highest point of the chamber when the mold is mounted on the injection machine. It is important that molds incorporating a bubbler cooling design be engraved with information about which way the mold should be mounted on the machine. The design can be successfully used for molds for which the production rates are not important and where mold costs must be kept to a minimum. A more recent development is the use of thermal pins (heat pipes) instead of copper pins. Heat transport in these heat pipes is based on the phase changes (liquid/vapor and vice versa) of an encapsulated medium. Figure 10-59 shows a thermal pin mold core cooling system. Cascade Water Junction Cooling System Cascade water junctions are mold temperature control systems commercially available on the market. They are compact in design, can be installed and removed in “one step”, and they have 360° seals for accurate predetermination of port locations.

635

10.13 Injection Mold Cooling For cascade type mold cooling applications, the cascade water junctions provide mold cooling design ﬂexibility and ease of operation. Their compact design makes them ideal for cooling inserted cores or spot cooling in hard to reach areas of the molds. They can be rotated a full 360° without affecting their positive seal and are easily connected and disconnected even when installed internally. Final location of the ports on the body of the water junction can be accurately predetermined, thus assuring proper lateral alignment with pipe clearance holes. Water lines may be connected to the same side or opposing sides of the water junction. A slot on the water junction shows a body end port position and can be turned with a screwdriver to align the ports with pipe clearance holes.

Out In

The brass tube has the rigidity to maintain uniform spacing inside the water channel and is threaded into the body for ﬁrm support. Cascade water junctions are ideal for cooling molds where cooling channels through the mold plate are not possible due to interference. The nipple type water junction provides low cost rigid installation. The 2.00 in long pipe nipple can be replaced with a longer pipe nipple to suit the application.

IInn

OO uut t

The socket type water junction is more easily connected and disconnected when the mold is assembled. The socket is equipped with internal Viton® seals. Figure 10-60 shows two DME cascade water junction cooling systems.

Figure 10-60 Mold cascade water junction cooling system

Cooling Systems for Cores Greater than 0.375 in Diameter In cores with a diameter of 0.375 in and above, straight or spiral bafﬂer cooling systems are used. Figure 10-61 shows a straight brass bafﬂer core cooling system. Figure 10-62 shows a spiral brass plug bafﬂer core cooling system.

Figure 10-61 Straight brass plug bafﬂer core insert cooling system

Figure 10-62 Spiral brass plug bafﬂer core insert cooling system

636

10 Injection Mold Design The core is channeled down the center (blind hole) and a straight or spiral brass plug bafﬂer with an outside diameter smaller than the channel is inserted. The coolant is fed to the bottom of the core, and the channel is split by the plug bafﬂe, entering the front half and controlling the core temperature during the return ﬂow. The result is a nonuniform temperature distribution in the core. The spiral brass plug bafﬂer system is more efﬁcient than a straight brass plug bafﬂer system. Cooling System for Cores Greater than 0.750 in Diameter The core inserts greater than 0.750 in diameter permit the use of an outer core and an inner core with a drilled channel at the center interconnected at the top with a machined internal spiral groove. The inlet and outlet channels are drilled into the inner core, as shown by the dotted lines in Figure 10-63. The spiral creates turbulence in the ﬂowing temperature control medium and thus improves its efﬁciency. The temperature control medium ﬂows to the tip of the core in one half of the channel and returns through the internal spiral groove that is connected to the outlet channel. The core insert is ﬁtted into a solid support and “O” rings are used to prevent leakage. This is a very efﬁcient design that ensures that the cooling ﬂuid follows a precise path and no “dead cooling spots” are possible. The inner core forms a positive support to the insert outer core or shell; therefore, the wall thickness of this shell can be small. A more rapid transfer of heat from the melt is achieved as the cooling ﬂuid passes relatively close to the mold cavity surface. The advantage of this design is that it does not weaken the core to any major extent. Such cooling systems may be justiﬁed for large production of high quality and precision injection molding applications. Figure 10-63 shows an internal mold core spiral cooling channel system.

Figure 10-63 Core insert internal spiral cooling groove system

Core Cooling System for Large Double Walled Parts With cores greater than 1.50 in diameter for an injection molded end cup with double circular walls, a single core cooling system cannot satisfy the requirements of the application. To achieve the desired requirements, two separate core cooling systems must be employed. This design basically consists of a bubbler circuit

637

10.13 Injection Mold Cooling 0.375 inches minimum

Spiral cooling groove

“Inlet (1)”

“Outlet (2)” Cooling bubbler

“Inlet (2)” “Outlet (2)” Inner core “Outlet (1)”

Round molded part with two walls

Outer core

Figure 10-64 Double cooling core insert for round two walled parts

placed inside the inner core and a spiral circuit machined between the inner and outer cores. This type of core cooling satisﬁes the requirement for parallel ﬂow cooling. Unfortunately, in practice this double cooling system is both difﬁcult and expensive to produce. Figure 10-64 shows a core insert with double internal cooling systems used for injection molding a rounded end cup with two parallel walls. Core Valve Ejector Cooling System Bubbler cooling systems are used to remove heat from a valve ejector mounted as an integral component of the core. The head of the valve ejector is found on the top surface of the core, opposite to the gate. The valve ejector normally forms a relatively large part of the core insert top surface. Facilitating the dissipation of heat from this component is desirable. The stem of the valve ejector is bored to hold a bubbler system unit. The connectors are coupled to the supply and return lines via ﬂexible hoses to allow for the ejector valve movement. The cooling ﬂuid passes, via the inlet, down the center of the pipe, and back to the outlet via the outside of the pipe. This is the simplest method of cooling the valve type ejector, particularly if a commercial water junction unit is used. If passing the coolant through the head of the valve ejector is necessary, a more complex circuit is needed. The support feature is provided by a central column that can be integrated with the inner core, as a separate cooling circuit. Obviously, this is needed if the depth of the bubbler requires a column that is relatively long. The design has two primary objectives: • To support the central region of the core against possible deﬂection. • To have a solid central region of the core insert to permit the valve ejector to be incorporated. An additional “O” ring must be incorporated to avoid cooling ﬂuid leakage through the core assembly.

638

10 Injection Mold Design

Cavity three level cooling Valve ejector bubbler cooling Valve ejector Ejector plate

Molded part

Figure 10-65 Mold cavity, core insert, and valve ejector cooling

The disadvantage of the bubbler cooling system is the ﬂow rate drop as it enters the annulus with possible air pockets being formed. To overcome this cooling problem, it is recommended to incorporate a large diameter ejector valve and bubbler system, having its own coolant circuit in the mold design. This will ensure that an efﬁcient coolant circulation system is used at the critical area where it is required; that is, at the hottest location of the core insert, namely the ejector valve head surface where the gate is located. Figure 10-65 shows a cross section view of a valve ejector bubbler cooling system, core insert, and cavity cooling. Mold Cooling Summary These basic principles lead to very simple procedures for designing efﬁcient mold cooling systems. • Place the cooling channels at a proper distance from the mold cavity surface • The minimum cooling channel center distance to be three to four diameters in length • Do not use small diameters for cooling channels, specify 0.4375 in diameter minimum • Cool directly in the cavity and core inserts • Cool auxiliary mold plates if required • Drill through cooling channels and use pipe plugs • Be sure that the Reynolds ﬂow will occur throughout the mold. The Reynolds number should not be less than 3,500 • Treat the cooling channels with electroless nickel or equivalent. If not possible, clean them regularly • The cooling ﬂuid should be a blend of distilled water and special corrosion resistant additive (25% minimum to 65% maximum) so that no scale deposits are found in the cooling channels

639

10.14 Injection Molding Machine Nozzle • Use a portable digital thermometer with a surface and needle adapter to measure the cavity surface and process melt temperature during molding without interruptions to the process • Hook up the cooling ﬂuid on the mold so that it must ﬁll all the cooling channels before it over-ﬂows. Adjust the cooling ﬂuid temperature and ﬂow rate on the mold cooling system • Vary the cooling ﬂow rate, the mold temperature, parallel or series hookups, and direct or counter ﬂow • Take samples under each condition, test them and make the best compromise between economics and quality • Record this information for future runs.

10.14

Injection Molding Machine Nozzle

The nozzle that screws into the barrel end cap is a tubular shaped component of various lengths and inside diameters, long enough to reach the mold. It is intended to bridge from the injection unit to the mold, providing passage for the melt through the stationary platen to the mold’s sprue. It commonly consists of a one-piece unit or removable separate nozzle tip, which screws into the main body of the nozzle. The main nozzle body extends from the barrel end cap to the sprue, using an independent heater band and a thermocouple at the rear end. The nozzle tip provides the ﬁnal transition from the larger diameter of the nozzle body to the sprue bushing inlet. With a cold sprue, the outlet oriﬁce of the nozzle tip is slightly smaller than the inlet oriﬁce of the sprue. This ensures that the frozen sprue can be ejected. A removable nozzle tip is desirable for changing molds with different sprues and molding different plastic materials. The oriﬁce of the nozzle tip can also be designed for various materials and processes. However, these standard nozzles are not designed to be temperature-control devices. Figure 10-66 shows three nozzle designs. The Kona nozzle uses four heat pipes mounted around the entire streamlined ﬂow channel of the nozzle. The heat is drawn from the barrel and its own heater band, using a thermocouple for uniform heat distribution of the nozzle. The one piece nozzle design has a front-mounted thermocouple and heater band. The Mold Master injection nozzle has a front mounted thermocouple and strategically distributed heaters that have been cast around the nozzle body providing precise and uniform internal temperature control through the entire ﬂow channel, eliminating internal hold up spots.

10.14.1 Mold Cold Runner System The cold runner system transfers the thermoplastic melt from the injection nozzle to the cavities. It consists of sprue, sprue puller, runner, cold slug pockets, and gates. All these components will be discussed in detail in this chapter. 10.14.1.1 Cold Runner Sprue The sprue transfers polymer melt and heat from the molding machine nozzle to either a runner system or directly to a cavity. A cold sprue is formed inside a sprue bushing, which must withstand repeated impacts from the injection nozzle each time it engages the sprue bushing. The sprue bushing must have good wear and corrosion resistance, a good heat transfer coefﬁcient, surface ﬁnish, and heat resistance. Most sprue bushings are made from SAE 6145 hardened tool steel.

Plug

Heat Pipe

Heater band

Removable nylon tip Thermocouple

Heater junction

Kona nozzle with heat pipes, heater and thermocouple Front thermocouple

Standard Heater band nozzle tip One piece nozzle design with a special front mounted thermocouple and front heater band Front thermocouple

Cast-in heater coil Removable nylon nozzle tip

Heater junction

Mold master injection nozzle with a fused heater coil strategically distributed with front thermocouple

Figure 10-66 Injection nozzle designs with better temperature controls

640

10 Injection Mold Design Moving half

PL

Locating ring Full round runner Mold fixed half

Front thermocouple

Heater band

End cap Barrel

Nylon nozzle Free flow check valve Sprue puller

Screw

Sprue bushing Contact radius surface

Heater band Magnified view

Figure 10-67 Contact radius surfaces between injection nozzle tip and sprue bushing

Magnified view

Nozzle

Sprue Sprue bushing

Flashing

Poor contact of radius surfaces does not provide a good seal flash limits sprue ejection Magnified view

Nozzle

The contact radius surface between the injection nozzle and the sprue bushing helps with alignment of the ﬂow channels between these components. If there were any misalignment, an undercut would be created and inhibit the sprue from being pulled from the bushing. The radius of the sprue bushing should be slightly larger than the injection nozzle tip to ensure sufﬁcient sealing without ﬂashing. However, if either contact radius surface of these components is in poor condition, ﬂashing will retain the sprue and manual removal will be necessary. The sprue bushing “O” diameter should be a minimum of 0.031 in larger than the injection nozzle tip oriﬁce diameter to avoid molding cycle interruptions to free the sprue from the bushing. When the mold opens it is essential that the sprue breaks away from the injection nozzle and is pulled completely from the sprue bushing. With single-cavity molds, the sprue feeds directly into the base of the molded part and the sprue is pulled while the molded part is ejected from the cavity. Figure 10-67 shows the interface between the contact radius surfaces of the sprue bushing and one piece injection nozzle design.

Sprue Sprue bushing

Flashing

Nozzle contact orifice larger than sprue “O” dia. flash limits sprue ejection

400% Melt flow area runner restrictions

90˚ Rotation sprue bushing

Sprue bushing rotates and runner flow area shears the melt limiting the flow rate

Figure 10-68 Interface problems between injection nozzle and standard sprue bushing

Figure 10-68 shows three common problems between these components. In the top illustration a lack of interface contact radius surfaces between the sprue bushing and the injection nozzle tip causes ﬂashing problems. The middle illustration shows a larger diameter for the nozzle contact oriﬁce than the sprue “O” diameter, which causes sprue undercuts and ﬂashing that also contribute to ejection problems of the sprue. The main runner is machined in the base of the sprue bushing, linking the sprue to the runner system. There is rarely any positive anti-rotation provided on the sprue bushing to ensure runner alignment. This commonly results in these two sections of the runner becoming misaligned, causing melt ﬂow restriction during the production run, as shown in the bottom illustration. A retaining dowel pin located in the shoulder of the sprue bushing ensures runner alignment. The interface between the sprue bushing and the nozzle must not be deformed as ﬂashing or ejection problems might occur. The temperature controls or thermal interfaces between the injection nozzle and sprue bushing contact surface radius is one of the major problem areas in the injection molding process. The machine’s injection unit plasticizes and delivers the polymer melt at the correct process viscosity, temperature, and injection pressure to the injection nozzle to ﬁll the mold. Standard sprue bushings and nozzles have difﬁculties maintaining a uniform melt viscosity, if only one heater band in the middle and a thermocouple at rear of the nozzle are used.

641

10.14 Injection Molding Machine Nozzle The temperature of the injection nozzle tip is the most critical and difﬁcult to control. If either the nozzle or the sprue bushing is not thermally insulated, heat is dissipated from the nozzle tip to the atmosphere when idle, or through the sprue bushing, while the sprue bushing becomes colder because the heat is absorbed by the mold base and machine ﬁxed platen. Injection nozzles with removable tips (different heat transfer coefﬁcient) have oriﬁces of different diameters, designs, and volumes. Each section requires various heat levels, which one heat source cannot deliver. The lack of positive contact surface seals, alignment and temperature control between these components is an injection molding process problem. The injection nozzle provides a force to seal with the sprue bushing. This interaction causes heat control problems and cooling of the melt as it enters the mold. Following are some guidelines when designing or specifying a sprue: • The sprue must not freeze before the runner system and cavities. This is necessary to complete the molding process screw forward time. • The sprue must be ejected easily and reliably. • The sprue interfaces with the injection machine nozzle must not have any ﬂash to avoid sprue ejection problems. • At the base of the main runner, in line with the sprue, a sprue puller pocket should be provided to act as a cold slug well. • Sprue dimensions should be calculated based on the sprue length and total shot size. • The outlet sprue diameter should be at least 0.062 in larger than the molded part wall thickness, depending on the sprue length, shot size, polymer crystallinity, and cold runner system total pressure drop. • The sprue bushing “O” diameter should be at least 0.031 in larger than the outlet oriﬁce of the injection nozzle. • The sprue bushing inside diameter must be tapered to 2.38° per side or 0.50 in/ft. The tapered surface reamed and hardened for the sprue to be pulled out of the oriﬁce when the mold is opened. • Sprue bushing radii should be modiﬁed, slightly increasing the outside edges to reduce the contact radius with the nozzle face for a better alignment and seal surface. • The sprue bushing should have a dowel pin positioned to prevent rotation and misalignment with the main runner. • Reduce the sprue length to a minimum to control the molding cycle.

0.125 r.

“O”

Standard Sprue Bushings Standard sprue bushings are most commonly used in the construction of thermoplastic injection molds. They are commercially available in various designs and sizes. These standard sprue bushings are used for simple 2- and 3-plate mold designs and for injection molding commodity thermoplastics of low or medium crystallinity rates, good thermal stability, and low melt process temperatures. Figure 10-69 shows a typical standard commercially available sprue bushing.

R. 0.50 Taper/foot (included angle) 0.187 “O” = 0.156, 0.218, 0.281 “R.” = 0.50, 0.75 Figure 10-69 Standard commercially available sprue bushing

642

10 Injection Mold Design “O” Dia. = 0.156; 0.218; 0.281 (inches) “R” = 0.50; 0.75 (inches) 420 Stainless steel hardened Ampcoloy 940 (beryllium free)

0.50 (inches) taper per foot (included angle) 0.187 (inches)

Performance Alloy Sprue Bushings Figure 10-70 shows a new type of sprue bushing commercially available in designs and dimensions found in standard sprue bushings, but with better temperature control and performance required for more critical molding applications. These sprue bushings are made of Ampcoloy 940 (high thermal conductivity alloy) to improve heat transfer; however, this alloy is very soft. To improve the toughness required at a contact surface radius, an insert made of 420 stainless steel hardened is fused inside the alloy body. TranziSprueTM Engineering thermoplastics for high melt process temperatures with high melting point temperatures, fast crystallization rates, requiring high mold temperatures, and using fast molding cycles will cause the sprue bushings to show poor performance. A new type of sprue bushing with the same standard dimensions, reasonably priced, has shown excellent performance when molding engineering polymers. Figure 10-71 shows the TranziSprueTM design. It uses a heater band and thermocouple to control the contact surface radius temperature and insulates the sprue from the mold base plate by using thermal barriers. Extension Nozzle Sprue Bushings

Figure 10-70 Performance alloy sprue bushing to improve temperature control

Extension nozzle sprue bushings are used to reduce the length of the sprue (economical reasons), improve the cold runner system melt injection efﬁciency, reduce the cold runner system pressure drop, and lower the injection pressure required for the molding process.

“O” Dia. = 0.156; 0.218; 0.281 (inches) Heater band Thermal barrier

Variations of the extension nozzle sprue bushings are used in horizontal 2-plate molds feeding directly to a single cavity or into a runner feeding several cavities in 3-plate molds and in vertical inserts encapsulation molds. Figure 10-72, top shows a commercially available extension nozzle sprue bushing, the bottom illustration shows an extension nozzle sprue bushing mounted in a 3-plate mold, reducing the length of the sprue.

0.50 (inches) taper/foot (included angle)

“R” = 0.50; 0.75 (inches)

0.187 (inches) “R” = 0.50; 0.75 (inches)

0.156" Dia. + 5˚/Side

0.375 (inches)

Extension nozzle sprue bushing

Short sprue Extension nozzle

Figure 10-71 TranziSprue, a sprue bushing with temperature control

Magnified view Short sprue, 3 plate mold & extension nozzle

Figure 10-72 Extension nozzle sprue bushing and short sprue mold

643

10.14 Injection Molding Machine Nozzle Hot sprue bushing to reduce sprue length Hot sprue bushing reverse taper gate

Short cold sprue

Second parting line

Conical sub runner Center gated molded part Main parting line

Magnified view

Figure 10-73 Short sprue in a 3-plate mold, using a hot sprue bushing

Hot Sprue Bushing A hot sprue bushing is an alternative to an extension nozzle sprue bushing. The hot sprue bushing directs the melt from the injection molding machine nozzle directly into a single cavity or into a cold runner feeding either multiple cavities or multiple gates feeding a single cavity. In applications where a long sprue and large diameter control the molding cycle (prolonged cycle time), the hot sprue bushing provides an excellent alternative. The hot sprue bushing with a good gate design (reversed taper) has few of the drawbacks of a hot runner system. They are relatively inexpensive, easy to install, operate, and maintain. They require only one additional temperature control. Figure 10-73, shows a hot sprue bushing application used to reduce the length of a cold sprue in a 3-plate mold. How to Size the Sprue Knowing the shot weight passing through the sprue and the length of the sprue bushing, the sprue size is calculated using Eq. 10-22. DAvg. =

W 0.5 × L0.25 = 16

W × 16

4

L

(10-22)

Where: DAvg = Average diameter of taper sprue (in) = Average inlet/outlet sprue diameters to the nearest 0.06 in W = Total shot weight (oz) L = Sprue length (in)

Example 10-6 Determine the thermoplastic sprue diameters for a total shot weight of 4.0 oz and a sprue length of 4.00 in. DAvg. =

40.5 × 40.25 2 × 1.414 2.828 = = ≈ 0.187 in Dia. 16 16 16

Taper sprue diameters: Dinlet = 0.125 in, Doutlet = 0.25 in

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10 Injection Mold Design 10.14.1.2 Sprue Puller Nozzle tip

The cold slug well or sprue puller pocket is one of the most effective traps for the cold slugs or colder melt front. When the melt stops ﬂowing, the heat transfer is reduced, the tip of the nozzle becomes colder and a cold slug is formed inside the injection nozzle’s front oriﬁce. Its function is to stop the ﬂow of the melt front at the nozzle each time the mold opens and the cold sprue breaks from the nozzle tip. During mold opening, the cold sprue is pulled from the sprue bushing by an undercut in the sprue puller pocket. The sprue puller is an automatic ejection device that functions in coordination with the sprue puller ejector pin. When the ejector pin moves forward, it pushes the cold slug well, forcing the sprue out of the undercut allowing the sprue and runner to fall freely.

Sprue puller ejector pin

Figure 10-74 illustrates the sprue puller operational sequence.

Mold closed, nozzle injects melt inside sprue bushing, sprue puller pocket and runner

Cold slug

Main runner

Sprue puller Mold opens and breaks the sprue from the nozzle tip by action of sprue puller undercut

Sprue bushing

The size of the sprue puller is often the determining factor that controls the length of the molding cycle. A big sprue puller requires a longer cooling time to cool off and become rigid to eject the sprue from the sprue bushing without breaking. If the sprue puller’s large volume (mold cooling is not provided in this area) is not frozen sufﬁciently at the time of mold opening, it will break and the sprue will remain lodged in the sprue bushing. To avoid this, the molding process cycle must be longer to allow additional cooling time for the sprue puller. When designing cold runner molds, the ejection of the sprue and the sprue puller runner system should be analyzed to make the best selection to ensure that the runner system does not stick to the ﬁxed half of the mold, causing interruptions in the molding process. The sprue pullers are found directly on the opposite side of the sprue entry and on the lower side of the main runner. The thermoplastics melt from the nozzle ﬂows through the sprue bushing, the front cold slug ﬁlls the sprue puller pocket, and upon solidifying, provides sufﬁcient adhesion to pull the sprue away from the nozzle and the sprue bushing as the mold is opened. There are many techniques and designs of sprue pullers; however, only the recommended sprue puller designs that have shown excellent performance and reliability in the ﬁeld will be discussed here. Figure 10-75 shows three small and efﬁcient sprue puller designs recommended for controlling cooling problems and to reduce the injection molding process cycle time. Illustration (a) shows two sizes of reverse taper sprue pullers (based on mold size). It is the most common design used for molding both high and low melt temperature materials, unreinforced, impact modiﬁed, ﬁber glass, and mineral reinforced resins.

Cold sprue with runner system

Fully opened the sprue puller ejector pin pushes the runner

Figure 10-74 Ejection sequence of the cold sprue and runner system

Illustration (b) shows the “Z” type sprue puller. It is used in similar applications as the reverse taper. It is not recommended with brittle materials. The internal radius in the ejector pin should be smooth and polished. The ejector pin should not rotate to avoid hang-ups of the runner and the ejector pin. The ejector pin has a short operational life (breakage at the thin cross section). Illustration (c) shows two sizes of annular ring pullers (based on mold size); this design is recommended only for unreinforced resins. Designing the optimum runner layout with exact dimensions will maximize savings of both raw materials and energy consumption during the injection

645

10.14 Injection Molding Machine Nozzle molding process. At the same time, runner size is constrained by the amount of pressure drop and the injection capacity of the machine. This mold design section presents some simple procedures to follow in minimizing runner size, maximizing productivity, and molded part quality.

0.12 r.

“D” dia.

Cold runner systems consume material and energy. Since molding a runner system obviously costs money, it only makes sense to minimize the amount of unusable material molded into the cold runner system. Even though the cold runner will probably be reground and recycled, it is important to keep its weight and size to the absolute minimum because some thermoplastic polymers tend to degrade during repetitive processing.

C

0.06 r.

0.025

It is necessary to provide a ﬂow-way in the injection mold to transfer the melt from the injection machine nozzle to each mold cavity. This ﬂow-way is referred to as the cold runner system. Normally the runner comprises a sprue, runner, and gate. These terms apply equally to the molded parts that are removed from the runner and gates in the process of extracting the molded shot. A typical full shot for four cavities and runner is shown in Figure 10-76. The thermoplastic melt passes through the sprue, main runner, branch runners, and gate before entering the cavity. It is desirable to keep the traveling distance of the material to a minimum to reduce pressure and heat losses. Therefore, careful consideration must be given to the mold cavity layout. The purpose of the runner cold slug pockets is to catch the melt that has chilled at the front of the nozzle. The runner is a channel machined into the mold cavity plate surface (parting line) to connect the sprue with the entrance (gate) to the cavity. There are two key functions of a runner that always must be thought out carefully during the mold design stage. The ﬁrst obvious function of the cold runner is to convey melt into individual mold cavities. In this respect, the runner works like a ﬂuid transfer line and the principles of ﬂow govern the design of the runner. The second, often less obvious function of the cold runner system, is a mechanical one. For instance, in a three-plate mold, sucker pins must be designed so that the runner is retained on the front plate and is stripped off the sucker pins as the mold opens. Design for mechanical operation of the mold should also allow the runners to fall free if automatic operation is desired.

“A” dia.

A

B

C

D

0.250 0.276 0.250 0.250 0.375 0.414 0.375 0.375

Reverse taper

0.12 r. 0.062

Parting line 0.06 r.

A + 0.062

0.02 r.

A/3 45˚

A/3 “A” dia.

“Z” type sprue puller

0.12 r.

“A”

“C” r. 0.125

“A” dia.

“B” dia. Molded part Sprue puller

Main runner

Edge gate

A

B

C

0.250 0.258 0.004 0.375 0.386 0.006

Annular ring Figure 10-75 Recommended sprue puller designs Cold slug pocket

Sprue

Full shot molded on a 2 plate mold

Secondary runner

Description of components produced in a full shot

Figure 10-76 Illustration of molded components in a complete shot

646

10 Injection Mold Design If additional handling of a cold runner system is anticipated, such as breaking the gate from the molded parts, the runner diameter connected to the gate should be large enough to provide the rigidity to conform to the degating requirements. In such a case, sizing a runner to the absolute minimum ﬂow requirement, otherwise highly desirable, might be shortsighted from a material handling point of view. These mechanical functions of the cold runner system are important and can often determine the practicality of automatic operation or how well a mold functions. The surface walls of the runner channel must be smooth to prevent any restriction to ﬂow. Also, as the runner has to be removed with the molded part, there must be no machine marks left which would tend to retain the runner in the mold plate channels. The runner channels should be polished and harden in the ﬂow direction of the melt. Important considerations for designing the cold runner system: • Cold runner system layout • Cold runner cross section geometry • Cold runner dimensions. Cold Runner System Layout The layout of the runner system will depend on the following factors: • The number of cavities • The geometry of the molded parts • The type of mold (i.e., two-plate or three-plate mold) • The type, geometry, and size of the gates. Cold Runner Cross Section Geometries The cold runner cross section geometries used in mold designs are usually one of four conﬁgurations, as shown in Figure 10-77 (a full round, a parabolic, a trapezoidal, and a half round cold runner). While the half round runners were prevalent in the earlier days of injection molding, they are rarely encountered today. The half round runner design is the most inefﬁcient of the four designs in delivering a melt, because its ratio of pressure loss to runner unit volume is very high.

Full round

Trapezoidal

Parabolic

Frozen skin

5˚- 10˚ (Typ) Frozen skin

R

Half round

5˚ - 10˚ (Typ) Frozen skin

Effective melt flow area R

D D

D R

Effective melt flow area Best runner design

R

D

Frozen skin Effective melt Effective melt flow area flow area Good runner design Practical runner design Poor runner design

Figure 10-77 Typical cold runner cross section geometries

647

10.14 Injection Molding Machine Nozzle The full round, parabolic, and trapezoidal designs are the most efﬁcient. Each offers distinct advantages for speciﬁc situations. The parabolic and trapezoidal runners are machined only on one cavity plate, the side where the ejector system is located and where the runner must be retained as the mold opens. It is then stripped away during part ejection. The trapezoidal runner is commonly used with three-plate molds. The full round cold runner can be machined into both cavity plate surfaces, where the parting line is located and there are no mechanical obstacles. A full round runner offers the lowest pressure loss per unit volume of all designs. The criteria for efﬁcient cold runner design are that the runner should provide a maximum cross section area from the standpoint of pressure transfer and minimum contact on the periphery from the standpoint of heat transfer. As the thermoplastic melt progresses through the mold’s cold runner system, the melt adjacent to the cold mold surface rapidly decreases in temperature, forming an external solidiﬁed skin. The melt that follows will pass through the center of this solidiﬁed material and because of the low thermal conductivity that most thermoplastics possess, the solidiﬁed skin acts as insulation and keeps the temperature of the melt’s central core hot, allowing the internal melt to ﬂow. The cross section area of the runner must be sufﬁcient to permit the melt to pass through to ﬁll the cavity and for packing pressure to be applied for shrinkage control of the molded part before the runner freezes. The cross section area of the runner should not be able to control the injection cycle. The main objection to the full round runner is that this runner is formed from two semicircular channels machined separately in each of the mold plates. It is essential that these channels are accurately matched to prevent the development of an undesirable and inefﬁcient runner system. Accurate matching of the channels means that the mold cost for a mold containing round runners will be higher than for one containing trapezoidal runners.

Sub runner

Cavities

Sprue

Cold Runner Dimensions Considerations required for specifying the cold runner dimensions:

Main runner

• The wall thickness and volume of the molded part • The distance of the cavity from the main runner or sprue

8 Cavities unbalanced runner

• The mold cooling system for the runner and gates

Cavities

• Type of cold runner cross section design • The thermoplastic melt ﬂow rate • The thermoplastic viscosity and shear rate characteristics The runner length should always be kept to a minimum to reduce pressure losses and the cold runner system should be balanced. Runner balancing means that the distance, the volume, and heat transfer characteristics should be identical for each channel. When the thermoplastic melt travels from the sprue to each cavity, it ensures that all the cavities will ﬁll uniformly and without interruptions. Figure 10-78 shows the difference between balanced and unbalanced cold runner systems.

Main runner

Sprue

3th runner 2nd runner 8 Cavities balanced runner

Figure 10-78 Difference between balanced and unbalanced runners

648

10 Injection Mold Design Figure 10-79 shows three different balanced runner conﬁgurations that provide uniform injection pressure to all cavities. There have been a number of misconceptions about proper cold runner design. In the past, many injection molders and tool builders felt that the larger the runner, the faster the melt would be conveyed to the cavity. They also believed that the lowest possible pressure loss through the runner’s system to the cavity would be the most desirable. There are now techniques for computing the minimum runner size required to convey the melt at the proper rate and pressure loss to achieve optimum molding quality. 4 Cavities balanced runner

10.14.1.3 How to Design a Cold Runner System The ﬁrst step is to examine the mold layout and assign the cavity spacing. Always, the distance between cavities should be as short as possible. The second step is to ﬁgure out the gate locations, which depend on part design and performance criteria. In striving for a balanced cavity layout, most mold designers prefer 2, 4, 8, 16, 32, or 64 cavities. Sometimes, having balanced cavities is not possible, especially when the goal is to put as many cavities into the mold as possible to reduce molding costs. 8 Cavities balanced runner

If good general design principles have been used for determining runner length and cavity spacing, ﬁnding the most efﬁcient runner size using a relatively simple set of calculations is now possible. Computing Runner Diameters These computations are based on a key rheological property of the thermoplastic resin to be molded. This property is the material’s shear rate vs. its melt viscosity at several commonly encountered melt temperatures for that material. Usually, this information is available from the resin supplier and is frequently displayed in molding manuals for individual materials.

16 Cavities balanced runner

Figure 10-79 Three different balanced runner conﬁgurations

Since no single calculation will do the job, starting with a reasonable runner size estimate is necessary, based on a balanced melt distribution through the runner. In fact, we must select these diameters based on ﬂow and equal volume distribution to all cavities. The melt ﬂow must be proportionally distributed as shown in Table 10-7: Table 10-7 Balanced Cold Runner Melt Flow Distribution

Sprue

100% Flow

Main runner

50% Flow

2nd Runner

25% Flow

3th Runner

12.5% Flow

4th Runner

6.25% Flow

5th Runner

3.12% Flow

Table 10-8 was developed considering the balanced cold runner conﬁguration for 64 cavities as shown in Figure 10-80, equal ﬂow distribution, and the most popular cutter sizes used by the mold makers.

649

10.14 Injection Molding Machine Nozzle

Runner # 4

Runner # 1 Runner # 2 Runner # 5

Runner # 6

Runner # 3

Figure 10-80 64-cavity balanced runner (reference for Table 10-8)

Table 10-8 Equal Flow Runner Diameter (in)

2 Cavities

4 Cavities

8 Cavities

16 Cavities

32 Cavities

64 Cavities

Runner #1

Runner #2

Runner #3

Runner #4

Runner #5

Runner #6

0.375 Dia.

0.312 Dia.

0.281 Dia.

0.250 Dia.

0.187 Dia.

0.156 Dia.

0.312 Dia.

0.281 Dia.

0.218 Dia.

0.187 Dia.

0.156 Dia.

0.131 Dia.

0.281 Dia.

0.250 Dia.

0.218 Dia.

0.187 Dia.

0.156 Dia.

0.125 Dia.

0.250 Dia.

0.218 Dia.

0.187 Dia.

0.156 Dia.

0.125 Dia.

0.109 Dia.

0.218 Dia.

0.187 Dia.

0.156 Dia.

0.125 Dia.

0.125 Dia.

0.093 Dia.

0.187 Dia.

0.156 Dia.

0.131 Dia.

0.125 Dia.

0.093 Dia.

0.078 Dia.

0.156 Dia.

0.131 Dia.

0.125 Dia.

0.093 Dia.

0.078 Dia.

0.131 Dia.

0.125 Dia.

0.093 Dia.

0.125 Dia.

0.125 Dia.

0.093 Dia.

Solidified skin

D (Dia.) = T + 0.040

0.093 Dia.

T (Dia.) Parabolic runner next to gate

Gate Runner Larger than Part Wall Thickness The last runner connected to the gate and cavity should be 0.040 in larger than the wall thickness of the molded part. Figure 10-81 shows the cross section of the molded part, edge gate, and the last runner diameter in relation to the part wall thickness. With the calculated diameter from Table 10-8, select the column for the number of cavities, select the row closest to the diameter required for the application, and use the other recommended runner diameters for the preliminary pressure drop calculations. A minimum runner diameter of 0.078 in should be used to prevent premature runner freeze-off and high shear rate conditions.

L (Land) = 0.040 in. Rectangular edge gate

H (Depth) = (40 - 60%) T “T” Part wall thickness

Parabolic runner 45˚

Molded part

Figure 10-81 Runner diameter next to gate larger than part wall thickness

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10 Injection Mold Design

Example 10-7 3rd runner

2nd runner

Main runner

Calculate the optimum runner size of an eight-cavity mold as shown in Figure 10-82, the molded part is made of unreinforced nylon 6/6, with a speciﬁc gravity of 1.00, weighing 15.0 grams; an injection speed of 3.0 seconds is used.

Sprue

Figure 10-82 Eight-cavity balanced cold runner analysis

Using a tertiary runner diameter of 0.125 in to satisfy the molded part wall thickness, based on Table 10-8 for eight cavities and a tertiary runner diameter of 0.125 in, a secondary runner diameter of 0.131 in, and a main runner diameter of 0.156 in are needed. With the lengths of the runners as shown in the table below, we can now compute the total volume of the melt that must be conveyed through each runner segment. Since this is a balanced mold, we need to calculate only half of the mold. Table 10-9 Runner Dimension Analysis

Length (in)

Original diameter (in)

Modiﬁed diameter (in)

Main runner

5.00

0.156

0.140

2nd Runner

3.00

0.131

0.131

3rd Runner

1.00

0.125

0.125

Total runner weight (g)

7.37

6.78

Runner pressure drop (psi)

5,785

6,036

Viscosity, (lb-sec/inch2 --Poise)

Since the molded parts weigh 15.0 g each, we must convey through each half of the primary runner the weight of four parts plus the weight of the runner’s system itself. Here, we must convey 60.0 g (4 × 15) of material plus the volume of the runner’s system through the left side of the mold. In this example, we assume that we are using a material with a speciﬁc gravity of 1.0. If the volume of half the runner’s system is “X”, the primary runner must carry 60 + “X” g of material within 3.0 s, giving us a ﬂow rate (Q) of (60 + “X”) / 3.0 in3/s.

10.0 1.0 0.1 0.01

536˚F. 554˚F. 572˚F.

0.001

0.0001 0.01 0.1

1

10

Shear rate, (1/sec.)

100 1.000

10.000 100.000

Figure 10-83 Viscosity vs. shear rate for unreinforced nylon 6/6 (Courtesy: Du Pont)

Because the melt viscosity depends on shear rate and temperature, the only remaining data we need are the melt temperature and the viscosity vs. shear rate graph of unreinforced nylon 6/6. Using the graph shown in Figure 10-83, we select the melt viscosity of the material at the calculated shear rate. With the above information, we can now determine the shear stress on the resin in the runner’s system: shear stress equals melt viscosity times shear rate. Finally, we can compute the pressure drop through the primary runner; pressure loss being the key factor in determining an optimum runner diameter. This procedure can be used to calculate pressure drops occurring in the secondary and tertiary runners as well.

651

10.14 Injection Molding Machine Nozzle How to Calculate the Pressure Drop in the Runner Engineering a cold runner system requires an understanding of the pressure drop of the polymer melt as it passes through a channel. This pressure drop is controlled primarily by the volumetric ﬂow rate or injection speed, the melt viscosity, temperature, and the channel dimensions. While it is possible to reduce the melt viscosity by increasing the melt temperature, thus reducing the pressure drop, most injection molding resins have an ideal melt temperature that provides fast cycles and optimum part quality. A well engineered cold runner should be designed by assuming an ideal melt temperature. This temperature can be found in the resin supplier’s molding manual. The other assumption that must be made initially is how much pressure drop can be tolerated. If the injection molding machine can deliver 20,000 psi pressure, the runner system should be designed so that the pressure required is somewhat less than the machine’s capacity. A runner pressure drop below 10,000 psi should be used for difﬁcult parts to ﬁll (thin and long walled parts); a pressure of 5,000 psi is usually adequate to ﬁll and pack out most cavities. This means, that the runner’s system can be designed for a 10,000 psi pressure drop. Example 10-8 The sample calculations of an eight-cavity, balanced runner layout, with a main runner of 5.0 in length and a diameter of 0.156 in, a secondary runner with 3.0 in length and a diameter of 0.131 in, a tertiary runner with 1.0 in length and a diameter of 0.125 in. We assume that all runners are full round, a resin speciﬁc weight of 1.0, part weight of 15.0 g, and an injection time of 3.0 s. For eight cavities together, the total volume is 120 g or 7.32 in3. Volume (in 3 ) =

Weight (g) 120 = Specific Gravity × 16.38 1.0 × 16.38

(10-23)

= 7.32 in 3 Runner Volume = V = π r 2 L

(10-24)

Where: r = Runner radius (in) L = Runner length (in) Thus, Primary runner volume: Secondary runner volume: Tertiary runner volume:

V1 = 3.1416 × 0.0782 × 10 = 0.191 in3 V2 = 3.1416 × 0.06552 × 12 = 0.161 in3 V3 = 3.1416 × 0.06252 × 8 = 0.098 in3

Total volume (runner + parts) = 7.32 + 0.191 + 0.161 + 0.098 = 7.77 in3 Since the ﬂow splits at the intersection of the sprue and primary runner into two identical halves of the runner system, we need only calculate the pressure loss through one half of the mold. The volume of the melt, therefore, that must be conducted through the primary runner in this half of the system is 3.88 in3.

652

10 Injection Mold Design

The volumetric ﬂow rate = Q = The shear rate = Sr = γ=

Volume 3.88 = = 1.29 in3/s (10-24) Fill Time 3.00

4×Q π × r3

(10-25)

4 × 1.29 = 3,463 (1/s) 3.1416 × 0.0783

The melt viscosity at this shear rate and at the speciﬁed melt temperature must be read from the unreinforced nylon 6/6 graph (Figure 10-83). For this hypothetical example, the apparent melt viscosity, μ = 0.0070 lb·s/in2. The shear stress = τ = μ × γr

(10-26)

τ = 0.0070 × 3,463 = 24.248 psi The pressure drop = P =

P=

τ×2L r

(10-27)

24.24 × 2 × 5 = 3,107 psi 0.078

Now, the next runner segment must be considered. The total volumetric ﬂow through each secondary runner is 3.88 in3 minus the volume in the primary runner: (3.88 – 0.095) / 2 = 1.89 in3 (remember that the ﬂow splits in half again at the secondary runner). The volumetric ﬂow rates in each secondary segment = 1.89 / 3 = 0.63 in3/s. Thus, γ=

4 × 0.63 = 2,930 (1/s) 3.1416 × 0.0653

The melt viscosity at that shear rate is 0.0080 lb·s/in2. Therefore, τ = 0.0080 × 2930 = 23.44 psi P=

23.44 × 2 × 3 = 2,160 psi 0.065

Volumetric ﬂow through each tertiary runner can be calculated by subtracting the volumes of primary and secondary runners, or simply by adding the total tertiary runner volume and the total part volume and dividing by eight cavities: 0.096 + 7.32 = 0.927 in 3 8

10.14 Injection Molding Machine Nozzle

The volumetric ﬂow rate is 0.927 / 3 = 0.309 in3/s. γ=

4 × 0.309 = 1,611 (1/s) 3.1416 × 0.0623

The viscosity corresponding to this shear rate is 0.010 lb·s/in2, and: τ = 0.010 × 1,611 = 16.11 psi P=

16.11 × 2 × 1 = 515.2 psi 0.062

The total pressure loss from the sprue to each gate is the sum of the pressure losses through each segment: Pressure drop (total) = 3,107 + 2,163 + 515.2 = 5,785 psi These preliminary calculations show that the total pressure drop can be increased by using smaller cold runner diameters to hold a pressure drop below 10,000 psi, and reducing the total runner weight. By repeating the calculations for progressively smaller runner diameters until we reach the targeted pressure drop, we eventually come up with the ideal runner diameters, as shown in Table 10-9. In this cold runner analysis example (calculations omitted), the tertiary runner diameter cannot be reduced below 0.125 in, because of the restrictions of the molded part wall thickness. The secondary runner diameter should remain the same, because the pressure drop does not present a problem. However, the main or primary runner can be reduced from 0.156 to 0.140 in, causing the total pressure drop of the runner to increase from 5,785 to 6,036 psi (insigniﬁcant), while reducing the total runner weight from 7.37 to 6.78 g.

10.14.2 Determining the Injection Pressure Needed How much pressure drop can be tolerated by the time the melt reaches the cavity while making sure that there is adequate pressure left to pack out the part and achieve consistent dimensions? One assumption that seems to hold for most injection molds is that there should be at least 5,000 psi available to ﬁll a cavity. Since there are very few absolute rules in injection molding, this 5,000 psi pressure recommendation should be considered only as a “rule of thumb”. It is not recommended to run the molding machine at its maximum pressure capability. In our example, assuming an available pressure of 15,000 psi is a better selection. Because we have allowed 5,000 psi to ﬁll the mold, the runner system must be designed with a 10,000 psi pressure drop. With the calculations in the previous section we see that the pressure drop is 5,785 psi. This result also reveals that our initial estimate of runner diameter using the values from Table 10-9 were much too small for the length of the runners. We must go back and recalculate, based on larger runner diameters. We may start by using the pressure drop equation to recalculate pressure loss for a new primary runner diameter. To simplify the process, keep the shear-stress value constant and substitute a larger runner diameter value until we obtain a lower pressure drop. It is then necessary to go back to the equation for shear rate and recalculate to ﬁnd the actual pressure drop for the new runner diameter.

653

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10 Injection Mold Design As before, it may be necessary to recalculate new values for the secondary and tertiary runner diameters as well. The calculations should be repeated by assuming larger and larger runner diameters, until the targeted pressure loss is achieved. With some practice, this can usually be done manually in several hours or with the help of a runner pressure drop software program and a computer. The same runner system with larger runner diameters calculated gives an overall pressure drop of 5,000 psi. The result is that now we have a runner system engineered to perform the required task of conveying the melt at sufﬁcient velocity and low enough pressure losses to achieve the required part quality, while providing for the lowest volume of material in the runner itself. In conclusion, the above example is intended to show that properly sizing a runner system is a matter of simple arithmetic and trial and error. This computation might take several hours, but it might save many more hours during the start-up plus considerable downtime and expense if the mold does not have to be sent back to the tool builder for changes in the cold runner’s system.

10.14.3 Cold Runner Flow Tab The use of a ﬂow tab as shown in Figure 10-84 is helpful in maintaining consistent production. The ﬂow tab is made by machining a groove, longer than the expected ﬂow, about 0.25 in wide and 0.020 in thick that extends from a runner cold slug pocket. It is marked at 0.125 in intervals to show the extent of ﬂow into the tab. A venting at the end of the tab should be provided to avoid any restriction to ﬂow. The operator can check periodically to determine if the tab length changes, showing a change of molding conditions and part size. Runner vent

Runner Cavity

Runner vent

Sprue

A Flow tab

Figure 10-84 Mold cold runner ﬂow tab

1 2 3 4 5 6 7 8 9 10

0.020 inches A 0.25 inches Tab vent

A - A

10.15 Mold Cavity Gating

10.15

Mold Cavity Gating

The gate operates as a thermo valve; the valve opens allowing the hot melt to ﬂow at high injection speed and pressure to ﬁll the cavity and the valve closes when the melt stops ﬂowing. Designing the gate as small as possible to reduce its visibility on the molded product is always desirable. However, a small gate causes the most severe restriction to ﬂow into the cavity. Because of the especially high shear rates of the melt as it passes through the reduced gate area, the melt temperature increases and a substantial amount of the potential energy represented by the pressure on the material is converted to heat by friction. The effect on the material is drastic and, for shear-sensitive materials, there is substantial degradation in the molecular weight of the material as it passes through the gate. If a thermoplastic reinforced resin is compounded with ﬁber glass or mineral additives, the severe ﬂow patterns generated at the gate will break up the reinforcing materials and convert ﬁber to powder with a substantial loss in the reinforcing properties of the resin. The whole idea of gating is that the runners must form a good hydraulic system to allow sufﬁcient quantities of hot melt to reach and enter the gates with minimum heat loss. This allows the maximum injection pressure possible to be transmitted to the mold cavity without interruption of melt ﬂow through the gate. When the injection screw/check valve stops its forward travel, pressing against the melt channels, material in the small gates will freeze quickly, making it essential that all necessary material ﬂow through in one uninterrupted movement. Because the gates do freeze after the cavities are ﬁlled and packed under pressure, no additional material can be forced into the cavities. However, with large gates, forcing in additional material after the thermoplastic melt has partially set may result in a laminated condition around the gate as the hot melt moves over the relatively colder material that is already in the cavity. The gate size is important for the following reasons: • The gate freezes soon after the cavity is ﬁlled allowing the injection screw/ check valve to pack the cavity and control the quality of the molded product during the cooling cycle • The correct type of gate allows for simple gate separation from the molded product and automatic molding process • After de-gating, only a small witness mark remains on the molded part • Uniform quality control of multi-cavity molds can be achieved • The correct screw forward time (ﬁll and pack) with the gate open allows higher density, reduces mold shrinkage, and provides better dimensional control, increasing the molding process efﬁciency. The optimum type, geometry, dimension, and location of the gate are determined by the following factors: • The viscosity and shear rate characteristics of the resin to be molded • The shot size or volume of material to be injected • The melt and mold processing temperatures • The crystallinity rate or time required to freeze the melt in the mold cavity • The size, complexity, and wall thickness of the molded part

655

656

10 Injection Mold Design • Molded product requirements (ﬂatness, roundness, tolerances, surface ﬁnish, toughness, strength) • Type of injection molding process (manual, automatic, runnerless, two- or three-plate, vertical inserts, lost cores, air assist).

10.15.1 Types of Mold Cavity Gates To obtain the optimum injection molding process conditions, the type of gate must be carefully chosen. Gates commonly used are sprue gate, edge gate, tab gate, fan gate, diaphragm gate, external ring gate, internal ring gate, spider gate, ﬁlm gate, pin point gate, tunnel gate, and runnerless gate systems. Sprue Gate

Molded part

When the molded part is injected directly from a sprue, the feed section is called a sprue gate. The main disadvantage with this type of gate is that it leaves a large gate mark on the molded part, requires manual sprue removal, and single cavity molds. The size of the mark depends on the diameter at the small end of the sprue, the sprue angle, and the sprue length. The gate mark can be reduced by keeping the dimensions to a minimum. The sprue small diameter should be 0.031 in larger than the nozzle exit inside diameter to avoid sprue ejection problems and molding cycle interruptions. An extension nozzle sprue bushing (see Figure 10-72, top illustration) can often be used to an advantage, because it enters a recess in the mold and cuts down the overall sprue length. A sprue gate is illustrated in Figure 10-85.

Sprue

Figure 10-85 Sprue gate

Rectangular Edge Gate The edge gate is a general purpose gate with the simplest geometry; it has a rectangular channel machined in one mold plate to connect the runner to the cavity. The rectangular edge gate offers the following advantages: • The cross sectional geometry is simple and cheap to machine • Close accuracy in the gate dimensions can be achieved • The gate dimensions can be easily and quickly modiﬁed. The ﬁlling rate of the cavity can be controlled relatively independently of the gate freezing time. One disadvantage of this type of gate is that after gate removal, a witness mark is left on a visible surface of the molded part. This is more noticeable with certain materials when the molded parts are broken from the gates. Edge Gate Critical Dimensions Molded part L (Land) = 0.040 inches "W" (Width) “T” "D" Dia.

Because of the rectangular shape, the dimensions of the gate are given by width (W), depth (H), and land length (L) as shown in Figure 10-86. The pressure drop across the gate is approximately proportional to the land length and should be kept as small as possible to be consistent with the strength of the steel that remains between runner and cavity. A land of 0.040 in is recommended.

Runner H (Depth) = (40 - 60%) T

Figure 10-86 Rectangular edge gate

The minimum depth of the gate controls the time for which the gate remains open. This gate open time must be sufﬁcient for the melt to reach the extremities of the cavity. Providing the wall section of the component has been correctly

657

10.15 Mold Cavity Gating chosen with regard to the maximum length of ﬂow required, it is reasonable to expect a relationship between the gate depth dimension, the wall thickness of the molded part, and the crystallinity of the thermoplastic resin. In practice, a gate depth ranging from 40% to 60% of the molded part wall thickness is recommended. The cross sectional area of the gate (H × W) controls the rate at which the thermoplastic melt enters the cavity. Gate width (W) should be between the depth (H) and the runner (D) diameter, depending on the volume of the melt that must ﬂow through the gate. Overlap Edge Gate This gate is a variation of the rectangular edge gate and is used to feed certain types of molded parts. Figure 10-87 shows the overlap edge gate is machined into the plain mold plate so that it bridges the gap between the end of the runner and the end wall of the cavity. This gate may be used for all the common molding materials except rigid PVC. This gate, being attached to the molding surface, does require more careful removal and ﬁnishing than for edge gates. The size of the gate can follow the general pattern suggested for the rectangular edge gate with the same limitations. Fan Edge Gate The corresponding dimensions of the fan gate are not constant, the width increases from the runner diameter or from a perpendicular sub runner as shown in Figure 10-88, while the taper depth decreases to maintain a constant cross sectional area throughout the length of the fan component of the gate. The width of the gate at the cavity is relatively wide and, because of this, a large volume of material can be injected in a short cycle time. This type of gate can be used advantageously for large area, thin-walled molded parts. The fan shape appears to spread the ﬂow of the melt as it enters the cavity and a more uniform ﬁlling is obtained with fewer ﬂow lines and surface blemishes. The gate may be used for all common molding materials, except rigid PVC. The relevant gate sizes that must be chosen are the land length (L), the gate width (W), and the gate depth (H). The land length needs to be slightly longer than

Runner

Molded part “T”

"H" (depth) = (40 - 60%) T

"W" (width)

"L" (land) = 0.050 inches

Transition fan Sub runner

Figure 10-88 Various types of fan edge gates

"W" (Width)

"H" (Depth) = (40 - 60%) T Molded part

“T” "D" Dia.

Runner

"L" = (0.040 - 0.050) inches

Figure 10-87 Overlap edge gate

658

10 Injection Mold Design Molded part Runner Edge gate

for the rectangular edge gate, a suggested size of 0.050 in is ideal to trim the gate using a knife while the molded part is hot. The width used is often much bigger to obtain the full beneﬁts of the fan gate design. The main disadvantage of the design is that a large witness mark is left on the molded part that must subsequently be trimmed and ﬁnished. Designing the gate relatively narrow is therefore advantageous.

Remove tab after molding

To maintain a uniform pressure distribution inside the trapezoidal conﬁguration of the fan gate the depth must be blended, stream-lined, progressively increased, and tapered down accordingly.

Figure 10-89 Tab edge gate

The effective fan gate depth (H) between the transition sub runner and the cavity progressively increases from a minimum at the center line to a maximum at the outer gate width (W). To compensate for the pressure differential caused by a uniform fan gate depth, it is common practice to increase the depth of a wide fan gate at either side to provide for more even melt ﬂow through the fan gate. Tab Edge Gate

Molded part

A side tab perpendicular to the runner and edge gate is added to a molded part, as shown in Figure 10-89. The tab and gate are removed after mold ejection.

T + 0.040 inches "L" = 0.040 in.

This technique prevents the undesirable jetting problem by changing the melt ﬂow direction, forcing the melt to ﬁll uniformly in a smooth steady ﬂow depending on the shape of the cavity. The tab gate leaves large witness marks. This gate was developed for high viscosity amorphous resins, such as acrylic.

45˚ ”T“ "H" = (40 - 60%) T

Sprue Diaphragm Gate This gate is used for single-cavity tubular shaped injection molded parts. It may also be used for multi-cavity tubular parts using hot runner molds.

Figure 10-90 Sprue diaphragm gate

Diaphragm gates are ideal for molding round products requiring precision dimensional control (TIR, roundness) and optimum strength without weld lines.

Runner

The sprue leads into a circular distribution disc 0.040 in thicker than the wall thickness (T) and smaller than the part inside diameter, see Figure 10-90. This disc functions as a runner, allowing the melt to ﬂow radially for constant ﬁlling of the continuous circular gate. The center disc is sometimes tapered from the center toward the gate to save material, provide thinner wall thickness, efﬁcient mold cooling, and faster cycles. Sprue and gate are removed using a secondary post molding operation.

Molded part External ring

External Ring Gate Gate "H" = (40 - 60%) T “T” "L" = 0.040 in.

(T + 0.040) inches dia.

Figure 10-91 External ring gate

The function of this gate is similar in performance to the diaphragm gate, but it is recommended only for unreinforced, low melt temperature, low crystallization rate thermoplastic resins. This type of gate is used for tubular type molded parts when more than one cavity is required in a simple two-plate mold. The gate provides for feeding all around the external periphery of the cavity and permits the use of a conventional runner system to connect the cavities. A trapezoidal runner is normally used because this type of molding would be ejected using a stripper plate. The gate is like a concentric ﬁlm between the runner and the cavity. The dimensions of this gate are similar to the diaphragm gate. Figure 10-91 illustrates an external ring gate.

659

10.15 Mold Cavity Gating Internal Ring Gate

Internal ring

The internal ring gate functions similarly to the external ring gate. This gate is used for molding small tubular molded parts usually made of unreinforced or reinforced thermoplastic resins. The sprue is connected to the internal ring via two or more sub runners, feeding all around the internal periphery of the cavity through a continuous and concentric circular ﬁlm gate. Figure 10-92 shows a typical internal ring gate and its dimensions. Molded part

Spider Gate The spider gate functions similarly to the diaphragm gate but does not provide the same results. This gate is used for tubular molded parts. The spider gate uses two or more internal spider tunnel or rectangular edge gates to feed the internal periphery of the cavity (equal spaces) and permits the use of a sprue system (single cavity) or a hot runner mold to feed several cavities. Figure 10-93 shows a sprue connected to three spider tunnel gates and a sprue connected to four spider edge gates.

Spider sub runner (T + 0.040) inches dia.

"H" = (40 - 60%) T

Sprue

Gate “T”

Film Edge Gate

"L" = 0.040 in.

The ﬁlm gate may be considered a long rectangular edge gate. Film gates are used for injection molding large, thin-walled, good surface ﬁnishing thermoplastic components to help in the production of warp-free products.

Figure 10-92 Sprue, three sub runners with internal ring gate

Film gates are useful for those materials that exhibit differential mold shrinkage, for which reduced central feeding of the melt is impractical. The gate normally extends across the complete width of the molded part, although a smaller width may be used initially. The ﬁlm gate width should be increased and the molding process parameters optimized until the product is satisfactory and meets the speciﬁed requirements. The ﬁlm gate is similar in principle to the diaphragm, external, and internal ring gates that provide a uniform and continuous wide entrance for the melt ﬂow to ﬁll and pack the cavity without molding defects and operational problems. The ﬁlm gate depth may be less than for a corresponding rectangular edge gate if the melt ﬂow rate, the crystallinity rate, the shear rate, and viscosity do not exceed the boundaries and molding characteristics of the polymer. A parallel runner and a ﬁlm gate are provided at the side face of the cavity to feed the melt evenly. A slender steel wall exists in the mold, between the runner and the cavity; a minimum ﬁlm gate land of 0.050 in is suggested. A ﬁlm edge gate is illustrated in Figure 10-94.

Tunnel spider gaate (T + 0.040) inches dia.

"H" = (40 - 60%) T Molded part

Puller Sprue

"L" = 0.050 inches

Gate “T”

Sub runner

"W"

Runner

"L" = 0.040 in. Edge gate spider

Figure 10-93 Three-spider tunnel gate and four-spider edge gate "H" = (25 - 50%) T

Molded part

Figure 10-94 Film edge gate

“T”

660

10 Injection Mold Design Pin Point Gate Cold Runner System The pin point gate is a cold runner system that offers many molding process advantages, such as precision dimensional control of molded circular gears. This system causes little surface imperfections to the molded product. The circular gate marks are relatively small and insigniﬁcant for many applications. This system is used as an alternative method to gating into the center of a component. This gate system is often used for single or multiple pin point gating. It is desirable for resins that exhibit high differential mold shrinkage characteristics. The pin point gate connects the cavity to the conical sub runner, to a parabolic or trapezoidal runner, and to a sprue. A three-plate mold with three or four equally spaced pin point gates in each cavity has been used successfully. An extra plate is added behind the cavity plate to permit a runner system to be behind the mold cavities. This system is composed of a sprue, a sprue puller, parabolic or trapezoidal runners, ejection shoulders at the end of the runner, conical sub runners, taper/ round pin point gate, and sucker pins used to separate the molded part from the pin gate to eject the runner system automatically from the mold by using the same sucker pins. Because of the pin point gate geometry, it may be difﬁcult to control the melt ﬂow rate, melt process temperature, shear rate, and gate freezing time independently. A compromise between these melt behavior characteristics is needed to ﬁnd out which one of the thermoplastic resin’s properties are the most important for molding conditions to improve the process efﬁciency. Figure 10-95 details a pin point gate.

Sucker pin 15˚ T + 0.062 (dia.) Parting line Trapezoidal runner r.

r. 30˚ 2˚

T + 0.040 (dia.) T

0.045 inch Parting line

Molded part

30% T (dia.) 15˚/Side

Figure 10-95 Pin point gate runner system for a three-plate mold

661

10.15 Mold Cavity Gating Round Edge Gate The round edge gate is a matching semicircular channel machined in both halves of the mold, between the runner and the cavity. The round edge gate is recommended only for the injection molding process of melt processing rubber compounds (TPE), of molded parts with a minimum wall thickness of 0.150 in, and when dimensional control is not critical for the application. The round edge gate is an advantage for melt processing of rubber (TPE). The melt ﬂow rate is affected by the shearing force applied to the polymer melt during the injection molding process.

Runner Molded part

Round gate

Figure 10-96 Round edge gate

The round edge gate is an efﬁcient melt shearing device. By increasing the shear rate, causing the melt process temperature to increase by injecting the melt under high injection pressure and fast injection velocity, increasing the melt at higher ﬂow rates. The high melt process temperature and fast melt ﬂow rates used to ﬁll the cavities causes the round gate area to increase the temperature, forcing the round gate to remain open during the time required to complete the screw forward cycle to allow packing pressure to be applied to the cavity, thus increasing the density and reducing the mold shrinkage of the molded parts. Standard Tunnel Gates The standard tunnel gate is a circular or oval gate that submerges and feeds into the cavity below the parting surface of the mold. Feeding into the side of the cavity has several advantages: • No cavity matching problems using a two-plate mold. • The oval cross section controls the ﬁlling rate (cavity); the conical shape and the length (conical sub runner) control the freezing time of the gate. • The gate is sheared off from the cavity automatically during ejection. The runner stops at a steel safety margin distance from the edge of the cavity. A conical sub runner is machined at an angle between 30° and 45° and is stopped short of the cavity wall by a distance (gate land length). The gate is then machined or EDM (burned) at the same angle and tapered at 15° per side to join the conical sub runner and pin gate to the cavity. The molded parts and the runner systems are removed separately from the mold; this means that a separate runner ejection is advantageous, particularly as a certain amount of deformation of the runner is necessary to shear off and remove the cold runner system from the mold. The sequence of a tunnel gate operation is illustrated in Figure 10-97. The standard tunnel gates are found in two varieties: short tunnel and long tunnel gates. Generally, the short tunnel is preferred. When a long tunnel is used, the angle between the part and the tunnel is important. The angle should not exceed 30° to ensure gate shearing. The long tunnel gate should never be used with high crystalline resins or molded parts requiring precision dimensional control. The steel safety margin must be at least 0.078 in or greater. The short tunnel angle between the center line of the conical sub runner and the cavity wall is normally between 30° and 45°. The dimensions for the oval short tunnel gate cross section at the cavity side wall should be between 30% and 40% of the molded part wall thickness (T), based on the melt ﬂow rates, crystallinity rates, the modulus of elasticity, how much ﬁber glass or mineral reinforcement is used, and the melt viscosity of the thermoplastic resin.

Mold closed, runner, tunnel gate and cavity are injected

Moving half opens runner attached

Ejector pins shear off part and gate

Figure 10-97 Tunnel gate, sequence of ejection

662

10 Injection Mold Design

45˚

T

Steel safety margin

30˚

0.078 min. Taper sub runner 0.020 to 0.031 dia. + 15˚ taper/side

Gate

r. Parting line Sharp corner (40 to 60% x T) dia.

Ejector pin

0.062 r. 45˚

Short tunnel gate

Cavity

(1.20 x T) dia. Ejector pin

30˚

Figure 10-99 Du Pont type “C” tunnel gate

Taper sub runner

Ejector pin

Du Pont Type “C” Tunnel Gate

Gate

Long tunnel gate (not recommended) Figure 10-98 Standard tunnel gates

This type of tunnel gate was developed by Du Pont in response to the need for molding resins with high crystallization rates and to produce close dimensional tolerance molded parts and aid automatic ejection. This tunnel gate cannot be machined like standard tunnel gates. The tunnel type “C” oriﬁce is fabricated by ﬁrst building an electrode and then burning the tunnel as speciﬁed in Figure 10-99. The round gate cross section diameter is between 25% and 35% of the part wall thickness. The melt volume diameter behind the tunnel gate pocket should be 20% larger than the part wall thickness (T). This pocket is necessary to bring more melt heat and keep the tunnel gate open for mold shrinkage control, without producing high shear forces, gate surface marks, or ejection problems. Banana or Cashew Gate

Reverse taper ejector puller

Molded part

Gate insert Ejector pin

Figure 10-100 Runner with banana or cashew gate

Figure 10-100 shows a banana or cashew gate as a variation of a tunnel gate. Other than the standard tunnel gate, it can provide gating into the lower base of the molded component. Its primary limitation is that its curved shape forces the material in the gate to undergo considerable distortion during ejection. This requires the polymeric material to exhibit good ductility at the time of ejection. If a non-ductile material is used, its ductility can be increased by increasing the cross-section of the body of the gate to ensure that it is at an elevated temperature at the time of ejection. This method works best with amorphous materials that have a broad solidiﬁcation temperature. This type of gate can be purchased as a specially made, powder injection molded gate insert. It has a relief along the side walls, which will help thermally insulate the gate from the cooled mold. The warmer gate will remain more ﬂexible and thereby facilitate ejection. The dimensions for the gate openings should be 30 to 70% of the wall thickness. Parts formed with smaller gate openings will most likely have poor packing control and cause part shrinkage.

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10.15.2 Different Types of Hot Runner Gates The most critical part of all runnerless molds is the gate. The thermoplastic melt must be kept in a ﬂuid state up to the point of separation, while the area just beyond the gate in the cavity must freeze rapidly and without serious ﬂaws. Most runnerless injection molding problems occur in the gate areas. Runnerless Mold Hot Tip Torpedo Insert Gate The hot tip torpedo insert gate is a small conical spreader at the gate, which aids heat transfer to the melt and limits the formation of a vestige in the gate area. It contacts the inside diameter of the hot sprue bushing and draws heat from the torpedo cartridge heater. Several streamlined passages around the cone allow the passage of melt into the cavity. The purpose of this heated insert is to deliver heat to the gate separation point and to establish the exact location of the separation point. Hot Runner Mold Valve Pin Gate A valve gate is an open gate with a reciprocating valve pin. The valve is opened and closed during each cycle by means of a hydraulic, pneumatic, or spring mechanism.Valve gates provide maximum control of gate vestiges and cosmetics, but add complexity and introduce serious wear problems with ﬁber glass and/or mineral reinforced resins. Valve gates also tend to push cold slugs into the gate area of the molded part and these cold slugs may cause impact failure or surface appearance problems.

Figure 10-102 Two types of hot runner mold valve pin gates

Figure 10-101 Runnerless mold hot tip torpedo insert gate (Courtesy: Mold Master)

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10.16

Molded part

Gate Molding Effects

Ideally, the gate should be positioned to allow an even ﬂow of the melt into the cavity, so that it ﬁlls uniformly and the advancing melt front spreads out and reaches the various cavity extremities simultaneously.

Edge gate Runner

Weld line

Elongated in the melt flow direction

Figure 10-103 Molding effects caused by a single external edge gate

Runner Edge gate

The location and the type of gates used in the molds affect the geometry and dimensional size of the molded part. These dimensional variations become more difﬁcult to control with molding materials with high mold shrinkage characteristics. The differential values based on the ﬂow orientation and the types of reinforcements used to compound the polymers also affects dimensional control. The injection molding processing parameters and mold design also play an important role in controlling the ﬁnal dimensions of the molded parts. To illustrate the molding problems caused by the gate location, the following mold gate design cases have been prepared. To injection mold a thermoplastic bushing using a single edge gate, the melt is oriented parallel to the gate, and high pressure is applied at the opposite side of the gate. The round bushing becomes oval in the direction of the ﬂow and a weld line is formed. Figure 10-103 shows this molding effect caused by a single edge gate.

Molded part Edge gate Runner

Weld line

Weld line Elongated perpendicular to the gates

Figure 10-104 Effects caused by two external runners and edge gates

Molded part

If two edge gates placed at an 180° angle are used to mold a thermoplastic bushing; the molding effects are similar to the previous case. The injection pressure from both gates causes a dual high force 90° away from the gates. The molded bushing becomes oval or elongated (less than in the previous case), with two stronger weld lines. Figure 10-104, illustrates the molding problems caused by the edge gates and the locations. If two spider runners with edge gates are used to mold a bushing, the effect is opposite to the previous case. The injection pressure from both gates causes a dual high force in front of the gates. The bushing becomes oval, with two stronger weld lines 90° from the gates. Figure 10-105 shows the effects caused by the gates. If a four spider sub runner gate is used to mold a bushing, the effects are similar to the previous cases. The pressure from the gates causes four high forces in front of the gates. The molded bushing becomes wavy with four high points at the gates and low points between the gates, but the weld lines are stronger and the roundness much better. Figure 10-106, illustrates these molding effects on the part. To obtain optimum round dimensions of the injection molded thermoplastic bushing, the use of a sprue gate is recommended.

Edge gate Sprue Spider

Low point Molded part

Weld line

Edge gate

High point

Weld lines

Spider runner

Weld lines

Weld line Elongated in the melt flow direction

Figure 10-105 Effects caused by two internal spider runners and gates

Sprue

Low point High point Better roundness and weld line strength

Figure 10-106 Molding effects caused by four internal spiders and gates

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10.16 Gate Molding Effects Gate melt pressure causes core deflection Edge gate

Thinner walled side (melt deflect core) Center sprue gate

Runner

Runner

Weld line

Core Melt

Edge gate Flow direction elongated Poor type of gate selection and its location

Thinner wall thickness

Thicker wall thickness Good gate gesign

Core deflected

Figure 10-107 Molding effects caused by the type of gate and location

Such an ideal position for the gate is possible in certain geometries of circular cross sections, for example, a cup in which the melt is injected through a sprue gate at the center of the base. Figure 10-107, illustrates the molding effects caused by the type of gates and location on a molded cup. Another strong reason for central gating components, such as a pen cap, is that side edge gating may cause deﬂection of the core. Edge gating provides a more rapid melt ﬂow down one side of the cavity, resulting in a differential pressure that can move the core out of position. This creates a thinner wall in the cavity on the opposite side of the edge gate, adding another weakness to the weld line and incomplete molded pen cap or short shot. Central gating at the base of the pen cap cavity can overcome these molding problems; however, this solution may lead to complications in a multi-cavity three-plate mold or a hot runner mold. Figure 10-108 shows where gate type changes and relocation were necessary to eliminate the molding problems. For rectangular molded parts, there is no ideal gate position; even in this case, the central position gate is often preferable, unless the parts are thin-walled, particularly when the material used can exhibit differential shrinkage causing distortion. When the edge gate is used, and most molded parts are edge-gated for reasons of mold economics, it should be positioned so that the melt ﬂow immediately meets a steel cavity wall. When a rectangular cavity that is long and narrow is edge-gated at the center of the long end of the bar, the melt entering at high velocity and pressure produces a jetting mark. The melt front quickly sets on contact with the cool mold cavity walls. More melt then enters the cavity and ﬂows around the original jetted melt front. The resulting ﬂow lines are often visible on the ﬁnished molded bar.

Runner Edge

Runner

Short shot or incomplete part Poor type of gate selection and location Balanced melt forces around the core

Melt Balanced forces Core Uniform walled part Pin point gate (centrally located)

Three plate mold sub runner

Good mold design

Figure 10-108 Molding effects caused by gate type and location

Top side location

Runner The molded bar processing problems can be gate Overlap overcome by using an overlap gate or an edge marks gate located at the top side of the cavity as Jet flow shown in Figure 10-109. The ﬂow of material marks Edge gate from this type of gate or optimum edge gate location is forced to affect the side wall of the cavity causing the melt to form an advancing Melt flow front that progressively ﬁlls the cavity, displacing the air trapped inside the cavity. This The best melt flow distribution design Poor mold design Better melt flow distribution molding situation keeps the molded bar free of jetting marks on the surface. Figure 10-109 Molding effects caused by type of gate and location

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10 Injection Mold Design Both ends of bar warped

Hotter area

Hotter area

The edge gate located in the middle of the rectangular bar will cause bending problems, resulting from the high injection pressure of the gate at the middle far side of the cavity and the differential mold temperature between both sides of the cavity (see Figure 10-110, top illustration). If multi-edge gates are used on one side of the cavity (see Figure 10-110, bottom illustration), side surface waviness across the gates and weld line marks between the gates are formed.

Gate/runner

10.17

Poor mold design Bar side surface waved Weld lines

Four edge gates/runners Poor mold design

Figure 10-110 Effects caused by type, number, and location of gates

Mold Venting Systems

Injection molding of thermoplastics requires the material to be melted by frictional heat and then to be injected into a closed mold cavity at high pressure and velocity. As it enters the cavity, the melt must displace the gasses trapped inside the cavity at mold closing time. If these gasses are not expelled from the cavity, they could suddenly compress between 2 and 5% of the original volume, causing additional high pressure inside the cavity. This high pressure increases the temperatures well above those suitable for the injection molding process. These elevated temperatures cause local ignition (dieseling) of the thermoplastic melt and burning of the molded products. What is more dramatic are the mold deposits, the corrosion in the mold cavity tool steel caused by the ignition sparks (dieseling effect) and the degraded polymer gasses. The openings through which these gasses escape are known as vents; the ideal vent would be one that will allow gasses to expel freely from the cavity while completely blocking the escape of molten polymer, which would cause ﬂashing. Vents can take many forms and be strategically located at various areas of the mold. The type and location of the vent are determined by several factors, such as part geometry or product design, part end use requirements, mold design, injection molding machine, maintenance, process characteristics of the polymer, quality control, molding conditions, process efﬁciency, productivity, and part cost. When the mold has been completed without considering the venting requirements, incorporating the best type of venting in the mold may no longer be possible. Inadequate mold venting is a serious problem causing defective molded parts with ﬂashing, burned marks, weak weld lines, short shots, gas voids, sink marks, surface ﬁnishing ﬂow marks, and warpage (nonuniform shrinkage), creating quality control problems in the injection molding process. Lack of proper venting will cause excessive use of injection pressure for the molding process, which will cause a high degree of internal stresses. Vents not only provide an effective means of displacing the trapped air in a cavity by the thermoplastic melt, but also permit the escape of gasses caused by the degraded polymer. Improper mold vent dimensions used for a type of polymer or worn vents entrap gasses within the cavity that cannot be expelled before the injected melt solidiﬁes. Relief edge vents have to be small enough to prevent the polymer melt from entering the venting channels. A cavity can be considered sufﬁciently vented when the polymer melt is injected in the cavity at high speeds without any sign of burn marks on the molded part or corrosion spots in the cavity tool steel surfaces.

10.17 Mold Venting Systems In a mold with several gates or core pins, the polymer melt inside the front of the cavity is divided at each gate and around the core pins. These different melt ﬂows eventually converge and may therefore encapsulate the entrapped gasses causing weld line problems, gas voids, and burn marks. Weld lines often contain entrapped gasses causing weakness at the weld lines, especially if the melt fronts are colder after traveling a long distance inside the cavity. Venting at this location may be found necessary to improve the weld line strength, produce better surface ﬁnishing, and faster molding cycles. The venting problem is further improved by adding ring groove vents in the core pins and ejector pins. Venting the sprue puller, all the runner cold slug pockets, including runner ejector pins, also needs to be considered. When molding thick section parts, venting the mold using a variety of techniques may be desirable. Venting the mold is strictly an art, perhaps a lost one. Mold venting design is a function dependent on the geometry and end use requirements of the part, types of resin, mold design, molding machine, process conditions, and quality control speciﬁcations.

10.17.1 Product Design for Venting Much of the trouble can be avoided if venting problems are anticipated during the basic design of the part and the mold. The geometry of the part to be molded should be studied to decide the best location for the parting line, ejection system, and gating to achieve the best mold venting. In some applications, altering the design of the part to provide adequate venting may be necessary. Design modiﬁcations need not necessarily be extensive or affect the part requirements. For example, a slight increase or decrease in thickness of a section of the part may be sufﬁcient to create preferential ﬂow conditions for placing a desired vent. Several product design geometries can lead to venting problems: • A thermoplastic injection molded product designed with thin-walled sections connected and surrounded by thick sections should be avoided. Figure 10-111 shows that the last place to ﬁll the cavity using an edge gate would be the thin circular section in the center of the cavity. The melt inside the cavity starts ﬁlling the heavy sections next to the gate. When the melt front faces obstacles or restrictions such as thin walls, the melt temperature decreases and the melt ﬂow rate is lower. The injection pressure forces the colder melt to ﬁll the small openings or thin walled sections where the trapped air from the cavity has been moved. The trapped air is heated and compressed by the melt causing short shots and burn marks. If such geometries are needed, special parting face vents should be implemented through the cavity insert contact surfaces. • Blind, deep holes in the cavity should be avoided. Figure 10-112 shows this type of product design problem. When a center shaft cannot be eliminated, the shaft depth should be reduced and vented. • Avoid any geometry that could cause a preferential ﬁlling along the parting line or around core pins, where the melt ﬂow is divided and the unvented section of the cavity is last to ﬁll. Figure 10-113 shows how preferential ﬁlling will occur around the thicker upper cup lip portion of the part. Premature sealing of the parting line vents will result. This mold will be extremely difﬁcult to ﬁll if proper venting is not installed at the lower thin wall section of the cavity.

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Molded part

Gas voids Heavy wall section

Middle thin section

Edge gating

Gas voids

Figure 10-111 Venting problems, outer heavy/middle thin wall section

Molded part

Avoid deep blind holes

Burn mark

Edge gating

Figure 10-112 Venting problems with a cavity/center deep blind hole

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Edge gating

Thick upper wall section Deep thin base wall section

Short shot Figure 10-113 Venting problems of thick upper/thin base taper cone

10.17.2 Venting Characteristics of Thermoplastic Polymers The molding resin, the product geometry and end use requirements, and the molding process often inﬂuence the extent to which mold venting is needed. There are two characteristics of thermoplastics that lead to venting problems: • Thermoplastic resins that release large quantities of vapor when heated require more mold venting than materials that do not. • Thermoplastic resins that have sharp melting points, fast melt ﬂow rates, and high nucleation. For a fast injection, the mold must be well vented. The melt ﬂow rate inﬂuences the need for mold venting. A high melt viscosity or low melt ﬂow rate PE will require less mold venting than a more ﬂuid and fast setting polymer such as nylon.

10.17.3 Mold Deposit Problems Mold deposit is a hazy surface on the mold cavity; if the mold deposits are allowed to accumulate, they can blemish the surface of a molded part or reduce its glossy appearance, occasionally causing sticking problems. In severe cases, mold deposits can build up sufﬁciently to change critical molded part dimensions. Small core pins used in mold cavities generally function at higher mold temperatures than the rest of the mold cavity, thus being susceptible to mold deposits. Mold deposits can narrow an oriﬁce in a cavity, as in an aerosol molded part, or worse, seal it, causing part rejection problems. Reducing the mold deposit problem requires cleaning off the cavity or changing the molding resin to new low mold deposit thermoplastic compounds.

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10 Injection Mold Design Cleaning of Mold Deposits from the Cavity Where molding conditions cannot be changed to reduce mold deposits, a scheduled cleaning can prevent these deposits from being baked onto the mold cavities, making the mold deposit removal easier. • Vapor mold deposit removal has been successful by using aerosol cleaners like “Slide”, resin remover “The Stripper”, mold cleaner “Chem Trend”, the heat from a propane torch, or a hot air gun and chrome cleaners, and heavy duty cleaning compounds such as trisodium phosphates. If the mold deposit has accumulated for long periods, ultrasonic cleaning with trisodium phosphate or liquid detergents can be used • Additive deposits can also be removed with benzyl alcohol or a 50/50 mix of benzyl with isopropyl alcohol.

10.17.4 How to Avoid Venting Problems Flashing can also be caused by excessively fast closing of the injection machine moving platen and mold, trapping the cavities/runners’ air in dead/deep end sections in the mold cavities. This problem can be reduced by slowing the closing of the mold by controlling the clamping speed. • Set the return stop position control of the moving platen at a minimum clamp open position to eject the molded parts with a safe working clearance • Set the transition speed position control of the moving platen 0.50 in from the mold ﬁxed half • Set the last speed position control of the moving platen between 0.062 and 0.125 in before closing the mold • Set the platen to high speed and low pressure starting at the mold open position on the control placed 0.50 in from the mold ﬁxed half • Set the platen to low speed and low pressure for the travel distance between 0.062 and 0.50 in • Set the platen to very low speed and high clamping pressure during the ﬁnal travel. At this stage, the mold is closed by high clamping force and is ready for injection, where the mold vents allow the trapped air and polymer volatile gasses to escape into the atmosphere. However, improper venting can cause ﬂashing or poor weld lines on the molded part. These are usually a sign of late venting, after the polymer has started to become solid. When venting a mold, a slight increase in the shot size is necessary as material may be dragged out of the cavity by the partial opening of the mold. As a rule, if there is no puff of steam or vapor escaping the mold, then the mold vents have not expelled the volatile trapped gasses from the cavities. Another cause of ﬂashing is the excessive clamping force applied to the mold cavity inserts. This high force closes the parting line cavity vents and seals off the escape of the gasses causing a partial opening of the mold. Any molding condition that will result in increasing the rate at which the mold cavity will be ﬁlled will increase the need for better mold venting.

10.17 Mold Venting Systems These are the molding conditions affecting high shear rates: • Higher injection pressure • Faster injection speed • Higher polymer melt temperature • Higher melt residence time inside the plastifying unit • Larger injection molding machine melting capacity • Higher mold temperature • High compression plastifying screws • Higher nozzle temperature • Molding machine hold-up spots in the barrel, screw, check valve, adapter, and nozzle

10.17.5 Planning Mold Venting A typical product development problem results when the mold venting is neglected during the mold design stage by a forgetful mold designer or during its construction by a hurried toolmaker. A mold with inadequate or nonexistent venting will frequently be the cause of poor part surface appearance and/or performance, high reject rates, and lowered productivity. The project anticipated as a proﬁtable injection molding venture, starts affecting customer production requirements and consuming huge amounts of proﬁt. Mold venting design should be studied and carried out during initial design and construction of the mold. Location of the mold parting line, the type, dimensions, number and location of the gates, including the best venting system, needs to be addressed. Preliminary mold design layouts offer greater possibilities to install proper vents at correct locations to eliminate trapped gasses inside the mold cavities. This work should be completed before accepting the mold, because during production there is not enough time to take the mold out of the injection molding machine to incorporate the correct venting. Unless the molded parts are rejected for ﬂashing, burn marks, weak weld lines, short shots, poor surface ﬁnishing or voids, convincing anyone involved in the molding process that missing vents in a mold can be causing these kind of molding problems is often difﬁcult. An experienced molder knows that gasses trapped inside the cavities are the cause of these problems, but because production demands often allow only a makeshift form of mold venting (a portable grinder, masking tape, reduced clamping force, lower injection pressure), the molder is usually ignored. One method used to verify that the burning problems on the molded parts are caused by trapped gasses inside the cavities is called the “dieseling effect”. The trapped gasses are compressed very quickly by high pressures and high temperatures, causing the gasses to ignite (sparks). To determine if the mold has venting problems, use a natural color resin for testing the mold, spray an aerosol lubricant such as “WD-40” onto the mold cavity surface, then close the mold and inject the melt using high speeds. The ejected molded part should be blackened in the areas where the gasses and “WD-40” spray were trapped causing an explosion and burning of the molded part.

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10 Injection Mold Design If an injection mold is poorly vented, molded parts are likely to suffer from weak weld or knit lines, because the trapped gas keeps the melt ﬂow fronts from fusing together when the melt is split. Occasional short shots, internal gas voids or sink marks, burn marks, and surface ﬂow marks can occur. Progressively, these problems become more frequent and severe because improper vents tend to clog with condensed gasses from the polymer. What is not apparent is the mold corrosion problem that may occur in the cavity’s deep pockets or on the surrounding surfaces of the mold base plates or ejector pins. While most polymers’ melts may be totally non-corrosive, high melt temperatures, high injection pressures, and fast injection speed cause degradation of the plastic at the point where air and gasses are trapped in the cavities with each shot, causing an internal explosion (dieseling) and, in time, pitting corrosion may result in the tool steel of the cavities.

10.17.6 Mold Venting Process Problems Inadequate venting or a mold without vents causes many injection molding problems:

Parting line flashing

Figure 10-114 Poor mold venting causes ﬂashing molding problems

Inclomplete part

• Flashing problems are the most common defects in the injection molding process. Parting line ﬂashing is the result of a fast clamping speed while closing the mold, an excessive injection speed and pressure, high melt and mold temperatures, long residence time of the melt in the plastifying unit, wet material or high moisture absorption of the resin, excessive amounts of reground and contaminated resins. The ﬂashing problem is caused by the high temperature and pressure required by the melt to overcome resistance to ﬂow created by compressed gasses. The ﬂashing problem will not be eliminated by decreasing the mold temperature, because of the fast crystallization of the melt. Venting the mold correctly to remove the trapped gasses is necessary, using the recommended melt temperature, injection speed, and mold temperature. These molding conditions will allow an efﬁcient melt quality, uniform crystallization, faster cycles, and quality molded parts that comply with production requirements. Figure 10-114 shows a typical parting line ﬂashing molding problem.

Figure 10-115 Poor mold venting causes incomplete molded parts

• Short shots are a result of the inability to ﬁll the cavities completely. The main causes of this problem are poor venting or molds without vents. The check valve may not hold pressure, small nozzle oriﬁce, small gate size, low melt/mold temperatures, low injection speed and/or pressure and a thin walled part with long cross sections are also causes. Figure 10-115 shows a typical example of a short shot molding problem.

Internal gas voids

Figure 10-116 Micro structural analysis showing internal gas voids

• Internal gas voids are found in thicker wall sections and/or in thin sections surrounded by thick wall sections of a molded part. Figure 10-116 shows a micro structure analysis of a gas void problem found in a molded part. Sink marks and internal gas voids (porosity) in a molded part are caused by a thick walled section molded from a high mold shrinkage polymer at a high mold temperature. This causes the surface of the molded part to crystallize or to become solid (skins), while the melt inside the frozen surface wall is still in a liquid soft stage. Voids are also caused by high melt temperature, a short screw forward time or lack of packing the cavity, check valve leakage, small nozzle oriﬁce, and small gate size.

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10.17 Mold Venting Systems • Poor weld lines can be developed behind core pins or obstructions that split the melt stream, causing two moving melt fronts to come together during the ﬁlling of the mold cavity. Ideally, the melt should ﬁll the mold rapidly enough so that the melt front interfaces are hot enough to cause a strong weld line. Good venting in front of the weld line, proper gate size and location, fast injection speed, and resins without lubricants will insure a strong weld line. Avoid the use of mold release spray in the mold cavity surface; this lubricant is a main contributor to the formation of weak weld lines. When the cavity is inadequately vented, the cavity ﬁlling can be slowed if the interfaces cool off excessively, resulting in a weak weld line. Figure 10-117 shows two micro structure analyses of poor weld line problems found in two thermoplastic molded parts. • Burn marks are molding defects of molded parts caused by highly compressed hot gasses that could not escape from the mold cavity fast enough. The molded parts are burned at a location that was the last place to ﬁll the cavity, where the pocket did not have sufﬁcient venting. The compressed trapped gasses increase the mold temperature (diesel effect) causing the burning of the thermoplastic. High melt and mold temperatures, excessive injection speed and pressure are also contributors to this burning problem. Figure 10-118 shows a burn mark problem found in a thermoplastic molded part.

Weld line

Figure 10-117 Micro structural analyses showing the weld lines Molded part Burn mark

Figure 10-118 Burn mark problem of a thermoplastic molded part

• Corrosion of the metal used to construct the mold cavity insert can be caused by the degraded, decomposed thermoplastic melt produced when the trapped gasses within the mold cavity are highly compressed. This generates temperatures far above the thermal stability of the thermoplastic causing an ignition (sparks) every molding cycle, corroding the mold cavity tool steel. Figure 10-119 shows a corrosion problem found in a mold core insert. • Poor surface or rough ﬁnishing on a thermoplastic molded part is caused by the inability to ﬁll the mold cavity quickly. The trapped gas slows down the melt ﬂow speed rate, cooling off the melt because of poor venting or no venting. The process conditions such as low screw rpm, a low screw back pressure, check valve leakage, small nozzle oriﬁce, small gate size, low melt/ mold temperatures, and low injection speed/pressure may also cause poor surface ﬁnish of the molded parts. Figure 10-120 shows surface defect problems found in two thermoplastic molded bushings. • Mechanical, electrical, and chemical resistance properties can be decreased by poor mold venting. Locally burned areas, gas voids or weak weld lines caused by inadequate mold venting reduce the tensile strength, impact resistance, and dimensional tolerances of thermoplastic molded parts.

Weld line

Tool steel corrosion

Mold cavity core insert

Figure 10-119 Core insert corrosion caused by poor mold venting

Figure 10-120 Two molded bushings showing poor surface ﬁnishing Mold depsit spots

• Long molding cycles caused by higher melt/mold temperatures, extended cooling time, long melt residence time, and hold-up spots also cause mold deposits. • Mold deposits in various compositions accumulate on the mold cavity surfaces. There are different types of mold deposits: they can be clear or opaque and can range from white to dark, waxy to hard. Product geometry, type of resin, mold design, and molding process play an important role in controlling the mold deposit problems. Figure 10-121 shows a mold deposit problem found in a molded cup.

Figure 10-121 Molded cup showing mold deposit problems

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10 Injection Mold Design Vapor Deposits in the Cavity The most common deposit develops as vapors from low molecular weight resins react with the trapped gasses in the cavity. For example, when vapors from formaldehyde polymers (acetal resins) are trapped in a mold cavity, deposits usually form at the last place the cavity was ﬁlled in a poorly vented mold. The faster the molten resin enters the mold cavity, the higher its mold temperature and the higher the injection pressure, the higher the shear forces applied to melt; these factors contribute to the mold deposit problem. To reduce the velocity for compressing the trapped gasses in the mold cavity, the following molding conditions are suggested: • Slow the injection speed • Reduce melt residence (hold-up) time inside the plastifying unit • Reduce the melt temperature • Reduce mold cavities’ hot spot temperatures • Improve the nozzle temperature control • Improve the design of the mold venting system • Eliminate molding machine hold-up spots in the barrel, screw, check valve, adapter, and nozzle. Additive Deposits in the Cavity Additive deposits are caused when an ingredient in a speciﬁc thermoplastic resin formulation separates during processing and forms a ﬁlm on the mold cavity surface. Additive deposits usually appear at the gate or other ﬂow obstructed areas. Thermoplastic resins containing plasticizers, stabilizers, lubricants, toughening agents, incompatible pigment concentrates in the carrier master batch in combination with high shear forces are the causes for the separation of additives.

10.17.7 Mold Venting Design Several areas of the mold can be used to expel trapped gasses. The mold venting efﬁciency will depend on many factors, such as product geometry or part design, part end use requirements, mold design, injection molding machine, maintenance, process characteristics of the thermoplastic polymer, quality control, molding conditions, process efﬁciency, productivity, and molded parts manufacturing costs. Mold venting efﬁciency also depends on the types and number of vents used, locations of the vents, and physical dimensions of the vents used in the mold cavities. Also a factor is the type, geometry, dimensions, number, and locations of the gates; the type and size with a balanced runner layout system properly vented around all cold slug pockets and runner ejection pins. The type of equipment used for grinding or machining the vents, craftsmanship, and careful construction of the mold is also important. There are several venting methods available for thermoplastic injection mold applications, including parting line cavity venting in line or perpendicular to the ring groove vent. There are two types of cavity insert parting face venting, one horizontal and the other vertical. Both are perpendicular to the ring groove vent.

10.17 Mold Venting Systems A stripper plate with ring groove vents is found either at the horizontal parting face position or a taper parting face position (one thirty o’clock). Ejector rings are also vented like the stripper plate; also used are vented core and ejector pins modiﬁed with ring groove vents, vented ejector blades, ejection assist poppet air valves modiﬁed for venting the body using a ring groove vent, sintered porous vent, magic seal negative coolant vent, and mold vacuum vent. Optimum venting methods are selected after a complete study of molded part end use requirements, mold venting design methods, cost of the mold, and additional auxiliary molding equipment needed. The vents should be ground and polished in those areas where the relief edge vent sections will trap air or gas. The air in the mold is at 330 to 380 °F and will get much hotter when compressed by the fast injection speed of the melt, which is hot enough to oxidize and erode the tool steel from the mold cavity and core inserts. There have been instances where enough heat was developed with high injection pressures, fast injection speed and a mold without vents to cause molded parts to burst into ﬂame upon ejection from the mold cavities. Cavities used for the molding of parts of simple geometry, as for example, bars, discs, or plaques of uniform thickness, do not cause any venting problems. However, by venting thin-walled parts, the trapped gasses are expelled through the parting line cavity vents, parting face cavity insert vents, ejector pins, and runner’s cold slug vents. Deep draw items such as containers, boxes, etc., if center-gated, will not cause venting difﬁculties. Occasionally, it may not be possible or desirable to center-gate a deep draw item. When such parts are side- or edge-gated and the ejector pin marks are also undesirable, the problem of mold venting is not easily overcome, especially when the mold is already constructed. Even a well vented injection mold will not last forever and as aging of the mold occurs, some adjustments will need to be made. An injection mold cannot have too much venting. Molds are most commonly vented at the parting line of the cavity. Unfortunately, this may not be the location where the gasses are trapped. However, several other venting techniques are used to remove the trapped gasses in deep blind pockets. 10.17.7.1 Parting Line Cavity Venting Parting line cavity venting allows the trapped gasses from the cavities to escape without causing any melt ﬂow restrictions. The thermoplastic melt can ﬁll and pack the cavities well, providing a good surface ﬁnishing, strong weld lines, maximum mechanical and electrical properties, fast molding cycles, and optimum molding quality. Parting line ﬂashing is caused by the internal gas pressure being hotter and greater than the molding machine clamping force, pushing the moving half of the mold apart, releasing the compressed hot trapped gasses through this small opening with the melt causing parting line ﬂashing problems. A well designed mold venting system conveys the trapped air through the runners and gates entering inside the cavities after closing the mold. The trapped air and the polymer volatiles are expelled outside the mold by the melt front high injection speed and pressure. Parting line venting is accomplished by grinding or machining the mold cavity inserts contact area that forms the mold parting line. The cavity inserts should

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10 Injection Mold Design

Figure 10-122 Mold base plates and cavity insert mount clearances

be mounted between 0.003 and 0.005 in above the level of the mold base cavity plate. This insert contact area is used to create an effective seal around the cavity, making sure the parting line venting functions without any problems and reducing the chance of ﬂashing the cavity. Figure 10-122 shows mold base plates and cavity insert mount clearances. A parting line cavity relief edge vent is a precision engineered opening machined or ground starting at the parting line cavity outside wall until reaching the escape or bridge vent channel. The cavity relief edge vent consists of the following components: • Vent land length. The length of the vent is 0.040 in. This dimension has been used for venting the molds creating a proper pressure drop to allow the air to convey while stopping the melt ﬂow. This dimension has shown good venting results for many years. • Vent land depth. The depth of the vent is directly related to the polymer melt viscosity and ﬂow rate. A high melt viscosity with slow ﬂow rate requires a deeper vent land channel, than a low melt viscosity with a high ﬂow rate (see Table 10-10). • Vent width. The width of the vent varies with the cavity/runner volume and the number of vents used (0.093 to 0.187 in). There are two types of parting line cavity vents found parallel to the cavity plane surface; one vent is in line with the escape vent channel, while the other vent is perpendicular to the ring groove vent as shown in Figures 10-123 and 10-124. A parting line cavity vent (in line) is recommended for the following injection molding applications: • Irregular geometries of the cavities projected areas • Single or multiple edge gates • Melt ﬂow ﬁlling and packing the cavity from the edge gate wall to the cross section last to ﬁll; the melt ﬂows inside the cavity like parabolic waves, with the maximum stream velocity in the center and zero velocity at the cavity sidewalls • Melt ﬂow ﬁlls the thicker cross sections next to the gate ﬁrst and the restricted or thin cross sections last • Trapped gasses are conveyed directly to the outside of the mold through the parting line cavity escape vent channels.

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10.17 Mold Venting Systems Parting line cavity vents perpendicular to the ring groove vent are recommended for the following mold applications: • Symmetrical round cavities projected areas with uniform wall thickness • Center gates (diaphragm, spider gates) • Three-plate mold with pin point gates equally spaced concentrically around the center of the cavity • Cavity uniform melt ﬂow ﬁlling • Number of relief edge vents and bridge groove vents equally spaced and perpendicularly connected to the ring groove vent • Ring groove vents around the parting line cavity insert border area • Escape vent channels from the ring groove vent to outside the mold • Compression strength for sealing the cavities inserts contact area minus venting area is less than 20% of the yield strength of the metal used for the cavity inserts. Parting line cavity venting design details required for both types of vents and for various types of polymers are shown in Table 10-10 for common types of resins and in Figures 10-123, 10-124.

Rectangular edge gate

Relief edge vent

L = 0.040 land

Runner H = Depth (see table 10-10) 0.022 escape vent channel depth

Cavity 0.093 to 0.187 width Escape vent channel

0.06 r.

Figure 10-123 In-line parting line cavity venting (in)

0.022

Ring groove

0.06 r. (Typ)

Bridge groove

0.093

L = 0.040 land 0.375 min. increase if needed

Relief edge vent

PL

Cavity

H = Land depth (see table 10-10)

Figure 10-124 Parting line cavity venting perpendicular to the ring groove vent (in)

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10 Injection Mold Design Table 10-10 Cavity Relief Edge Vent Depth for Common Types of Resins

Cavity vent Hole

Cavity vent Runner

Edge gate

Cavity vent

Spacer

Cavity vents

Resin

Depth “H” (in)

Acetal ABS Acrylic Nylon unﬁlled Nylon glass reinforced Nylon mineral reinforced LCP glass reinforced PC PE PP PS PPO PSU PET glass reinforced PVC ridgid PVC ﬂexible

0.001–0.002 0.002–0.003 0.002–0.003 0.0005–0.001 0.0007–0.0015 0.001–0.002 0.0003–0.0007 0.0015–0.003 0.001–0.002 0.001–0.002 0.001–0.002 0.001–0.003 0.001–0.002 0.001–0.0015 0.001–0.003 0.001–0.002

Figure 10-125 shows four mold layout illustrations, where both types of parting line cavity venting systems are used in these mold venting designs.

Cavity vents

Mold Cavity Insert Parting Face Venting Vented ejector pins

Core pin ring groove vent (typ.)

Cavity vents

Runner

Fan gate Escape vent hole

Cavity vents

Cover plate

Overflow tab

Ring groove Bridge groove

Several insert components are used to form a complete vented cavity insert used in mold construction. This method is used when certain forms of a molded product require minor cavity changes to mold another product. Replacement of damaged or broken components of the cavity insert may require replacing only the critical components of the cavity without having to replace the complete cavity, or it can be done when installing vents in the cavities, because machining it as an integral component of the mold cavity is difﬁcult. These cavity insert components are fabricated separately from special steels or the same material as the cavity, and the inserts are installed in the proper position and become part of the mold cavity insert unit. Often, they are called integral cavity or core inserts. They are also used in the formation of narrow slots to produce vented support ribs on the molded part. One advantage of constructing the cavity with one or several inserts separated from the mold base plate is the cavity insert joint surface walls can be used for the installation of vents to improve the venting system of deep and complicated cavity geometries. Ring groove

Bridge groove Pin point gates

Tab vent

Relief edge vent Diaphragm gate Bicycle wheel

Escape vent hole Escape vent groove

Relief edge vents Escape vent groove Spur gear

Figure 10-125 Parting line cavity venting systems

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10.17 Mold Venting Systems The metal-to-metal contact area (parting line) outside the cavities between the two insert cavities mounted in each half of the mold base must not be reduced excessively by the venting channels. If the cavity inserts are raised above the mold base surface, the insert contact area outside the cavities should be large enough to keep the compressive stress less than 20% of the yield strength of the tool steel used for the mold cavity inserts.

Engraving insert

Insert parting face venting

Figure 10-126 shows four cavity insert parting face vent designs using various insert components to form a complete vented cavity insert. The illustration a) shows a replaceable engraving cavity insert. The vertical parting faces between both inserts can be used to install vents at the deep surface of the cavity and escape vent channels ending at the bore. The side of the bolt hole could be modiﬁed to machine escape vent channels.

Engraving cavity insert

a)

Insert parting face venting

The illustration c) shows a cavity insert with a bolted replaceable side insert and a bottom insert. The cavity inserts can be used to install parting face vents and escape vent channels. The illustration d) shows a cavity insert made by using two inserts. The contact area between these inserts could be used at the bottom of the cavity to install parting face cavity insert vents and escape vent channels.

Insert

The illustration b) shows a cavity for a deep and thin ribbed part made by using three additional insert components. These parting face cavity inserts at the bottom rib walls can be used to install vents and escape vent channels.

Insert

Deep, thin ribed cavity insert Replaceable insert

There are three types of parting face cavity insert vents, based on the plane of contact between the inserts: horizontal, vertical, and vertical tapered with respect to the cavity plane surface. The escape vent channel should be connected in line with the parting face cavity insert vents. Figure 10-127 shows the horizontal parting face cavity insert vent details. Figure 10-128 shows the vertical (or vertical tapered) parting face cavity insert vent details. The vertical tapered parting face insert vent is used for venting through the internal taper hole (matching the core insert shut-off dimensions) of an ejector ring or a stripper plate.

b)

Insert parting face venting

Bottom insert

Top replaceable cavity insert

c)

Figure 10-129 shows two different molds and their cavity insert venting systems; these two applications incorporate various venting designs, such as parting face cavity insert venting, parting line cavity venting, and core and ejector pin groove venting systems.

Bottom insert Insert parting face venting

Deep, recess cup cavity insert

d)

Figure 10-126 Different mold cavity insert parting face venting systems

Figure 10-127 Horizontal parting face insert cavity venting (in)

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10 Injection Mold Design H = Land depth (see table 10-10) L = 0.040 land

PL

Cavity Cavity insert

0.06 R. (Typ.)

0.375 min. increase if needed

Bridge groove

0.022 Ring groove

0.093

Cavity insert Escape vent groove

Figure 10-128 Vertical parting face insert cavity venting (in)

Escape vent groove Ring groove

Bridge groove Deep cavity with two fixed cores cavity insert venting design Parting line venting with escape groove

Parting line ring groove venting

Back groove escape vent

Vented cores

Insert face venting with escape groove

Back groove escape vent

Vented core Clearance vent (0.022)

Insert face ring groove venting

Vented ejector pins

Spur gear with two sets of teeth cavity insert venting design

Figure 10-129 Two mold illustrations of cavity insert venting systems

10.17 Mold Venting Systems

Figure 10-130 Ejector ring and parting line cavity venting system

Ejector Ring Venting System The ejector ring venting principle is similar to the vertical parting face insert or the stripper plate venting systems. To avoid the use of a large stripper plate, each cavity is provided with an individual vented ejector ring. This system is used to eject deep and thin-walled containers from the cavities. Figure 10-130 shows the parting line cavity venting, combined with a vertical taper parting face insert venting between the ejector ring internal taper wall and the core insert lower external taper wall. Core and Ejector Pin Ring Groove Venting The ring groove venting is the most widely used venting technique for round cores and ejector pins. This venting system is similar to the relief edge vent used to vent the cavities. The land (L) is 0.040 in, the depth (H) is the clearance between the ejector sleeve inside diameter and the pin outside diameter, the width is the circumference length of the pin. Round pins, such as stationary core pins and ejector pins, can be modiﬁed by machining a ring groove connected with a longitudinal ﬂat on the shaft and head surface of the pin.

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H = Clearance, ejector sleeve inside dia. and pin outside dia.

Longitudinal flats based on pin dia.

+ 0.0015 - 0.0000

0.022 depth, longitudinal flat escape vent (typ.)

Pin outside dia.

Pin head

+ 0.000 - 0.001

Ejector sleeve or insert block for both core and ejector pin 0.093

Cavity surface

0.022 0.040

Figure 10-131 Core or ejector pins ring groove venting details (in)

L = 0.040 land 0.062 r. groove

Figure 10-131 shows the design details for stationary core pins and ejector pins. This method allows venting the mold cavity in hard to reach areas away from the parting line. The ring groove vented pin is self cleaning and effective for removing the trapped gasses from the cavity without causing ﬂashing problems. Figure 10-132 shows a core or ejector pin ring groove vent projection view. For ejector pins ring groove vents, the length of the longitudinal ﬂats should only clear the base of the mold support plate to discharge the trapped gasses away from the mold base. The ejector pins are precision made of thermal shock resistant H-13 or CM-50 tool steel. The heads are hot forged for uniform grain ﬂow and high tensile strength, then annealed and ﬁnished for easy machining the ring groove venting. A ﬁnal ﬁnishing operation (including ring groove venting) is performed on the full length of the ejector pin to reduce wear and prolong pin life. The higher core hardness makes the ejector pins ideal for use in high melt/mold temperature injection molding applications. One or two longitudinal vent flats 180˚ from each other. Ejector pin length to clear support plate bottom wall

The following are recommendations for the selection of ejector pins and to improve mold performance for an efﬁcient molding operation:

Circumferential ring groove

• Avoid the use of small diameter ejector pins in high volume and critical injection molding applications. Small diameter and long pins wear off quickly and have the tendency to break during molding • Core hardness of 50–55 Rc reduces nicking, dishing, and bending

Cut one/two escape vent head channels only for the core pins

Figure 10-132 Core or ejector pin ring groove vent projection view

• Unchippable surface treatment of 65–74 Rc reduces ﬂashing problems • Annealed heads (30–35 Rc) and good ﬁnishing permit easy machining • Final surface ﬁnish (5–10 RMS) reduces wear and prolongs pin life.

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10.17 Mold Venting Systems

Figure 10-133 Two screw bosses using cores with ring groove vents

Figure 10-133, left illustration, shows how to vent a counterbore hole cavity using an interlock core with ring groove venting; the right illustration shows how to vent a boss (for self-tapping screws) cavity using a ring groove vented core. The vented cores improve the molded bosses’ weld line strength, sizes, ﬁnishing, and they are completely ﬁlled (good density) without burn marks. Rectangular Ejector Blade Venting Narrow rectangular ejector blades are used to push out ribs and thin sections of a molded part from the cavity. Ejector blades are thin components that do not have sufﬁcient area to install vents around their sides. The blade ﬁts into the sleeve ejector and is mounted into the ejector plate. As the mold ejector unit is activated, the ejector blade pushes the molded part off the cavity. One method for venting through these thin ejector blades is to taper grind two ﬂats starting from the top corners, as shown in Figure 10-134.

0.50˚ Taper

This venting technique has not been well implemented or proven in thermoplastic injection mold applications. This method could be acceptable for thermoset compression molds, because the transfer melt ﬂow, ﬁll pressure, and mold temperatures are relatively lower than the thermoplastic injection molding process.

0.003

0.006

0.062 Polished surface Length

Air Poppet Valve/Stripper Plate Cavity Venting The air poppet valves are designed to help with the vacuum problem often encountered during injection molding of deep draw and thin-walled parts. The air poppet valve prevents mold damage caused by product ejection problems. Air ﬂow (timed to coincide with the ejection cycle) opens or pushes the poppet valve to break the vacuum between the core’s deep surface and the inside wall of the molded product, starting the ejection process. To complete the part ejection

Width Polished surface

Figure 10-134 Vented rectangular taper ﬂats ejector blade (in)

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10 Injection Mold Design automatically, the poppet valve should work in coordination with other ejection systems, such as ejector pins, ejector rings, or stripper plates. The air poppet valve should never be used as the sole means of part ejection, or as an alternative to replace other standard ejection methods. The precision ground poppet valve body outside diameter can be modiﬁed by machining a ring groove and escape vent channels to provide venting at the deep section of the cavity, while the poppet valve seat prevents the melt from entering the valve during injection. Each poppet valve is matched to the body to assure reliable performance. Air Poppet Valve Operating Conditions The poppet valve air pressure and machine ejection should be activated simultaneously to relieve a vacuum build up in the cavity during part ejection. The air ﬂow to the poppet valve must be fully released to the atmosphere after each cycle to insure that the poppet valve closes. Thermoplastics melt injected into a partially open poppet valve could cause damage to the valve and/or the mold. Stripper plate ejection is generally used where ejector pin marks would be objectionable on the molded product and where a maximum ejection surface is required. Tapered parting face vented stripper plates are used on single and multiple cavity molds to eject round or rectangular parts from the male core

Figure 10-135 Air poppet valve/stripper plate cavity venting system

10.17 Mold Venting Systems cavity inserts. An angle of 5° is machined in the stripper plate parting face; the vent is located between the stripper plate internal tapered wall and the insert core lower walls. Figure 10-135 shows how to vent the air poppet valve and the stripper plate to avoid venting problems, automatically ejecting the molded parts without defects or interruptions in the molding cycle. UniLifter® Internal Undercut Ejection Venting Standard UniLifter® components (undercut releasing system) simplify mold design and construction for release of molded products with undercuts. A variety of sizes of standardized components are used in many mold applications. The smooth travel of U-Coupling in T-Gib eliminates heel binding. The radiused dovetail design lets the core blade seat automatically at the required angle. The UniLifter® ejector blades are made from H-13 tool steel with a hardness of 38–42 Rc, designed for conventional machining and grinding of the vents at the upper surfaces below the lower cavity surface. The type of venting used for this system is determined by the cross section of the ejector. For rectangular blades, thickness and width is supplied surface ground with a maximum of 0.010 in additional stock over nominal for ﬁtting into the insert pockets and/or to ﬁt a nominal size product. For round core blades, diameters are supplied with a tolerance of (+0.000/–0.001) in for ﬁtting in a bored hole or bushing. The following are mold design guidelines for the UniLifter® system: • General installation. It is recommended that the UniLifter® be installed with T-Gib mounted on top of the ejector plate • Angles. Designs using angles between 5° and 10° will typically yield the best results • Lifter guides. UniLifter® guides are recommended for designs with angles of 15° or whenever less than half the core blade is bored in the core insert • Guided ejection. It is recommended that guided ejection be used in all mold designs • Core blade venting. Vents should be ground in all permissible upper surfaces below the lower cavity surface. Venting should be added after establishing the ﬁnal core and guide dimensions. Vent dimensions for different types of polymers are provided in the cavity parting line or ejector pin ring groove venting detail drawings shown in Figures 10-123, 10-124, 10-127, 10-128, and 10-131. • Fit and ﬁnish. Recommended clearance for a core blade ranges between 0.001 and 0.0015 in where permissible. Additional performance can be obtained by applying a hard surface coating after the machining operation is completed (Ti N coating, chrome plating) • Locking angles. Locking angles may be designed in the mold if required to provide a locking surface to balance the cavity projected area force caused by the injection molding pressure. Figure 10-136 shows a mold venting application using the UniLifter® internal undercut ejection venting system.

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Mold closed Angle pin

Ring groove core pin vent

Figure 10-136 UniLifter® internal undercut ejection venting system (Courtesy: DME)

Single Angle Pin Slide Mold Venting When molded parts require side action to form different geometries, angle pin slides are commonly used for the designs of thermoplastic injection molds. Angle pins use the normal movement of the molding machine to remove the core pin as the mold opens at the parting line, using an angle pin and slide. The lock that is at an angle of 5° greater than the angle of the angle pin butts against the back surface of the slide. This positive locking device is used to ensure that the core pin is in the proper position and keeps the pin from retracting as the material is injected into the mold cavity. As the mold opens, the molded part stays with the movable half of the mold. The angle pin forces the slide to move toward the outer edge of the mold, pulling the core pin from the molded part. Once the core pin is free of the molded part, ejector pins push the molded part in the conventional manner. The slide is kept against the stop by spring tension. This keeps the hole in the slide in proper alignment for the angle pin when the mold is to be closed.

Stop

Ring groove vent ejector pins Mold open

Parting line cavity vents

Lock Slide

Molded part

Molds with slides offer several advantages for the installation of various types of vents. The slides are like additional parting line cavity surfaces where vents can be placed. Figure 10-137 shows a single angle pin slide mold with side core pin and ejector pin ring groove vents; the cavity vents at both parting lines and parting face cavity insert vents at the bottom walls of the molded part.

Insert parting face venting

Figure 10-137 Single angle pin slide mold venting system

Figure 10-138 shows a spool mold that requires two angle pin slides to form the end round walls, the parting line cavity vents are used between the slides. The main parting line is also vented around the front cavity ﬂange; the two interlock core pins are vented through the ring groove vent channels. Parting face venting is used between the slide base and the mold cavity plate surfaces.

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10.17 Mold Venting Systems

Mold closed Ejector sleeve

Angle pin

Ring groove core (interlock) vent

Mold open Slide

Spool

Parting line cavity vents

Stop Slide

Slide parting face venting

Ring groove core venting

Figure 10-138 Double angle pin slides mold venting system

Cold Runner Venting System The runner system begins at the machine nozzle tip radius with the mold sprue bushing, where the polymer melt front solidiﬁes forming a skin around the forward oriﬁce of the nozzle during the molding process. The quantity of solidiﬁed polymer melt front can be reduced by the proper design and temperature control of the nozzle tip. Operating the nozzle tip front inside oriﬁce as a thermo valve controls the sprue small diameter extension length inside the nozzle. When injecting the mold cavity at a high rate of speed, a runner system without vents can be responsible for unbalanced ﬁlling of the mold cavity causing burns and corrosion of the cavity metal. Since the volume of melt or shot size moving through the mold is in part dependent on the melt resistance to ﬂow, runners without vents will slow down the mold cavity ﬁlling rate.Venting the runners and sucker pin puller of a three-plate mold is as important as a two-plate mold. The ﬁrst cold slug pocket of the runner system is the sprue puller pin with ring groove vent channels. If an opening for the cold slug cannot be found at the sprue puller pocket, the runner system should be designed with additional safety

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10 Injection Mold Design protections. The sprue and all runners should have cold slug pockets, ejector pin ring groove vents, and runner vents. These items are used to catch the melt front cold slug and to discharge the gasses inside the mold runner system to the atmosphere. Without these items, the polymer melt front cold slugs may plug small runners, gates, or the cavities and runner vents. Venting the runner system is much more important than the sprue because all air left in the runner system must be exhausted through the cavity vents. Raising the cavity inserts between 0.003 and 0.005 in above the level of the mold base cavity plates is standard practice with some mold makers to obtain a better seal around the cavity and reduce ﬂash. A sprue puller is the least important area, because it is effective only in exhausting some air in the sprue. It is important for very long sprue bushings feeding small runners and low volume cavities. The following guidelines help avoid venting problems: Table 10-11 Runner Relief Edge Vent Depth for Common Types of Resins

• Install a generous pocket extension length between 0.25 and 0.50 in at the end of all runners.

Land depth “H” (in)

Resin

0.0015–0.0025

Acetal

0.0025–0.0035

ABS

0.0025–0.0035

Acrylic

• The mold runner parting line vents must not allow ﬂashing.

0.0010–0.0015

Nylon Unﬁlled

0.0015–0.0020

Nylon Glass R.

• The area of the runner vent (land length and land depth) must be large enough to prevent a rise in gas pressure in the mold cavity.

0.0015–0.0025

Nylon Mineral R.

0.0010–0.0015

LCP Glass R.

0.0020–0.0035

PC

0.0015–0.0025

PE

0.0015–0.0025

PP

0.0015–0.0025

PS

0.0015–0.0035

PPO

0.0015–0.0025

PSU

0.0015–0.0020

PET Glass R.

0.0015–0.0035

PVC Ridged

0.0015–0.0025

PVC Flexible

• Machine parting line vents from the end of each runner cold slug pocket to the outer edge of the mold. The size of these secondary runner cold slug pockets and vented ejector pins should be the same as the cross section area of the runner diameter.

• Install vented ejector pins in the runner channels and gate pullers. • Roughly grind the mold base cavity plates. • The land of the vents should be polished and have the same hardness as the cavities. In addition, the cavities and land of the vents could be chrome plated after the mold has been tested and approved for production. • The runner’s escape vent channel’s depth is used to convey the gasses outside to the atmosphere. The channel depth starts at the end of the runner vent land length to the external edge of the mold. Figure 10-139 shows the runner parting line venting design and Table 10-10 provides the values for the runner relief edge vent depth for common types of resins. Runner

Ring groove vent ejector pin

Parting line

Cold slug pocket L = 0.040 land H = Depth (see table 10-11) 0.022 escape vent channel depth 0.078 to 0.187 width

Secondary runner

Figure 10-139 Runner cold slug pocket parting line venting (in)

0.06 r.

Relief edge vent

Escape vent channel

10.17 Mold Venting Systems Figure 10-140 shows various types of ring groove vents for sprue puller pins, runner/gate ejector pin vents, and three-plate mold runner venting systems.

Figure 10-140 Additional cold runner system locations for venting

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10.17.8 Mold Venting Using Sintered Porous Insert Plugs Outside diameter Effective diameter

Pores

Effective length Full length

Figure 10-141 Sintered porous insert plug mold venting (Courtesy: DME)

Sintered vents are unique venting plugs composed of enough straight, parallel, and uniform pores made by a powdered metallurgy process. The pores allow trapped air or gas to escape from the mold cavity during the injection molding process, reducing the occurrence of defective parts. However, based on vent pore diameter and thermoplastic polymer melt characteristics, the pores usually ﬁll with mold deposits or other substances quickly, becoming inoperative and requiring maintenance. Sintered vent plugs should not be used in injection molding applications requiring mold cavity temperatures above 180 °F. Figure 10-141 shows a sintered porous self-contained standardized vent plug insert (DME) used for venting thermoplastic injection molds. The porous vent plugs are characterized by • Number of pores

• Pore diameter based on resin viscosity • Porosity % of effective diameter • Effective diameter • Effective length

Advantages of Sintered Vent Plugs • Venting the air from the mold or removing the trapped gasses by using sintered vent plugs reduces the occurrence of molding process problems such as short shots, ﬂashing, and burn marks of molded parts • Self-contained standardized vent plugs save engineering time in mold design, installation, and maintenance • A variety of standard size sintered vent plugs is commercially available • Fast and easy replacement or cleaning of sintered vent plugs from the production molds improves the process efﬁciency and productivity • Standard size sintered vent plugs have been tested in the ﬁeld to ensure product reliability 10.17.8.1 Types of Sintered Vent Plugs for Different Polymers Sintered vent plugs with a 0.03 mm vent pore diameter should be used with high melt ﬂow rate polymers, such as nylon, high melt ﬂow acetal, PBT, PET, polyethylene, or polypropylene. Sintered vent plugs with a 0.05 mm vent pore diameter are recommended for medium melt ﬂow rate polymers, such as polycarbonate, PPO, PS or ABS. When injection molding high molecular weight, high viscosity polymers (low melt ﬂow rates), such as HMWPE, high viscosity acetal, high impact resistance nylon, acrylic, or TPE, select the sintered vent plugs with 0.10 mm vent pore diameter. Stainless steel sintered vent plugs are recommended for polymers that have heat sensitive process characteristics, short melt residence time, and high processing temperatures. Melt degradation causes a chemical reaction discharging gaseous fumes. The melt sticking to the mold cavity walls is corrosive to the steels used for the construction of the molding machine and to the tool steels selected for the fabrication of the molds. Especially damaging are such polymers as PVC, acetal contaminated with

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10.17 Mold Venting Systems PVC, PC ﬂame retardant, or PTFE. Stainless steel sintered vent inserts are also recommended for ﬂame retardant thermoplastic compounds or thermoplastics using unstable thermo colorants. 10.17.8.2 Installation Procedures • The press-ﬁt is 0.01 to 0.02 mm for outside diameters of 10 mm or less, and 0.015 to 0.035 mm for outside diameters more than 10 mm. • Use a plastic or wooden hammer for installation. Do not tap the pore surface of the sintered vent plug with a metallic or hard tool. The use of hard tools will result in clogging or chipping of the sintered vent plugs. • Do not grind, machine, or cut the pores of the sintered vent plugs. Figure 10-142 shows a mold design application using a sintered vent plug to remove the trapped gasses from the mold cavity. Back escape vent channel

Vent plug insert

Sintered vent plug

Spiral baffler

Stripper plate

Gate runner P ipe

Coolant “OUTLET”

Coolant “INLET” Core

Spiral plug

Figure 10-142 Sintered vent plug insert mold venting application

10.17.9 Logic Seal (Negative Coolant Pressure) Mold Venting This type of venting makes use of the negative pressure found in the mold temperature controller (vacuum circulator) causing a vacuum (suction force) to remove the trapped gasses from the mold cavity. This mold venting system, also known as “Leak Stopper”, was developed and manufactured by Logic Seal. Inadequate mold venting poses a major molding problem when the gasses are trapped inside the cavities. The mold must eliminate the air trapped in the cavity after the mold closes, and also the gasses generated by the thermoplastic melt itself. As the melt enters the runner and cavity, the melt must displace these gasses pushing them ahead of the melt front. If these gasses are compressed they may exit through the mold parting line, cores, and inserts causing ﬂashing problems. Designing molds using negative water cooling pressure to vent the mold has several advantages: cycle time improvements, faster injection speed, lower melt/mold temperatures, better injection molding processing, and reduction of rejected molded parts.

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10 Injection Mold Design The introduction of new thermoplastic engineering resins has made mold venting more difﬁcult. As they are processed, these new compounds produce more gasses than the old resins, because they contain increased levels of ﬁllers, reinforcements, stabilizers, plasticizers, impact modiﬁers, pigments, and other additives, all of which can add to the amount of compressed trapped gasses that cause mold venting problems. 10.17.9.1 Logic Seal Venting Through Sintered Insert Plugs A solution for improving the mold venting problems is to increase the mold venting area and to control the mold temperature required by the thermoplastic polymer. The Logic Seal venting technique uses sintered vent plugs that allow the cavity’s trapped gasses to pass through the sintered vent plugs, but not the thermoplastic melt. The negative coolant pressure regulates the mold temperature and cools the sintered porous plug directly without causing coolant leakage inside the cavity. If the sintered metal inserts are not cooled, they overheat causing the sintered vent pores to ﬁll with mold deposits or other substances and quickly become inoperative. Figure 10-143 shows a typical mold application using sintered vent inserts for removing cavity gasses. Mold venting using negative coolant pressure and sintered vent inserts means a deep draw cavity can be center-gated at its base, allowing the top to retain a perfect ﬁnish. Several sintered metal inserts are pressed into the vertical wall of the core insert, forming the inside wall of the molded part. Gate Sintered vent plug

Gasses flow

Molded part

Sintered vent plug

Coolant Coolant suction gasses thru sintered plugs

Stripped plate Coolant flow

“O” Ring

Coolant “INLET”

Figure 10-143 Logic seal sintered plug mold venting application

These sintered metal insert vent plugs are connected with a discharge oriﬁce to the negative pressure waterline of the core, bringing coolant right up to the sintered vent plugs, always keeping them at the same or close to the coolant temperature. The sintered vents do not overheat or become plugged. 10.17.9.2 Logic Seal Cavity Parting Line Cores and Ejector Pin Venting Another mold venting method uses vented cores and ejector pins (with ring groove vent channels). These pins are mounted directly in the middle of the negative pressure waterline. This method has several advantages: because the

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Escape vent hole suction

Parting line cavity vent

Ring groove vent

Cavity Coolant “INLET”

Coolant suction vents gasses, cools cavity, cores & ejector pin

10.17 Mold Venting Systems

Gas flow Ejector pin escape vent gasses to suction line

Water flow Core ring groove vent gasses connected to water line

Figure 10-144 Logic seal cavity parting line, core and pin venting

vented pins run directly through the coolant, the pins will not overheat. Standard pins without direct contact with coolant generally have higher temperatures than the cavity surface temperature. The core and ejector pin ring groove vents will not become plugged with mold deposit or other substances reducing the mold venting performance. Another important advantage is that the tool makers are not constrained by the placement of pins, which can be located for maximum venting and ejection effectiveness. The cooling water negative pressure ﬂows around the pins providing for maximum cooling and the vacuum suctions to expel the trapped gasses from the runner system and mold cavity into the cooling water line. Tool makers, already familiar with the conventional technique of venting cores and ejector pins by machining ring groove vent channels to convey the gasses to the atmosphere, may feel more comfortable with this method before applying the mold venting technique through sintered vent plugs. Figure 10-144 shows a typical mold application using cavity parting line vents, vented cores, and ejector pins for removing the trapped gasses with negative coolant pressure through mold vent channels.

10.17.10 Mold Cavity Vacuum Venting System Mold vacuum venting systems are widely used for thermoplastic injection molds. This system eliminates the trapped gasses from the mold along with volatiles that may be generated by the thermoplastic melt during the molding process. This method has several advantages: a properly constructed and working system will allow the ﬁll of dead end sections of the mold cavity since there will be no gasses to impede the thermoplastic melt ﬂow. Poor weld lines and molded-in stress concentrations are also reduced as there is no excessive heat build up by the dieseling effect of compressed gasses and there is no need to wait for the gasses to escape from the mold cavities. The molds can be run at higher temperatures; as the thermoplastic melt is injected faster into the cavities, there is more frictional heat developed at the gate, decreasing melt viscosity, which also promotes a faster ﬁlling time. While the mold vacuum venting system will not show impressive gains on all injection molding applications, this technique has proven to facilitate faster injection cycles, to eliminate molding reject parts, and given better quality control.

694

10 Injection Mold Design The process is quite simple and inexpensive when molders have vacuum lines available to each molding machine. Portable vacuum systems are used when the unit is moved from machine to machine as required. A commercial mold vacuum venting unit is self-contained, with all controls and gauges mounted in a common panel for immediate installation and operation. Installation requires mounting the unit on the injection molding machine or on a separate stand for portability. It is recommended to connect the mold vacuum venting to the mold using a Teﬂon® steel braided hose when mold temperatures exceed 325 °F and to mount the two limit switches on the mold, platens, or rear rails of the injection molding machine that automatically begin the vacuum and blow back cycles. A vacuum gauge should be mounted as close to the mold as possible next to the molding machine to measure the vacuum. Leaks in ﬂexible hoses or connections may give a faulty picture of what is happening in the mold, if the gauge is on the vacuum reservoir. A ﬂexible hose connects to a trap mounted between the mold and the reservoir to keep the thermoplastic melt volatiles out of the vacuum pump oil. The vacuum reservoir is required because most vacuum pumps do not have the capacity to obtain the necessary pressure drop in the time it takes to ﬁll the mold. Mounting the sealing grooves and vacuum lines on the stationary half of the mold is advisable. The sequence of operation is to close the mold and when the reciprocating screw/check valve stops plastifying the resin and moving it forward for injection, to start the vacuum system by a limit switch or a manually operated valve. When the screw forward time (ﬁlling and packing the cavity) is completed, the vacuum is turned off. With injection, the system is activated when the screw starts forward. The vacuum pump runs constantly to maintain the vacuum in the reservoir. 10.17.10.1 Required Mold Modiﬁcations Minor mold modiﬁcations are required to use the mold vacuum venting system. A mold parting line venting installation consists of an “O” ring and a ring groove vent channel around the cavities. The “O” ring is required to seal the mold and prevent ambient air from entering the cavities, which, in turn, helps to create the vacuum. The gasket material should be Viton®, silicone rubber, or other materials that will withstand the mold temperature and pressure without melting or crushing. The gasket material should be soft enough to conform to the other half of the mold and provide a seal, but at the same time not hold the mold open. Between the sealing ring and the mold cavity is another ring groove channel tied into the bridge groove channels and the cavity relief edge vents. The most common method is to use a vacuum line, a ring groove vent channel, bridge groove channels, cavity relief edge vents, and “O” ring. This simply involves cutting a groove around the mold cavity face 0.105 in deep with a 0.135 to 0.140 in end mill to install a 0.125 in diameter “O” ring (the larger the diameter the better its performance) and a ring groove vent connecting the cavity relief edge vents. The ring groove vent channel should also be as deep and wide as possible. A 0.093 in wide by 0.022 in deep machining with a 0.125 in diameter end mill is recommended. To keep the “O” ring in place in a parting line venting application, it can be glued in the mold cavity insert by using RTV (room temperature vulcanizing) silicone. A Neoprene® “O” ring can withstand temperatures up to 160 °F; a Viton® “O” ring can withstand temperatures up to 420 °F. Figure 10-145 shows the mold vacuum venting cross section and top view of a mold lower half parting line cavity vacuum venting.

695

10.17 Mold Venting Systems

Mold lower half "top view" Ring groove vent

Cavity "O" ring Vacuum line

Moving mold half

Relief edge vent

Bridge groove

Figure 10-145 Mold vacuum parting line venting system

For a thermoplastic injection molded product with deep thin ribs, a second mold cavity insert is used to form the rib and to provide a contact area between the insert walls to machine desirable parting face cavity insert vents. This vent should be connected to the mold vacuum unit close to the bottom of the cavity insert. It is recommended that both vertical walls of the mold cavity insert be sealed by using Viton® “O” rings or be coated with a ﬁne coat of silicone grease or RTV silicone to prevent vacuum leakage. Figure 10-146 shows how the mold vacuum venting technique is accomplished by using the parting line cavity vents; the mold parting face cavity insert vents are connected to the mold vacuum venting unit. Another mold vacuum venting method is the use of a modiﬁed ejector pin machined with ring groove vent channels connected to the mold unit vacuum line. This mold vacuum venting technique requires that the vented pin be sealed at the support plate of the mold to prevent vacuum leakage. A Viton® “O” ring seal is recommended due to its durability and heat resistance.

696

10 Injection Mold Design Parting line cavity Vent “O” Ring

Cavity

Mold cavity plate

Escape vent hole Parting face insert vent Vacuum line

Insert

Core

“O” Ring

Figure 10-146 Parting face insert vacuum venting system application

Parting line cavity vent

“O” Ring

Escape vent hole Cavity

Mold cavity plate Lead-in clearance

Vacuum line

Ejector pin ring groove venting

Support plate “O” Ring

Figure 10-147 Mold vacuum venting a cavity blind hole by ejector pin

Figure 10-147 shows a mold cavity with a deep blind hole in the middle of the cavity. Venting this part is accomplished by having a vented ejector pin with ring groove vent channels connected to the mold vacuum venting unit. 10.17.10.2 Evacuation Time Analysis To ﬁgure out if the evacuation or vacuum cycle interferes with the injection molding cycle time, it is necessary to evaluate the following conditions: • Screw forward time required to ﬁll the mold cavity • Total shot weight, including mold cavities, runners, gates, and sprue • Speciﬁc gravity of the thermoplastic polymer. The analysis is based on evacuating 1.00 in3 of mold cavity volume and inlet air pressure of 60.00 psi at 5.5 cfm (maximum efﬁciency). The mold vacuum unit can draw 26.00 in of mercury in 0.131 s. These characteristics are obtained with a sealed mold by using a hose of 0.375 in inside diameter with a length not to exceed 60.00 in.

697

10.17 Mold Venting Systems

Example 10-9 Calculate the evacuation time for processing 33% ﬁber glass reinforced nylon based on the following conditions: : Screw forward time required to ﬁll mold cavities is 2.00 s : Total shot weight is 250.00 g

: Speciﬁc gravity of the resin is 1.38. Volume =

Shot Weight (g) 250 = = 11.00 in 3 (10-8) Specific Gravity × 16.38 1.38 × 16.38

To calculate the evacuation time, use the vacuum draw versus evacuation venting time graph in Figure 10-148. Consider drawing a vacuum of 24.00 in of mercury, the evacuation time is calculated by multiplying the total shot volume 11.00 in3 by the 0.039 s to draw a vacuum of 24.00 in of mercury (see Figure 10-135), it would take (11.00 × 0.039) = 0.43 s. If 26.00 in of mercury vacuum is required, the evacuation time would increase to (11.00 × 0.131) = 1.45 s. The evacuation time (0.43 and 1.45 s) at both levels of the vacuum is less than the screw forward time to ﬁll the cavities in 2.00 s; therefore, the mold vacuum venting unit will not interfere with or increase the molding cycle time.

26 24

21

15

12

9

0.008

Evacuation venting time, (seconds)

Figure 10-148 Vacuum draw vs. evacuation venting time

0.131

0.039

0.021

0.001 0.003

0.013

6

0.005

Mold vacuum draw, (inches of mercury)

18

698

10 Injection Mold Design

10.18

Mold Cavity Insert Contact Area Strength

The parting line of a mold cavity insert is subjected to the repeated impact loads during mold closing. This type of action can cause changes in dimensions, molding ejection problems, and ﬂashing at the parting line. To ﬁnd out the clamping force required for a mold that is going to be used in the injection molding process, multiply the total parting line projected area (including the runner and venting system) by a clamp force factor of 5.0 t/in2. To calculate the maximum clamping force of an injection molding machine based on the strength of the cavity insert contact area, the use of two cavity insert parameters is required: The cavity insert contact surface area, minus the mold venting areas at the parting line and the cavity insert strength to support the maximum clamping force without deﬂecting and/or compressing the cavity insert sidewalls. However, because of the operating conditions to which cavity inserts are subjected (repeated impact action, high clamping forces, high injection pressures, fast changes in temperatures, wear, corrosion, close tolerances, long service life), a high safety factor is used. It is recommended to use only 20% of the yield strength of the steel used for the construction of the cavity insert. In the following the strength of the cavity insert contact area is determined: • Determine the clamping force required for injection molding • Select only 20% of the yield strength for the cavity insert steel • Calculate the cavity and runner system projected contact area • Calculate the mold insert surface contact area at the parting line Example 10-10 Determine the clamping forces required based on the injection molding conditions and on the strength of the cavity insert contact area. Figure 10-149 shows the cavity insert dimensions. Assume the cavity insert is center-gated and is made of P-20 tool steel with a yield strength of 125,000 psi. The total venting area, projected cavity area, and the surface contact area of the cavity insert are calculated as follows: Total venting area

= 16 (cavities) × 0.125 (width) × 1.0 (length) = 2.0 in2

Projected cavity area

= 3.0 (width) × 6.0 (length) + 2.0 (vent) = 20.0 in2

Insert surface contact area

= 8.0 (length) × 5.0 (width) – 18.0 (cavity contact area) – 2.0 (vent) = 20.00 in2

Cavity venting 8.00 6.00

1.00

5.00 3.00

Cavity opening

Center gate

1.00

0.125 Insert surface contact area

Figure 10-149 Mold cavity insert, “top view” (in)

The clamping force required for holding the cavity and the maximum load recommended for the cavity insert contact area is calculated as follows: Clamp force (molding)

= 20.0 (cavity projected area) × 5.0 t/in2 = 100 t

699

10.18 Mold Cavity Insert Contact Area Strength

Work stress (steel)

= 125,000 (yield strength) × 0.20 (safety factor) = 25,000 psi

Clamp force (insert strength) = 25,000 (stress) × 20 (contact area) = 500,000 lbs. = 500,000/2,204 (lbs./t) = 226.86 t.

10.18.1 Cavity Insert Sidewall Strength Because the cavity insert is subjected to high internal pressures during the injection of the melt, calculation of the cavity inserts sidewall thickness is very important in keeping the cavity insert deﬂections below a speciﬁed maximum requirement. While the compression stress of the parting line contact surface area in the mold must remain within safe limits, it is the actual physical deﬂection of the cavity insert sidewalls of the mold that is very important in the operation of the mold. If these requirements are maintained within the speciﬁed limits, generally the stress values will also be satisfactory. The formulas used for calculating the cavity insert sidewalls are suitable for square or rectangular geometries only where the cavity insert length exceeds the cavity depth. As the depth approaches the length or exceeds it, its effect as a cantilever beam should be taken into account when calculating the total deﬂection. Insert

The following assumptions are made to calculate the strength around the cavity: Cavity

• An insert sidewall is a ﬁxed beam with a uniform distributed load. • An insert sidewall is a uniform internally loaded picture frame insert. • Each cavity insert sidewall is considered a freely supported beam with a uniform load applied to a rectangular mold cavity plate. • The effects of mold base steel strength, clamping force, the mold support plate thickness and the number of support pillars are ignored. • The applied load for combined sidewall thickness of cavity plate and insert is 20,000 psi, and for the cavity insert sidewall only 10,000 psi. Figure 10-150 shows two methods for mounting cavity inserts in the base mold cavity plates. One is the picture frame in which the opening is cut through the base mold plate; the other is the blind pocket in which a recess is machined to receive the insert cavity. The picture frame insert design is selected due to the high cost of accurate machining of a blind deep pocket insert.

Insert

Mold cavity plate

Picture frame insert

Insert

Cavity

The deﬂection should not allow the clearance between the components to increase to such a degree that ﬂashing can occur at the mold parting line. As the thermoplastic melt cools and shrinks, the cavity injection pressure is reduced and the sides of the mold cavity inserts are not loaded, therefore they return to their original straightness. If the cavity insert walls are bent, this causes ejection problems, the molded part becomes trapped between the steel insert cavity faces of the male and female halves of the mold. The greater the mold cavity insert deﬂection, the greater the difﬁculty in opening the halves of the mold.

Mold cavity plate Blind pocket insert

Figure 10-150 Mold cavity insert types and mounting

700

10 Injection Mold Design

10.18.2 Methods to Calculate the Strength of Cavity Insert Sidewall The ﬁrst method to calculate the strength of the cavity insert sidewall is based on the standard ﬁxed beam deﬂection equation: δ=

W × L4 384 × E × I

(10-28)

Where: δ (in) = Deﬂection, insert sidewalls W (lb.) = Load inside cavity = P (psi) Pressure of cavity × H (in) Depth of a cavity L (in) = Length of a cavity E (psi) = Modulus of elasticity (H-13 is 30 × 106 psi); I (in4) = Moment of inertia of cavity insert sidewall = (H × t3) / 12 H (in) = Depth of a cavity t (in) = Thickness of cavity insert sidewall. The necessary wall thickness (t) to permit a speciﬁed maximum deﬂection can therefore be obtained by transposing Eq. 10-28: 1/ 3

 W × L4  t =   32 × δ × H 

(10-29)

This calculation is ﬁrst made for the long side of the cavity insert. For the short sides, a similar calculation can be made, or the same wall thickness determined for the long side is used, as this wall thickness will give an even smaller deﬂection. Using Eq. 10-15, the minimum sidewall thickness (t) for a given deﬂection is obtained. In practice, however, the sidewalls of a mold cavity insert do not completely fulﬁll the basic requirements of end ﬁxing, on which the ﬁxed beam formula is based, and this method can lead to much higher deﬂections in service. If this formula is used, it should be restricted to molds in which the two mold halves ﬁt in each other so that the moving male half of the mold prevents the opening of the mold by acting as an interlock. The second method is to consider the sidewall of the insert cavity as a freely supported beam, according to the basic Eq. 10-30: δ=

5 × W × L4 384 × E × I

(10-30)

Then the sidewall thickness for a given deﬂection (δ) is calculated by using the following Eq. 10-31: 1/ 3

 5 × W × L4  t =   32 × δ × H 

(10-31)

Equation 10-31 is commonly used, giving the greatest sidewall thickness for a given deﬂection compared to previous methods. In this respect, it provides the

701

10.18 Mold Cavity Insert Contact Area Strength greatest margin of safety, and in many molds a greater sidewall thickness above that strictly necessary is often of little consequence. Perhaps, a more rational approach is offered by considering the mold cavity insert as a “picture frame” insert, considering each sidewall of the cavity insert independently as a uniformly loaded beam, with ﬁxed ends and freely supported ends respectively. However, this does not hold true, because the ends are neither ﬁxed nor free; nor does any insert sidewall deﬂect independently of the others, i.e., the effects of runner pressure drop and the corners must be taken into account. The third method considers each sidewall of the cavity insert as a freely supported uniformly loaded rectangular plate. To fulﬁll these conditions of support, the base of the cavity insert must be either solidly or rigidly registered into the side, while the open end of the insert cavity must be completely and ﬁrmly tied across by substantial registers in the moving half of the mold. Then, from a formula derived by Timoshenko, the approximate maximum deﬂections for the cavity insert sidewall are found by using Eq. 10-32: δ=

K × P × H4 E × t3

(10-32)

Where: K = a constant having the values as shown in Figure 10-151, for the ratio L / (H or F) L = longest length of the cavity insert sidewall H = depth of the cavity insert F = mold “B” cavity plate thickness P = injection pressure inside the cavity E = modulus of elasticity of the cavity insert t = cavity insert sidewall thickness.

“L” Length of cavity wall (inch)

“D” or “F” Depth of cavity wall (inch)

Ratio

6.00

5.00

4.00

3.00 2.66 2.25 2.00 1.80 1.50 1.21 0.94 0.70

0.02

0.04

0.06 0.08 “K” Constant

0.10

Figure 10-151 Cavity insert (length/depth) ratio vs. constant “K”

0.111 0.13 0.12 0.14

702

10 Injection Mold Design

Figure 10-152 Two cavity inserts, picture frame mounted in cavity plate

Figure 10-152 shows how to calculate the cavity insert sidewall and the combined sidewall thicknesses of the mold “B” cavity plate and the cavity insert.

Example 10-11 Two cavity inserts made of H-13 tool steel are mounted in line in two individual mold cavity plate openings made of P-20 tool steel as shown in Figure 10-142. Calculate the cavity insert sidewall thickness (t) that meets the requirements of a maximum deﬂection (δ) of 0.001 in, when 10,000 psi injection pressure (P) is applied inside the cavity. The cavity insert sidewall thickness (T) is calculated by transposing the Eq. 10-32 in a function of T: Where: K = Constant per Figure 10-141, where L / H (8.0 / 3.0 = 2.66) = 0.13 E (H-13 ﬂexural modulus) = 30 × 106 lb./in2 T =

3

K × P × H4 = E×δ

3

0.13 × 10,000 × 34 = 1.52 in 30 × 106 × 0.001

Example 10-12 Calculate the combined sidewall thickness (N) for the previous example for a maximum deﬂection (δ) of 0.0001 in, at 20,000 psi injection pressure (P). By using the previous equation, the sidewall thickness (N) can be calculated with: K = Constant per Figure 10-141, where L / F (8.0 / 4.0 = 2.00) = 0.111 E (P-20 ﬂexural modulus) = 25.56 × 106 lb./in2 N =

3

K × P × F4 = E×δ

3

0.111 × 20,000 × 44 = 6.06 in 25.56 × 106 × 0.0001

The cavity plate width (J) = (2 × N) + C = (2 × 6.06) + 3.50 = 15.62 in or standard mold base cavity plate width = 15.875 in

703

10.18 Mold Cavity Insert Contact Area Strength

Let us assume that a sectional mold is being built by cutting one large opening in the mold base cavity plate for mounting the two cavity inserts. An insert spacer is mounted in the middle opening between the two cavity insert internal sidewalls and a smaller runner injects in the middle of both cavities in the same direction, as shown in Figure 10-153. Calculate the combined cavity plate sidewall deﬂections, one at location (a) where 20,000 psi injection pressure is applied inside the cavity and maximum deﬂection (middle) of the mold base cavity plate. A beam load diagram representing the cavity plate’s large opening conditions is developed to ﬁnd the corresponding deﬂection equations used in this example.

T = 1.52 L (8.00)

S (Spacer) = 3.00 J (15.875)

Example 10-13

W

W

A

A V

T = 1.52

W

H = 6.06 F (4.00)

E (P-20 ﬂexural modulus) = 25.56 × 106 psi

Figure 10-153 One hole cavity plate, two inserts mounted in line

I (Moment of inertia) = (F × N3) / 12 = (4.00 × 6.063) / 12 = 74.18 in4 W = L × F × P = 8 × 4 × 20,000 = 640,000 lb. a = (L + t) / 2 = (8 + 1.52) / 2 = 4.76 in b = 2 × L + S + 3 × t = 2 × 8 + 3 + 3 × 1.52 = 23.56 in Deﬂection at “a”: W × a2 δa = (3 × b − 4 × a) 49 × E × I 640,000 × 4.762 = (3 × 23.56 − 4 × 4.76) 49 × 25.56 × 106 × 74.18 = 0.0080 in (80 times the recommended deflection). Deﬂection at the middle: W ×a (3 × b2 − 4 × a2 ) 196 × E × I 640,000 × 4.76 (3 × 23.562 − 4 × 4.762 ) = 196 × 25.56 × 106 × 74.18 = 0.0130 in

δM =

The deﬂection at the point of the load (a) is 80 times and at the middle 130 times greater than for the two openings for mounting the cavity inserts shown in Example 10-13. The maximum deﬂection of 0.0130 in of the mold cavity plate sidewall thickness would permit some thermoplastic melt to ﬂash, the mold parting line opening would become progressively worse, and associated problems would increase. Thermoplastic polymers that cannot ﬂow in such narrow spaces interfere with the mold opening, because an interference of 0.0080 in per side is a sizable amount for the ejecting force to overcome.

Beam load diagram V = 23.56 W (640.000 lb.)

A = 4.76

W (640.000 lb.)

A = 4.76

704

10 Injection Mold Design Cavity plate deflection mold parting line opens High load Beam bending diagram

Cavity insert

Runner Flashing

There have been cases where the deﬂection of the plate was large enough to cause a sizable increase in wall thickness of the molded products. When the cavity injection pressure is reduced, the mold cavity plate steel springs back to the original position with a force equal to the one that caused it to deﬂect. The pressure from deﬂected steel on the increased wall thickness of the molded part could be so great that mold opening is impossible. Only by completely disassembling the mold into its basic components while the mold is hanging in the molding machine would it be possible to separate both halves of the mold. After clearing out the thermoplastic melt, the mold would need to be reassembled and reestablished into operation. Such problems can be avoided by calculating the required cavity insert and mold cavity plate sidewall thicknesses and strength correctly.

Flashing

10.19

High pressure drop low force, not flashing High injection pressure deflection and flashing

Figure 10-154 Mold layout out of balance with parallel cavities High force deflect/flash

High force deflect/flash

Runner Flashing

Flashing

Cavity insert

Figure 10-155 Balanced mold layout, cavity rotated, long runners Cavity plate permissible deflection without flash Beam bending diagram Uniform and thicker insert sidewall Cavity insert

Film gate

Short runner

Width reduction

Low differential forces between both inserts inside sidewalls

Figure 10-156 Symmetrical balanced mold with short runner and ﬁlm gates

Mold Layout Case Studies

The following mold layout illustrations show three schemes for positioning the cavity inserts and the cold runner system in the mold cavity plate. The last two mold layouts were modiﬁed to reduce the out-of-balance forces that cause the parting line ﬂashing. Various other mold layout schemes (hot runner) can be set up to achieve the same objectives as the last mold layout illustration. Although some small out-of-balance force may be tolerated in practice, it is essential that the greatest care is exercised to make sure these forces do not become excessive, otherwise, ﬂashing can result. Correcting this problem can only be accomplished by changing the molding parameters, employing less than optimum melt and mold temperatures, and reducing injection speeds and pressures. Figure 10-154 shows an out-of-balance mold layout that causes ﬂashing problems at the right side of the mold. The mold is considerably out of balance because of the mold layout position of the maximum projected area portions of each long cavity and runner. The parts were molded of glass reinforced PET (fast melt ﬂow resin), using high process and mold temperatures, fast injection speeds, and the maximum clamping capacity to ﬁll the cavities. Figure 10-155 shows a mold layout improvement by rotating the lower cavity insert 180° to balance the forces at both sides of the mold. This modiﬁed mold layout requires a longer runner and small edge gates in the same areas as the previous lay out. It also requires high injection pressures for the long runners and restricting gates to ﬁll the thin and long wall cavity. Almost half the injection pressure is lost through the runners and gates. The remaining injection pressure is used to ﬁll the cavity. The highest injection pressure occurs at the entrance of the cavity where the projected area is the largest, causing forces greater than the clamping force available, pushing the moving half or opening the parting line of the mold, causing ﬂashing problems at each side of the cavities and runners. Figure 10-156 shows a mold layout scheme for symmetrical balancing and reduction of cavity pressure, using short runners, bigger gates (ﬁlm or fan type of gate) positioned at the same location in both cavities. This is the preferred mold layout to balancing the cavity melt ﬂow length for a uniform melt ﬂow of the cavities. The injection pressure conditions will be symmetrical and in balance for both cavities, while the overall size of the mold could be reduced allowing the smaller mold to use a lower capacity injection molding machine, thus reducing costs of the molded parts.

705

10.21 Tolerances for Thermoplastic Molded Parts The optimum mold lay out provides the following beneﬁts: • Optimum mold venting performance • Faster and uniform molding cycles • Reduction of injection pressure, melt, and mold temperatures • Reduction of cavity deﬂections without ﬂashing • Thinner cavity insert sidewalls • Reduction of overall mold size • Runner length, diameter, and pressure drop reductions • Larger gate size closer to sprue to reduce ﬂow path and ﬁlling the cavity uniformly using lower injection speed and pressure • Balanced melt ﬂow in the cavities (part dimensional control)

Cavity plate

LOAD

• Reduction of rejected molded parts

10.20

Mold Support Pillars

The mold cavity plate surface should be between 0.003 and 0.005 in lower than the cavity insert surface to provide a parting line contact surface area in both mold halves. Figure 10-157 shows the beneﬁts of support pillars; they greatly increase the capacity of the mold for supporting the projected area of the cavities and runner system. By providing additional stiffness to the support plate, they prevent deﬂection of the mold cavity inserts.

Support plate

Ejector housing No support pillars Cavity insert

LOAD

For example, a 11.875 × 15 in mold base with a maximum permissible load on the support plate without support pillars will allow 14 in2 of projected cavity area without ﬂashing (top illustration). The use of one row of support pillars (middle illustration) will quadruple the projected cavity area to 56 in2. Two rows of support pillars (bottom illustration) will provide nine times the projected area or 126 in2. Support pillar

10.21

Tolerances for Thermoplastic Molded Parts

Does my company need and can it afford the tolerances I have speciﬁed? A realistic view of the purchase cost of tolerances often helps avoid high manufacturing costs without affecting the performance of the thermoplastic injection molded products.

One row of support pillars increases the capacity to support projected cavity area four times Cavity

LOAD

It may be unreasonable economically to specify ﬁne production tolerances on a molded product when it is designed to operate within a wide range of environmental conditions. Dimensional changes of the product caused by temperature variations alone can be three to four times as great as the speciﬁed tolerances. Specifying ﬁne production tolerances does not make too much difference in many applications such as bearings and gears. Fine tolerances with thermoplastics are not as vital as with metals because of the resiliency of thermoplastic materials.

Two rows of support pillars increase the capacity to support projected cavity area nine times

The following are some general product design suggestions for the selection of production tolerances:

Figure 10-157 Increase in load capacity using mold support pillars (Courtesy: DME)

Support pillars

706

10 Injection Mold Design • The product design for a thermoplastic molded part should indicate the conditions under which the dimensions shown must be held. • In a product design drawing, the overall tolerances for a thermoplastic molded part should be shown in in/in, not in ﬁxed values. A title block should read,“All decimal dimensions ±0.00X in/in (±0.00X mm/mm), unless otherwise speciﬁed”, not ±0.00X in (±0.00X mm). • Only ﬁne tolerances should be speciﬁed in the design drawing.

• When compromises in tolerances could be acceptable from a performance standpoint, the tolerances in question should be discussed with the molding process engineer. Figure 10-158 shows the tolerances developed by the plastics molding industry for a thermoplastic injection molded part. The purpose of these speciﬁcations is to assist the product designer in obtaining a quick and preliminary analysis for the different molding tolerances found in thermoplastic injection molded parts. A

5.00

A, B, or C

Internal dimensions

ENGLISH UNITS Suggested tolerances (inches)

6.00

F

4.00

D E AR SE FIN TAND AR S CO

3.00 2.00

D

1.00 0.002 0.004 0.006 0.008 0.010 0.012

0.014 0.016

E

Tolerances ( inches)

Tolerances ± (inches) Fine Standard Coarse 0.002 0.004 0.006 0.003 0.005 0.007 0.002 0.003 0.004 0.002 0.003 0.004 0.001 0.002 0.003 0.002 0.004 0.006 0.004 0.005 0.007 0.006 0.007 0.008

For dimensions over 6.00 inch add (±) 0.001 (Fine), 0.002 (Standard) or 0.003 (Coarse) inch per added inch Single Cavity D = lower than 1.00 Multiple Cavity D = lower than 1.00 D greater than 1.00 add per inch For C = lower than 1.00 C greater than 1.00 add per inch E = lower than 0.10 E = 0.10 to 0.20 E = 0.20 to 0.30

External height D Sidewall thickness F = lower than 0.30 Bottom wall thickness E

A

125.00

A, B, or C

Internal dimensions

150.00

METRIC UNITS Suggested tolerances (millimeters)

C

B

F

100.00

D E AR SE FIN TAND AR S CO

75.00 50.00

0.10

0.15

0.20

0.25

0.30

0.35

0.40

Tolerances ( millimeters)

Sidewall thickness F = lower than 7.50 Bottom wall thickness E

D

25.00 0.05

External height D

C

B

For dimensions over 150.00 mm add (±) 0.02 (Fine), 0.05 (Standard) or 0.08 (Coarse) mm per added mm Single Cavity D = lower than 25.00 Multiple Cavity D = lower than 25.00 D greater than 25.00 add per mm For C = lower than 25.00 C greater than 25.00 add per mm E = lower than 2.50 E = 2.50 to 5.00 E = 5.00 to 7.50

Figure 10-158 Tolerances for thermoplastic injection molded parts

E

Tolerances ± (millimeters) Fine Standard Coarse 0.05 0.10 0.15 0.08 0.13 0.18 0.002 0.003 0.004 0.05 0.08 0.10 0.001 0.002 0.003 0.05 0.10 0.15 0.10 0.13 0.18 0.15 0.18 0.20

10.21 Tolerances for Thermoplastic Molded Parts

10.21.1 Factors Affecting Dimensional Control Tolerances Thermoplastic injection molded product ﬁnal tolerances are more complicated to predict or calculate. Tolerance is a result of product design geometries, product size, wall thickness, type of thermoplastic, resin molding characteristics, linear coefﬁcient of thermal expansion, mold/post molding shrinkage of resin, amount of ﬁber glass or mineral reinforcements (orientation), material handling, molding machine, auxiliary equipment, mold design, molding conditions, and end use environment conditions. The following variables are responsible for dimensional changes in thermoplastic injection molded product and the inability to hold dimensions: Thermoplastic resin • Thermoplastic melt viscosity variations require process temperatures adjustments. This can occur by deliberate or accidental mixing of lots of virgin material having different viscosity properties. • Virgin material moisture content variations affect the drying process conditions and cause changes in melt viscosity. • Types and levels of reinforcements, melt ﬂow rates, impact resistant additives, lubricants, and colorants may not be uniformly dispersed. • Variations in size and geometry of the virgin resin pellets. • Different amounts of reground material (sprues, runners, defective parts) of different sizes (longs and ﬁnes) with different moisture content that are poorly blended with the virgin resin. Material handling • Drying process affects the melt ﬂow rate and molding parameters. • Condensation of virgin resin (cold warehouse, warm molding area). • Resin contaminated with ﬁnes and longs or other materials. • Molded parts after ejection are subjected to different cooling methods: ejected parts fall in a box, parts placed facing up or down on the work table while others are immediately packed, etc. • Transporting the molded parts through hot, dry, wet or cold locations. • Fans or open windows causing drafts in the work area. • Considerable atmospheric changes in humidity and temperature. Injection molding machine • Erratic performance of the electrical and hydraulic systems. • Nozzle temperature problems; cold nozzle causes short shots; hot nozzle burns the melt changing its viscosity • Worn clamp platen bearings (mold moves causing concentricity and vertical dimension variations). • Worn plastifying screw, check valve, or barrel causing melt variations and back ﬂow leakage (short shots and poor density of molded parts). • Shot-to-shot variations in the quantity of material fed to the cylinder.

707

708

10 Injection Mold Design Mold • Gates too small and/or too long (premature melt freeze-off caused by a pulsing of the melt ﬂow rate within the cavities). • Runner system too long or too complex (excessive injection pressure drop causing melt ﬂow pulsations during ﬁlling the cavities). • Poor or improper mold venting. • Deﬂections of the core insert, cavity insert sidewall, or support plate. • Poor temperature control of the cores and cavities surfaces. • Worn mold components such as leader pins, ejection system, wedge blocks, binding, etc. Auxiliary Equipment • Mold temperature ﬂuctuations of the mold thermolator, or the size of the thermolator too small. • Mold temperature control equipment can not maintain a uniform temperature of the cavities at required levels. • Defective dryers. Defective desiccant beds position control, desiccant pellets do not absorb moisture (too old and/or contaminated), defective thermocouples, shut off valves, ﬁlters or blowers. Dryer temperature too low for drying the resin or variations in drying temperature and time exist. Processing dry air return too high for the efﬁciency of the desiccant material. Drying equipment needs an after-cooler on the processing air return to improve the drying efﬁciency of the desiccant beds, allowing the dryer to operate at higher drying temperatures (250 to 350 °F). • Inadequate chipper design or poor maintenance. The same chipper is used to grind different materials (contamination). Reground material of different sizes (longs and ﬁnes). • No vibrating screen equipment to separate the reground material in large sizes and ﬁnes. • Nonuniform metering, blending, and reground material ratio used with the virgin resin. Mold Operating Conditions • Molding cycles not uniform (change of shifts, delays between cycles). • Deliberate change in any controllable molding variable (melt/mold temperatures, screw back pressure, screw return velocity, injection pressure, packing pressure, screw forward time, injection speed, etc.). • Molding cycle too short, resulting in some unmelted material that could block the gates or enter inside the cavities (low back pressure, low compression screw, low barrel temperature proﬁle). • Injection molding ﬁlling rate too slow. • Mold surface cavity temperature variations.

10.22 General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins

10.22

General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins

Special consideration must be given to mold construction, mold quality, and mold performance. The mold vendor is responsible for calculating mold shrinkage, pressures, and mold stresses for each mold. All the mold components have to be sized and constructed accordingly. The mold vendor is responsible for any corrections that are required if the mold does not function properly.

10.22.1 Mold Design Requirements Metric and English units are used for dimensioning the thermoplastic injection molded product designs and details. All components used in the mold must be to the American National Standards in inch for molds that will be operating in the USA manufacturing facilities. For molds that will be operating in other countries, metric standards (millimeter) are required (metric to English conversions are not acceptable). Unacceptable Examples: ⇒ 10 mm ejector pin converted to 0.3937 in diameter. ⇒ 50 mm dimension converted to 1.9685 in. Unless otherwise speciﬁed, allowable tolerances on all the cavities, cores, slides and the related components are: ⇒ Two (2) place decimal dimensions = ±0.01 in. ⇒ Three (3) place decimal dimensions = ±0.001 in. ⇒ Four (4) place decimal dimensions = ±0.0005 in. All cavity and core pockets must be machined in and must be square and perpendicular within ±0.001 in. All cavity and core blocks must be square and perpendicular within ±0.001 in. The ﬁtting must be done by surface grinding and must ﬁt correctly in the cavity or core pockets. Hand grinding the cavities or core pockets or the cavities and core blocks is prohibited.

10.22.2 Mold Drawing Standards All drawings can be on roll size “Velum” (maximum of 52.00 in long) or on a computer ﬁle. The company border and title block drawing format is used. The ﬁrst sheet of each set of drawings is a “Table of Contents”, listing what is drawn on each sheet in the set of drawings. There must be at least one plan view for the ejector half of the cavity and one plan view for the ﬁxed half of the cavity. One view showing the ﬁxed half plan and half ejector plans are not acceptable. Label plan views with “Top” to show which side is the top of the mold when set in the injection molding machine. A minimum of two section views must be shown.

709

710

10 Injection Mold Design Assembly drawings will be ballooned, showing frame and cavity detail numbers. If a detail of the mold is drawn, the balloon will have a horizontal dividing line with the detail number above the sheet number, on which the detail is drawn below. On each sheet, below the title block, the detail numbers and description must appear in the margin. All components that are not standard should be detailed. Any standard component that requires modiﬁcation must also be detailed (e.g., ejector pins, core pins). The shut off height of mold must be at least 1.00 in larger than the minimum shut off height of the molding machine and at least 1.00 in shorter than the maximum daylight of the machine. For a family type mold, where parts are very similar, and a master cavity and frame are used to produce different parts, the company tool engineer can advise on what type of molds are required. The vendor will supply detail cavity changeover instruction sheets. Outlines of the molding machine tie bars need to be shown in phantom lines in the plan views of both the ﬁxed and moving halves of the mold. All water cooling bafﬂes and diverter plugs must be detailed. All water cooling bafﬂes and diverter plugs must be detailed. All runner systems (sprue, sprue puller, runner with cold slug pockets and ejectors, and gates) must be detailed. All venting systems must be detailed. A full set of approval drawings will be required for all assembly sheets; design detail drawings for some hot runner mold systems will be required at the completion date. The tool engineer will date and initial a full set of these drawings at this time. Documentation for Mold Modiﬁcations The tool drawing change letter has no relationship to the product change letter, only the change notice number does. The changed area of the drawing is identiﬁed with the change letter in a triangle. There may be changes not related to an engineering change, such as for mold improvements. These are recorded in the same manner. The company change notice block is required on every sheet, placed in the upper right hand corner of the sheet. Provide a minimum of eight lines of entry (tool engineer to provide format.) Mold Shrinkage Mold shrinkage values for each speciﬁc thermoplastic resin are the responsibility of the tool vendor to provide! The tool engineer or the resin supplier does not make shrinkage recommendations to the mold builder. Center-to-center distances between cored holes, slots, bosses, etc. are to be held to the part print mean dimensions plus the applied mold shrinkage rate for a particular dimension of the thermoplastic resin.

10.22 General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins Cored holes and openings in the part should be dimensioned on the tool drawing to the mean dimension, plus 50% of product print tolerances above mean, plus mold shrinkage rates. This will allow for metal removal on mold insert, if necessary. Product design having outside ﬁt conditions will be dimensioned to mean dimensions, minus 50% of product print tolerance below mean, plus the thermoplastic resin mold shrinkage factor. Molded Product and Mold Cavity Identiﬁcation Each cavity must be identiﬁed as required by the engineering product design print (tool engineer to advise). The company trademark is required on molded products and positioned as shown on the product design drawing. Cavity numbering sequence must be on the assembly drawing. In cases when there is more than one mold of the same tool number, the following mold and cavity identiﬁcation will be required: ⇒ ⇒ ⇒ ⇒

Mold # 1, Tool Number – “A” cavity 1 and 2. Mold # 2, Tool Number – “B” cavity 3 and 4. Mold # 3, Tool Number – “C” cavity 5 and 6. If cavity numbers go up to nine, underline cavities 6 and 9.

10.22.3 Required Types of Tool Steels for Mold Construction Types of Tool Steels Mold base custom designs, use P-20/P-21 prehardened 34–36 Rc. For all special structural components, use AISI 4140 prehardened 28–32 Rc (extension nozzle bushing, clamp plates). After the tool vendor has shown that the mold has been debugged and the preproduction results have been reviewed and approved by the tool engineer, the mold components can be plated. Electroless nickel plate all mold bases, clamping plates, locating rings, and ejector plates, at the completion of the ﬁrst molding trials, after making the necessary adjustments and corrections to the mold. Preharden and polish the cavities/cores/slides. DME, HASCO, NATIONAL mold components are to be used (e.g., core pins, return pins, extension nozzles, etc.). Types of tool steels, hardness and their applications are shown in Table 10-2 and 10-12 (the usage of other types of tool steels requires the approval of tool engineer): The tool engineer reserves the right to ask for formal certiﬁcation of any type of steel and hardness used in the mold by the mold vendor. Nitrating is required for all the cavities, cores, inserts, slides, and slide carriers. Typical case depths between 0.015 in and 0.025 in are recommended. Special cases must be reviewed with the tool engineer before nitrating. Before nitrating, all standing steel, all shut-off, or any special areas of the cavity or core must be digitized. The related components must be measured and documented on the

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712

10 Injection Mold Design Table 10-12 Types of Tool Steels for Molds

Material

Hardness

Application

H-13 / S7 / A2

Rc 48–52

Cavities/cores inserts

P-20 / P-21

Rc 28–36

Mold base

Nitrated H-13

Rc 65–74

Ejector/leader pins

D-2 / CPM 10V

Rc 60–62

Gate inserts

A2 / A6 / P6 /O2

Rc 48–52

Slides

Ampco 940

Rc 28–30

Inserts/cores in poor cooling areas

A2 / A6 / P2

Rc 56–58

Wear plates

SAE 1020

Rc 30–35

Ejector plate

tool drawings. (Note: Break all unnecessary sharp edges, 0.030 in × 45°, on any nitrate components to reduce chipping!). Welding No welding will be done without the prior approval from the tool engineer. Before any welding can be approved, the vendor must provide the proper procedures in writing to the tool engineer. Those procedures are to be in line with the type of steel to be welded and the details of the items, such as: • Surface preparation and cleaning • Preheat treatment or annealing • Preheating

• Post-weld surface treatment • Post-weld heat treatment • Post-weld cleaning • Filler material

Before welding, the vendor must submit an updated copy of the “Request to Weld” document to the tool engineer. Welding will not be allowed, except in those cases where the tool engineer has signed approval on the document before the welding operation. A copy of the signed document will be given to the tool engineer immediately following each approval or denial of the “Request to Weld” documents. The original of all “Request to Weld” documents used during the complete mold building, tryout, and “MQ1” will be supplied to the tool engineer with the tool drawings. Each original sheet is to be signed and dated by the vendor. This will be the vendor’s statement in writing regarding the weld history of the mold. Any welds done in the mold outside this speciﬁcation and/or not in concert with its requirements will result in replacement of the welded items at the vendor’s expense including any ﬁtting, tryout, debug, or re-qualiﬁcation costs incurred. Finishing Cavities, inserts, and cores must be ﬁnished as stated on the speciﬁcation sheet. Surfaces must be free of nicks, scratches, undercuts, and all ﬁnal polishing is to be done in the direction of the draw.

10.22 General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins Mechanically remove all EDM ﬁnished areas and thermally stress-relieve per material suppliers recommendations. Meta-Lax is not an acceptable method for stress relieving EDM residual internal stresses.

10.22.4 Mold Construction Requirements Locating Ring Locating rings need to be the nominal size required by the type of injection molding machines to be used. Locating rings are to be made of SAE 6521 or 6524 steels, hardened to Rc 28-32, and documented on the Q.C. Plate. Locating rings are not required for vertical clamp and horizontal injection type molds. Tool Engineer will provide molding machine information. Sprue Bushing The nozzle radius for sprue bushings should be 0.75 in or 0.50 in spherical (highly polished), made of SAE 6145 steel hardened Rc 43-45, ground and polished, unless otherwise speciﬁed by the tool engineer. The sprue length must be kept to a minimum (extension nozzle type of design is recommended). Circumstances requiring a longer sprue length must be made known to the tool engineer. All sprue bushings for horizontal clamp and horizontal injection type of molds, in which radial orientation is critical, must be keyed to keep sprue bushing from rotating. The sprue bushing small inside diameter should be between 0.031 in and 0.062 in larger than the front inside diameter (oriﬁce) of the injection molding machine nozzle. Vented Cold Runner System The cross section area of the runners must be fully round (modiﬁed trapezoidal type or other geometries must be approved by the tool engineer). Layout and size of runners are to be calculated for a speciﬁc mold layout and type of thermoplastic resin by joint effort between the tool vendor, tool engineer, and resin supplier’s technical representative. All cold runner layouts on the molds require a balanced melt channel distribution, unless otherwise approved by the tool engineer. A step down can be made in runner size diameters to obtain an efﬁcient pressure drop of the runner, and an ideal cavity pressure required by the thermoplastic melt to mold the products. All cold slug pockets required for the runner system should be vented at the sprue puller, at the ends of the main and the secondary runners. All runners should have vented ejector pins; they must be smooth, free of nicks or imperfections and polished in the direction of ﬂow. Runners along mating surfaces of two inserts are not acceptable. In such cases, a runner plate must be provided. Use of Mold Flow and Mold Shrinkage analysis software programs is highly recommended and encouraged whenever possible. Gates and Mold Venting Systems The types of gates and mold venting systems are determined by the type of mold ejection system, the part geometry, resin type, and dimensional control required by the application.

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10 Injection Mold Design A separate set of mold venting system detail drawings are required, showing the vent channels, geometries and dimensions for venting the cavity/core inserts, all runner’s cold slug pockets, sprue puller, and ejector pins. The venting layout should show venting channel intersections and end at the edge of mold or venting connected to side of mold, via a clear port designed for this function. “Ejector pin venting”, “sintered vents”, “logic seal venting system” or “vacuum venting system” are venting systems developed by several companies. These venting systems are commercially available and should be considered for mold design applications. All are highly recommended for deep/blind and thin (less than 0.020 in) walled products, difﬁcult to ﬁll applications to improve weld line strength, reduce ﬂashing problems, reduce injection pressure, and to improve surface ﬁnishing. Vented Sprue Puller Reversed taper or “Z” type sprue pullers with ring groove vent channels are recommended for horizontal clamp and horizontal injection, or for vertical clamp and vertical injection type molds. Sprue pullers are not required for vertical clamp and horizontal injection type molds. The sprue puller is replaced by a vertical ejector pin system to free the runner and the cavities from the lower cavity plate. Layout of Mold Cavity and Core Inserts A symmetrical mold cavity layout in the middle of the mold will provide uniform distribution and balanced clamp force to counteract the polymer injection pressures that may cause a mold cavity shift. Allow space between cavity inserts for runners, waterlines, return pins, ejection system, support pillars, etc. for mold construction strength. Each cavity insert should be mounted in a single opening “picture frame” of the cavity plates; two or more cavity inserts cannot be mounted in a single large opening, separated by spacers. The cavity plate sidewalls will have excessive large deﬂections causing serious molding problems. Lift holes are required for each cavity insert half. Holes will be sized correctly for the weight to be lifted. Each cavity and core insert requires two precision construction holes, holding the size and location to within 0.001 in. Hole sizes are determined by the size of the cavity insert blocks. Only standard sized holes are acceptable. The cavity inserts should be mounted at least 0.003 to 0.005 in above the mold base cavity plates. The cavity insert sidewalls should have a maximum deﬂection of 0.001 in, when the cavity is subjected to 10,000 psi injection pressure. The cavity plate compounded sidewalls should have a maximum deﬂection of 0.0001 in, when the cavity is subjected to 20,000 psi injection pressure to avoid deﬂections at the parting line causing ﬂashing and ejection problems. Cavity Insert Surface Contact Area (Shut Off) Cavity inserts contact surface areas minus the runner and venting areas should require less than 20% of the yield strength of the steel used for the cavity inserts. These values will provide the maximum clamping force or size of the injection molding machine for the mold that can be used. Picture frame type inserts are recommended for the construction of the mold. If the cavity inserts are offset, support pillars must be added to the mold base frame to balance the loading on the machine tie bars.

10.22 General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins Straight or severe angle vertical cavity insert contact areas are not acceptable. The tool engineer is to be notiﬁed of any deviations from the horizontal cavity insert contact area or face to face shut off. Draft Angles Usually, draft allowance should be per product print speciﬁcations. The responsible tool engineer must always conﬁrm the need for draft, or how much draft to be used in each speciﬁc part design. In cases where a draft angle has not been speciﬁed, or the required draft angle may be the cause of problems during the molding process, the tool engineer must be consulted. This may be the case, for example, if the speciﬁed draft angle interferes with the geometry of the part, the draft angle is too small to eject a textured wall, the core draft angle is too small (the molded part cannot be removed from the core, or the part inside diameter sticks to the core, producing blisters, rough surface and poor dimensions). Chamfer Edges Chamfers along parting surface break lines should be 0.125 in × 45°. Provide pry bar slots at main parting line of the mold. Speciﬁcations for Pipe Tapped Holes and Pipe Plugs All external and internal pipe threads are to be NPT type, also known as “National Pipe Tapered” or “American National Standard Taper Pipe Thread”. NPT threads have superior strength and sealing abilities. These attributes are due to the interference that occurs at all engaged roots and crests of the threads. This interference or sealing action will prevent spiral leakage and allow for pressure tight joints without totally relying on the use of sealing compounds. NPT style threads are tapered 0.75 in per foot and do require greater care and accuracy. When designing and manufacturing the mold it is necessary for external and internal threads to have full thread height for length of engagement. After the threads are assembled, the top of the NPT dry seal type pressure plug must be ﬂush with or below the top of the threaded hole. If a design or manufacturing compromise needs to be incorporated into the tool and there is not enough room for full length engagement of a standard NPT dryseal type pressure plug, a ﬂush type NPT pressure plug may be substituted. The “Dryseal pressure tight ﬂush” type plug has a taper of 0.875 in per foot and provides a high pressure seal through the differences in taper and is designed to remain ﬂush with the top of the 0.75 in taper NPT threaded hole. All holes to be threaded NPT style should be countersunk, drilled, and taper reamed. Incomplete or torn threads in an NPT threaded hole must not be accepted. Mold Temperature Control Cooling channels must be drilled from one side of the mold to the other. The plugs need to be easy to remove and the channels easily accessible for maintenance and cleaning. The mold cooling layout proposal must be approved by the tool engineer. The use of a mold cool temperature control analysis software program is recommended whenever applicable. All waterline openings other than outlets or inlets will be closed with brass pressure plugs, using Teﬂon® tape as a thread sealant. Plugs must be below the surface to eliminate any interference. Modifying the plugs is not acceptable. A separate drawing is required showing only the mold cooling systems in the cavity plates, cavity and core inserts.

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716

10 Injection Mold Design All the internal diverter passages must be plugged by slip ﬁtting 416 stainless steel or brass plugs, braised to a 0.125 in diameter, 416 stainless steel, or brass rods, that will hold the plugs in place. The rods will have a brazed threaded attachment at the pipe tap end of the hole. To ease plug removal, use “Neverseize” or an equivalent thread compound on plugs at assembly. Waterline connections need to be on the side of the mold and are to be tapped with a standard NPT thread to connect with the mold temperature control hoses’ male plug connectors. Special waterline ﬁttings of correct sizes should be used on the mold cooling channels. These ﬁttings do not restrict the coolant ﬂow rates. The mold builder is responsible for providing and installing all waterline connectors and pipes. These connectors should be compatible with those available at the mold tryout. Use only brass pipes and ﬁttings. (Common types of steel pipes and pipe ﬁttings of any type are not acceptable). Waterline entrances and exits are to be stamped on the mold base with sufﬁxes indicating “IN” and “OUT”. Mold drawings must also show the layout of cooling channels and same designations for “In”s and “Out”s. Permanently identify hose ﬁttings on both ends. Use Parker Hanniﬁn Teﬂon®/stainless steel braided hose # 919 with manufactured end ﬁttings also in stainless steel or equivalent. The size of the hoses will be determined by the mold cooling analysis and the turbulence ﬂow required by the application. Leader Pins and Bushings Leader pin diameters should be sized based on the leader pin length and mold weight. All leader pins must be placed in the side of the mold that will provide the least restriction to part removal. Leader pins must engage bushings by at least one pin diameter before cam pins engage and/or telescoping mold sections engage, or any other mold sections engage. Provide slug clearance behind leader pin bushings to clear out possible slugs from mold. One leader pin is to be offset to ensure proper assembly of the mold halves. Core Pins • Mounting: Core pin bodies must be held to the shortest practical length, and held in place by the head. Where radial orientation is critical, a positive lock at the head of the pin must be provided. Locking by means of pinning into a recess is not acceptable. Core pins must be mounted to provide access for replacement without major mold disassembly. If this is not practical, the problem must be made known to the tool engineer. Core pins mounted in assembly with sleeve ejectors need to be retained in the ejector bottom plate, backed up by a plate held by two screws. Do not use set screws to back up core pins. The bottom plate must be doweled to the ejector retainer block to establish and maintain core alignment. All core pin holes must be jig ground. The core pin outside diameter will be located both on the bottom of the core pin and next to the counterbore for the core pin head. Use letters for identifying core pins. Lettering of core pins must also be shown on the mold drawings. • Material: Use DME “CX” core pins (Rc. 50-55) and nitrated moving core pins.

10.22 General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins Vented Ejector Pins Location and size of vented ejector pins (ring groove vent channels) will be approved by the tool engineer. Contoured ejector pins must be approved by tool engineer. Contoured pins will be positively locked radially. Locking by means of pinning is not acceptable. All ejector pin holes must be jig ground in the cavity block. An ejector pin outside diameter will be located both on the bottom of the ejector pin and beside the counter bore for the ejector pin head. Use numbers for identiﬁcation and include on the mold drawings. Use “THX” Thru Hard DME ejector pins (Rc 65-74). Ejector Blades Ejector blades are used only in areas speciﬁed by the tool engineer. Ejector Sleeves Use DME or equivalent ejector sleeves. Return Pins Standard, hardened and nitrided, DME or equivalent return pins should be used. The minimum diameter of pins should be 0.625 in for molds designed for machines smaller than 550 t, 1.00 in diameter for larger molds; unless otherwise approved by the tool engineer. A minimum of four (4) pins per mold is required. Return pin contact surfaces will be ﬂat and perpendicular to mold travel, 0.062 in × 45° chamfer is required at the end of the pins. The maximum allowable span from pin to pin is 18.00 in. Ejector Guide Posts and Bushings All ejector plates are guided by a minimum of four hardened guide posts and bronze plated steel bushings. Guide posts are to be mounted in the cavity holder block and bushings in the ejector plates. Ejector Plate Return Ejector plates are to be spring returned. Accelerated or positive return mechanisms will be used in addition to the return spring when needed to prevent possible mold damage. Use a “Square D” Limit Switch (Class 9007 Type B 62G) mounted to sense that the ejector plate has returned. Support Pillars Support pillars should be placed as close as possible to the points of maximum loads, under the cavities, under the projected surface of runner and sprue systems, and under all slides. Support pillars are to be fastened by 3/8″-16 socket head screws to the clamp plate. Preload, if necessary, will be determined for each mold. Support pillars should be of adequate size and number to reduce movement of the mold during all phases of the molding cycle. Stop Buttons Standard “Stop” buttons should be used.

717

718

10 Injection Mold Design Mold General Assembly All loose components that are to be assembled with the mold must be screwed, doweled, or keyed non-symmetrically for error proof assembly. Wherever possible, cavity details are to be made removable from the mold without taking the mold out of the machine. Cavity and core inserts must extend from the parting line into the cavity pockets. Angled Pin Slides Mechanically actuated slides, properly heeled, are preferred to hydraulically actuated slides and should be used where possible. All sliding surfaces have a hardened steel-to-bronze or hard-on-hard surface contact and have provisions for lubrication. Heel surfaces will have hardened steel wear plates acting against bearing surfaces. Slides opening to the outside of the mold must have positive stop blocks to prevent accidental drop out from the mold. A positive system to keep the slides in the proper orientation when the mold is open must also be incorporated. DME or HASCO standard slide retainers or custom built spring plunger locks are to be used (tool engineer to advise). All sliding mold components must have a provision for lubrication. Use an angled lead-in on cam pins instead of a full radius. Hoist Rings and Hoist Ring Holes Safety is the number one priority when selecting and installing hoist rings onto the molds. Select hoist rings rated for a minimum of one and a half times the total weight of the mold. Hoist rings must be permanently mounted to the mold and should have means of securing them on the mold during the normal molding operation. The mold must balance within 5° in any direction when suspended on the hoist ring. If more than one hoist ring is installed on the mold, each hoist ring must be rated above the total weight of the mold. If a spacer block is required to elevate the hoist ring, the socket head cap screws securing the spacer block to the mold must have threaded engagement a minimum of twice the diameter and be sized so that each bolt can handle twice the total weight of the mold. Spacer block mounting bolts must be torqued to the recommended torque loads. Use DME hoist rings for molds up to 10,000 pounds and Danly hoist rings for molds between 10,000 and 20,000 pounds. Stamp in the location of the main lift ring (“Main lift ring location”), also stamp in the correct lift ring installation torque requirement (“Torque XXX foot-pounds”). Stamp tool number on the lift ring spacer block, if the mold is so equipped. Each mold will be provided with a minimum of two lift holes on each face of the cover and ejector half of the mold; i.e., top, operator side, bottom, and side opposite the operator. Table 10-13 Recommended Lift Holes and Hoist Ring Tapped Hole Sizes

Thread size

4-ft capacity

3/4″-10 UNC

500.00 to 2,000 lb

1″-8 UNC

2,000 to 6,000 lb

1-1/4″-7 UNC

6,000 to 12,000 lb

1-1/2″-6 UNC

12,000 to 18,000 lb

10.22 General Speciﬁcations for Mold Construction for Thermoplastic Injection Molding Resins Four (4) lift holes must be provided in each of the Q.C. clamp plates. Each lift hole must be rated for the total weight of the mold. Any modiﬁcations to the lift rings or the lift rings spacer block mounting bolts are strictly prohibited! Slides and sections that weigh more than 50.00 pounds must have lifter holes (drilled and tapped) for removal and installation purposes. Number and location are dependent upon speciﬁc application. Each mold half must be individually balanced for set-up purposes. In instances where putting lift holes on the centerline of gravity is physically impossible, three or more lift holes may be required for adequate balancing. Mold Halves Interlocks To ensure proper mold halves alignment, always use DME or HASCO interlocks. A minimum of four interlocks will be required per mold. Interlocks are to be positioned to ensure alignment of the mold halves, top to bottom and left to right. The tool engineer has to agree with the vendor, as to type, size, and location of the mold interlocks to be used. Mold Cavity Pressure Monitoring Button type pressure transducers are to be used. The tool engineer will select what type and how many pressure transducers to be used in the mold, if required. The connectors for the pressure transducers will be part of the system. Each connector will be identiﬁed by a number showing the transducer’s location in the mold. A bolt on the protective plate is required over the connectors. The controller will be mounted on the operator’s side of the mold. A metal plaque will be attached to the operator’s side of the mold showing the location of each pressure transducer. Mold Identiﬁcation Tag A metal mold identiﬁcation tag is to be completed and mounted on the operator side of the mold. All metal plaques and identiﬁcation tags are to be mounted on the mold using 1/4″-20 UNC screws. Cavity or Core Inserts Sidewall Thickness Wall thickness must be considered with extreme care to keep the cavity or core insert from deﬂecting, cracking, or breaking. The cavity insert wall thickness should be calculated based on the maximum length and depth of the cavity and the type of tool steel. The maximum sidewall deﬂection of 0.001 in is recommended for a cavity injection pressure of 10,000 psi. This cavity insert contact area should have a width large enough to hold the dimension requirements for the mold runner, gate, and venting system selected for this application. Water bafﬂe holes or any internal channels must have a full radius. Any cavity insert thin walls or a weak structure area deemed to be a potential problem must be reported to the tool engineer. Standard Mold Tool Steels All standard tool steel and hardness areas must be reviewed with the tool engineer. A decision must be made if the standard tool steels are acceptable or different types of materials (stainless steel, high thermal conductivity copper alloys, wear resistant, etc.) are required for the mold base, cavity/core inserts, or other components of the mold.

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10 Injection Mold Design

10.23

Mold Tryout – Debug – Approvals – “MQ1” Requirements

10.23.1 Mold Tryout or Evaluation The mold vendor delivers parts from the ﬁrst molding run for evaluation. To identify the severity of injection molding problems, ten samples are required of the molded component in its print speciﬁed material and ten samples in its natural color version. A complete layout inspection of those parts molded in print speciﬁed material will accompany the samples. The mold vendor will identify any known defects and show whether they will be addressed by mold modiﬁcation or requests for dispensation. Be advised that samples with missing or undersized ﬁllet radii are unacceptable. Dispensation for noncritical features and modiﬁcations to mating parts will be considered. The tool engineer will also consider vendor recommendations to improve the molded part’s functionality or molding performance of the parts.

10.23.2 Mold Debug Procedures The mold vendor is required to eliminate all dimensional, mechanical, electrical, hydraulic, and pneumatic errors, plus all water leaks and malfunctions in the tool. The mold vendor will develop the operating parameters with the tool engineer and modify the tooling as required to provide approved parts. Please note that some part drawings specify a moisture content and thermal conditioning before inspection. The mold prints must be revised to show the ﬁnal details of any modiﬁcation made to produce molded components. At the conclusion of this phase, all molds will operate entirely as speciﬁed and designed and make molded components to the speciﬁed dimensions. The mold vendor is responsible for running as many tests as are necessary to approve the mold.

10.23.3 Approval of Molded Parts and Pre-Production Molding Process Molded components supplied by the mold vendor upon completion of the debug phase are to be made in the mold under conditions closely approximating the intended molding production conditions for that mold. These molded components should not be supplied to the tool engineer until the mold vendor feels that they conform to the part drawing(s) or computer ﬁles. Support layout documentation must accompany these molded parts. The mold must be running at the speciﬁed cycle time and all molding parameters must be documented at this time. A 100% part layout of each cavity is required when the mold is running to the correct molding conditions. Inspection and documentation of all molded component print dimensions are required.

10.23.4 Mold Cavity and Core Surface Temperatures The mold vendor is required to provide a mold surface temperature map for both the cavity and the core for each half of the mold. Readings need to be taken when the mold is running in a molding production mode.

10.23 Mold Tryout – Debug – Approvals – “MQ1” Requirements

10.23.5 “MQ1” Requirements The “MQ1” production run is conducted after the mold vendor has veriﬁed that the mold has been properly debugged and the pre-production results have been reviewed and approved. The “MQ1” run is a continuous run for a six hour period, starting from the time when the mold and molding machine are production ready (wide adjustment range of the molding parameters). Ten subgroups of ten consecutive samples are to be pulled and labeled from each cavity at 30 minute intervals (the total number of samples from each cavity is 100 pieces). Material certiﬁcation, including moisture analysis, is required for “MQ1” before run. Measure all “KPC” dimensions with the same equipment being used at the company’s production facility. Provide 100% layout conﬁrming the molded components to print from two molded components, for each cavity. The “KPC” dimensions must have the mean centered with a “PPK” greater than 1.67 from all the above samples and be statistically shown. Any failures in the “MQ1” production run, or the sample parts dimensional correctness, will require a new “MQ1” production run. The tool engineer must be present for all mold tryouts, debug and the “MQ1” molding production run. The “MQ1” needs to be run on equipment that accurately reﬂects the intended production equipment. Parameters are those intended to be used in production. Any deviation will require written approval by the tool engineer. The mold supplier is responsible for conducting as many mold trials as necessary to bring the molded components up to an “MQ1” level. One additional mold tryout will be required after the steel hardening process of the mold is completed and the mold is reassembled for the ﬁnal time.

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11

Performance Testing of Thermoplastics

Thermoplastic materials cannot be subjected to the same tests developed for metals. They have their own special test requirements to comply with the demands of the plastics industry. While some standard testing methods may be applicable, in most cases, the results from the analytical tests developed for metal materials have very little value in deﬁning the properties and the process characteristics of the thermoplastic polymers. The tests commonly used in the plastics industry are established and used to describe many properties and the process characteristics of these polymers. This allows resin producers, product designers, molding processors, and end users to have a common understanding about what thermoplastic polymers are and what type of properties can be correlated with the end use results or representing the conditions found in the application. Tests are not ends in themselves, but rather a means of extracting knowledge about the thermoplastic polymers. The real test of a material comes with actual service. Once a thermoplastic end product is taken home and used by the consumer, it no longer matters whether the tensile strength is 3,000 or 28,500 psi. The ﬁnished product succeeds entirely or it fails. To assure success in automotive components and industrial products, the properties of the thermoplastic materials are studied by design engineers, who, through experience and judgment, balance material characteristics and service requirements against the geometries, sizes, functionality and performance needed in a product to give an adequate safety margin in the application. Reliable, useful tests for thermoplastics must be based on an understanding of the properties, processing characteristics, and the performance of these materials in similar applications. The following testing procedures are used for thermoplastic materials: • Mechanical, and creep analysis • Viscosity, chemical, and thermal analysis • Electrical, and ﬂammability analysis • Characterization of temperature resistance • Determination of chemical resistance • Characterization of weather aging and radiation resistance • Optical testing and characterization • Statistical analysis of test data These tests may also be quite different for the two major classes of plastics, thermosets and thermoplastics. The testing methods are used to characterize chemical composition, physical, and structural morphology and processibility. An important aspect of mechanical testing is the interpretation of results. Satisfactory mechanical characterization of an engineering thermoplastic requires not only specimens and test procedures of good quality, but also the proper interpretation of the results.

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11 Performance Testing of Thermoplastics Characterization of temperature resistance includes such diverse concerns as dimensional stability, effects on properties, and thermal decomposition. Electrical testing of thermoplastics deﬁnes the dielectric properties, arc resistance, tracking resistance, insulation, conductivity, and service end use temperature. Flammability classiﬁcation, ignition temperature, limited oxygen index, and smoke densities are some of the tests used to deﬁne the properties of engineering thermoplastics. Chemical resistance is a diverse area that requires many types of tests, most of which have been developed speciﬁcally for thermoplastics. Aspects of chemical resistance include the absorption and transport of solvents, acids and bases, the effects of many types of additives, and various types of degradation of the thermoplastics. Degradation can result from various sources, including chemical attack, thermal decomposition, photo oxidative deterioration and environmental damage. Weather exposure and radiation susceptibility are not strictly separate from chemical resistance and degradation. Different tests are used to measure the resistance of engineering thermoplastics to temperature, moisture, ultraviolet radiation, pollution, ozone, oxidation, and microbiological attack. Optical testing is used to characterize such factors as transparency, clarity, reﬂectance, hazes, color, gloss, and refraction. Many of the tests used in this area are identical to those used for other materials. However, some of the unique properties and applications of polymers have led to the design of special testing procedures. Statistical analysis of test data is probably the most important aspect of testing and characterization. For any type of thermoplastics testing, neither the tests nor the resulting characterizations are valid unless an appropriate statistical analysis test program has been established and the data is correctly interpreted.

11.1

Property Data Sheet for Thermoplastics

The American Society for Testing and Materials (ASTM) in cooperation with resin suppliers, product designers, and other institutions has developed the current ASTM test procedures employed for testing thermoplastic resins properties. These test results are used to produce the properties data sheet of thermoplastic resins and the marketing needs of competitive speciﬁcations. Product design engineers, who have experience with the polymer molding process characteristics and understand the complexity and signiﬁcance of the resin properties, are able to develop a special product geometry that satisﬁes the injection molding process and the end use service requirements. Selecting the ideal thermoplastic materials to be used for injection molding the new product requires taking into account all these factors. The properties data sheet shows the important properties of a thermoplastic material that have been developed or used for a speciﬁc application. This document is used in design studies, to compare properties with other materials, among others. However, it does not provide the information to understand the fundamental behavior, the viscoelasticity, or temperature sensitivity of thermoplastic polymeric materials.

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11.2 Tensile Testing (ASTM D-638) Thermoplastic injection molding resins are developed and commercialized for a speciﬁc market and application. The product data sheet has the properties tested in compliance with the ASTM test procedures. All the properties data sheets are different for each thermoplastic polymeric material. Remember that the resin is manufactured to comply with the speciﬁcations of the application, by providing the minimum number of tests needed for that market. When more properties of the resin are needed for a new application, the resin supplier’s representative can provide the additional data important for the design analysis of the new product that has not been published.

11.2

Tensile Testing (ASTM D-638)

Tensile strength and tensile modulus measurements are among the most important indications of strength in a material and are the most widely speciﬁed properties of thermoplastic materials. A tensile test is a measurement of the ability of a material to withstand forces that tend to pull it apart and to determine to what extent the material stretches before breaking. Tensile modulus is an indication of the relative stiffness of a material and can be determined from a stress-strain graph. Different types of thermoplastic materials are often compared based on tensile strength, elongation and tensile modulus data. Many thermoplastics are very sensitive to strain and environmental conditions. Therefore, the data obtained by this method cannot be considered valid for applications involving load time scales or environments widely different from the conditions used in the test. The tensile property data at several temperatures is more useful in comparing thermoplastic materials from a data base; they are used in actual ﬁne element computer analysis design of the product. This computer analysis program takes into account the time-dependent behavior of the product and the thermoplastic materials under simulated loads at various temperatures.

11.2.1

Tensile Testing Equipment

Commercially available tensile testing equipment can test a single station, or up to ﬁve independent stations, stretching the tensile test bar at a constant rate with a crosshead movement. It has a ﬁxed or essentially stationary member, carrying one grip and a movable member carrying a second grip. Self-aligning grips employed for holding the test specimen between the ﬁxed member and the movable members prevent alignment problems. A controlled velocity drive mechanism is used. Some of the commercially available machines use an independent closed loop servo-controlled drive mechanism to provide a high degree of speed accuracy. A load-indicating mechanism capable of indicating total tensile load with an accuracy of 1% of the indicated value or better is used. Figure 11-1 shows two commercially available tensile testing machines. Lately, the inclination is to use digital load indicators that are easier to read than analog indicators. An extension indicator, commonly known as the extensometer, is used to determine the distance between two designated points located within the gauge length of the test specimen as the specimen is stretched.

Standard tensile testing equipment

Automatic tensile testing of five specimens

Figure 11-1 Two types of tensile testing equipment

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11 Performance Testing of Thermoplastics 0.125 inches wall thickness 8.50 inches

0.75 inches

The advent of new microprocessor technology has virtually eliminated time consuming manual calculations. Stress, elongation, modulus, energy, and statistical calculations are performed automatically and presented on a visual display or hard copy printout at the end of the test.

0.50 inches

Figure 11-2 Test specimen ASTM D-638 for tensile tests

11.2.2

Tensile Test Specimen

Test specimens for tensile tests are prepared in many different ways. Most often, they are either injection molded or compression molded. Test specimen dimensions vary considerably depending on the requirements and are described in detail in the ASTM book of standards. Figure 11-2 shows an ASTM D-638 tensile test specimen most commonly used for testing rigid and semirigid thermoplastic materials.

11.2.3

Specimen Conditioning

The tensile test bars are conditioned using standard conditioning procedures. The tensile properties of some thermoplastics change rapidly with small changes in temperature and humidity. Procedure “A” for conditioning test specimens calls for the following periods in a standard laboratory atmosphere (50 ± 2% relative humidity, at 73° ± 1.8 °F). Specimens under 0.25 in thick are conditioned for 40 hours. Specimens over 0.25 in thick are conditioned for 88 hours. Adequate air circulation around all specimens must be provided. The temperature and moisture content of thermoplastics affect the mechanical and electrical properties. To get comparable test results at different times and in different laboratories, this standard has been established.

2 in.

11.2.4

Tensile Strength Test Procedures

The speed of testing is the relative rate of motion of the grips or test ﬁxtures during the test. There are basically four different testing speeds speciﬁed in the ASTM D-638 Standard. The most frequently employed speed of testing is 0.2 in per minute. The tensile test specimen is positioned vertically in the grips of the testing machine, the grips are tightened evenly, and ﬁrmly to prevent any slippage. Elongation measurement

Figure 11-3 shows the tensile test specimen grip, elongation, and the different types of ruptures. The speed of testing is set at the proper rate and the machine is started. As the tensile test specimen elongates, the resistance of the specimen increases and is detected by a load cell. This load value (force) is recorded by the instrument of the testing machine. Some machines can also record the maximum load obtained by the specimen that can be recalled after the completion of the test. The elongation of the specimen is continued until a rupture of the specimen is observed. Load value at break is also recorded.

Specimens after testing Figure 11-3 Tensile test specimens initial, elongation, and rupture

Figure 11-4 shows two typical tensile test output diagrams, one for oriented polypropylene, the other for polyethylene. Some materials, such as polymethacrylate, will break when they have been strained (stretched, elongated) only a small amount and while still showing

727

11.2 Tensile Testing (ASTM D-638)

The yield point is reached when the material continues to elongate (strain) with no increase in the strength of the pull (stress). When this takes place, the material yields, or in other words, is permanently deformed by elongation. The tensile stress at which yield takes place is called the yield stress.

Oriented polypropylene

Time

essentially elastic behavior. Other materials, such as nylon and polyethylene, can be stretched many times their original length before they break. The yield point and the corresponding yield stress are determined.

Load time curve Intercept

2 in. = 1% elongation Tangent Chart speed = 20 in./min. Crosshead speed = 0.5 in./min. Grip separation = 5 inches

Four results from a standard tensile test are generally reported: yield stress, tensile strength, ultimate strength, and percentage of ultimate elongation.

0

200

The tensile strength at yield and tensile strength at break (ultimate tensile strength) are calculated.

400

600

800

Load, (lb)

Tensile strength (psi)

Tensile strength at yield (psi)

=

Tensile strength break (psi)

Load at break (lb.) = Cross section area (in 2 )

Max. load (lb.) Cross section area (in 2 )

Time

Polyethylene

Force (lb.) = Cros section area (in 2 )

Load time curve Intercept

2 in. = 1% elongation

Tangent

Chart speed = 20 in./min. Crosshead speed = 0.5 in./min. Grip separation = 5 inches 0

50

100

150

200

Load, (lb)

11.2.5

Figure 11-4 Two examples of tensile test curves (elongation vs. load)

Tensile Modulus and Elongation

Tensile modulus and elongation values are derived from a stress-strain curve. An extensometer is attached to the test specimen as shown in Figure 11-5.

Stress, σ (psi)

The extensometer is a strain gauge type of device that magniﬁes the actual stretch of the specimen considerably. The simultaneous stress-strain curve is plotted as shown in Figure 11-6.

H

T

8.000 7.000

Tensile modulus

6.000

Yield point Break

5.000

B

Y

4.000 F P

2.000

Elongation at yield εY

1.000 A

Strain,

ε (inch/inch)

0.16

0.06 0.069

C 0.02

O

Elongation at break εB

0.12

2.835

Figure 11-6 Tensile stress-strain curve of a thermoplastic resin

Figure 11-5 Extensometer attached to the tensile test specimen

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11 Performance Testing of Thermoplastics The following procedure is used to make the calculations: • Mark off the units of stress in psi on the vertical axis of the chart. This is done by dividing the force by the cross section area of the test specimen. • Mark off the units of strain in in/in on the horizontal axis. These values are obtained by dividing the chart value by the magniﬁcation selected. • Carefully draw a tangent O-T to the initial straight line portion of the stressstrain curve. • Select any two convenient points on the tangent (points P and T). • Draw a straight line P-A and T-C connecting points P and T with vertical axis of the chart. • Draw a straight line P-F and T-H connecting points P and T with the horizontal axis of the chart. • Stress at H = 8,000 psi, corresponding strain at C = 0.069 in/in • Stress at F = 2,835 psi, corresponding strain at A = 0.02 in/in Tensile modulus

= Stress (H – F) / Strain (C – A) = (8,000–2,835) / (0.069–0.02) = 105,408 psi

Elongation at yield

= Strain at yield (εY) × Original length (2.0 in) = 0.06 × 2.00 = 0.12 in

Percentage elongation at yield = 0.12 × 100 = 12% Elongation at break

= Strain break (εB) × Original length (2.0 in) = 0.16 × 2.0 = 0.32 in

Percentage elongation at break = 0.32 × 100 = 32% For accuracy, the tensile modulus values should not be determined from the results of one stress-strain curve. Several tests should be conducted on the thermoplastic material and the average tensile modulus should be calculated. From a practical point of view, the determination of tensile modulus is subject to a number of errors associated with the stress-strain curves. They seldom have a truly straight line initial portion and considerable judgment must be exercized in drawing a tangent line through the stress-strain curve of thermoplastic polymers.

11.2.6

Molecular Orientation Effects

Molecular orientation affects the tensile strength values. A load applied parallel to the direction of molecular orientation may yield higher values than the load applied perpendicular to the orientation. The opposite is true for elongation. The process employed to prepare the specimens also has a signiﬁcant effect. For example, injection molded specimens generally yield higher strength values when molded under compression than when molded under tension. Another important factor affecting the test results is the number, location, and size of the gate on the molded specimens. This is especially true in the case of glass ﬁber reinforced tensile test specimens. A large gate located on top of the tensile specimen will orient the ﬁbers parallel to the applied load, yielding higher tensile strength. A gate located on one side of the tensile test specimen will disperse the ﬁber glass in different directions.

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11.2.7

Crosshead Speed Effects

As the strain rate or crosshead speed is increased, the tensile modulus of the polymer being tested also increases. However, the elongation rate is inversely proportional to the strain rate. The effect of the crosshead speed on the tensile modulus is shown in Figure 11-7.

11.2.8

Temperature Effects

Tensile modulus, (psi x 105)

11.2 Tensile Testing (ASTM D-638) 4.5

4.0

Acetal copolymer

3.5

0

0.1

0.2

0.3

0.4

0.5

0.6

Crosshead speed, (in/min)

The tensile strength and modulus properties of thermoplastic materials decrease proportionally to the temperature increase rate, while the elongation rate is increased as the temperature increases.

Figure 11-7 Crosshead speed effects on tensile modulus

The temperature effects on the tensile stress-strain at various temperatures curve of TFE thermoplastic polymer is shown in Figure 11-8.

0° F.

Stress, (psi)

The temperature effects on the mechanical properties of the thermoplastic materials cause severe changes, depending on the type of resin. Tensile testing of thermoplastic materials is performed at different temperatures. Figure 11-9 shows a commercial environmental test chamber used to study the effects of temperature on tensile, compression, and ﬂexural properties of the materials.

-100° F. 6.000

4.000

60

. °F

. °F 75 ° F. 100

2.000

20

400

0°

F.

° F.

0

11.2.9

Strain, (%)

Moisture Absorption Effects

If a nylon part absorbs moisture, dimensional changes will occur with increasing moisture content. To assume that dimensional increases will occur with increasing moisture content is unrealistic for most design purposes, as stress relief tends to counteract growth due to moisture absorption.When nylon products are exposed to higher humidity for long periods of time, the part dimensions will eventually increase, and this growth must be compensated for in product design.

Figure 11-8 TFE tensile stress-strain effects caused by temperature

Similarly, in dry applications, such as automotive engine parts, dimensional decreases due to stress relief must be considered. For typical nylon applications not exposed constantly to water, such as automotive body applications, an allowance of 0.005 to 0.007 in/in for possible growth due to moisture absorption has proven sufﬁcient. It is important to remember that moisture absorption and dehydration are slow processes, and the heavier the wall thickness the slower the process. The stressstrain graph in Figure 11-10 shows the equilibrium moisture content at various relative humidities for unreinforced nylon 6/6 resins. 13.000

Tensile stress, (psi)

11.000

Dry as molded

9.000 7.000

50 % Re 5.000

Figure 11-9 Tensile test apparatus with environmental chamber

y lat ive hu mi dit

10 0% Re lat ive

hu m id ity

3.000 1.000 0 0 0.2

0.6

1.0

1.4

1.8

Strain, (%)

2.2

2.6

3.0

Figure 11-10 Nylon 6/6 stress-strain effects caused by humidity

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11 Performance Testing of Thermoplastics

1 hr 10 h r s. 100 h r s. 1.00 0 hr s. 10 .00 0h r s. 10 0.0 00 hr s.

Another dimensional effect resulting from the environmental temperature at various relative humidity changes can be determined from the coefﬁcient of thermal expansion.

3.000

Tensile stress, (psi)

2.500

11.2.10 Stress-Strain Effects Caused by Creep The results of experiments in ﬂexural creep for long terms at 73 °F air temperatures and different levels of stress have been plotted and are shown in Figure 11-11. Data from this long term creep test shows a signiﬁcant reduction of tensile stress for acetal homopolymer at the beginning of the test compared to after 100,000 hours. Figure 11-11 shows the acetal homopolymer creep curves that may be used to predict creep or relaxation phenomena in acetal homopolymer parts under either tensile, compressive, or ﬂexural loadings by substitution of the stress-temperature-time dependent (creep or relaxation) modulus in the engineering equations.

2.000

1.500

1.000

500 0 0

0.5

1.0

1.5

2.0

2.5

Strain, (%) Figure 11-11 Acetal homopolymer tensile stress-strain effects of creep under load at 73 °F

It is suggested that the data from Figure 11-11 be used as required with acetal homopolymer data for unstressed exposure in air. This creep data could be used for product design strength analysis and prototype parts for intermediate stress-temperature-time conditions.

11.3

Flexural Testing (ASTM D-790)

The stress-strain behavior of polymers in ﬂexure offers another type of strength properties, closer to the type of loading, stiffness, and bending found in many applications. Flexural strength is the ability of the material to withstand bending forces applied perpendicular to its longitudinal axis. The stresses induced due to the ﬂexural load are a combination of compressive and tensile stresses. Flexural properties are reported and calculated using the maximum stress and strain that occur at the outside surface of the ﬂexural test specimen. For thermoplastic materials that break under ﬂexural load when using this test, the specimen is deﬂected until a rupture occurs in the outer ﬁbers. Many ﬂexible thermoplastic materials do not break, even after large deﬂections of the ﬂexural test specimen. The ﬂexural characteristics of these resins make the use of this test impractical. In such cases, the common practice is to report ﬂexural yield strength when the maximum strain in the outer ﬁber of the specimen has reached 5%. The ﬂexural strength test has several advantages over the tensile strength test. If the geometry of the application is a structural beam, and the thermoplastic component service failure occurs in bending mode, a ﬂexural test is more practical for design or speciﬁcation purposes than a tensile test. The tensile modulus value is generally about 10% higher than the ﬂexural modulus calculated from the outer ﬁber of a bent beam. The tensile specimen alignment is more difﬁcult to set up in the tensile test. Also, the tight clamping of the tensile test specimens creates stress concentration points. The small strain or deﬂection produced by the ﬂexural test is sufﬁcient enough to obtain accurate measurements. There are two basic methods used to determine the ﬂexural properties of thermoplastic materials. Procedure “A” is a three-point loading system utilizing

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11.3 Flexural Testing (ASTM D-790) P 2

P 2

Loading nose Specimen

T (thickness)

Loading nose

Loading nose R.

T

R. 2.00 inch support span

W (width) Specimen

L 3 Support span

Load span

L 3

L

6

5

4

3

2

1

0

1

4 5 2 3

Compression cell

Figure 11-13 Flexural test, four-point configuration – procedure “B”

center loading on a simple supported beam structure. A ﬂexural test specimen of a rectangular cross section rests on two supports, 2 in apart, and is loaded by means of a loading nose midway between the supports. The maximum axial ﬁber stresses occur under the loading nose. Flexural test conﬁguration Procedure “A” is shown in Figure 11-12. Procedure “A” is especially useful in determining ﬂexural modulus and yield stress properties required for product design, quality control, and for general speciﬁcation purposes. Procedure “B” is a four-point loading system utilizing two load points equally spaced from their adjacent support points with a distance between load points of one third of the support span. In procedure “B”, the ﬂexural test specimen rests on two supports and is loaded at two points (by means of two loading noses), each an equal distance from the adjacent support point. The ﬂexural test conﬁguration for procedure “B” is shown in Figure 11-13. Procedure “B” is used for testing materials that do not fail at the point of maximum stress under a three-point ﬂexural loading system. The maximum axial ﬁber stress occurs between the two loading noses. Procedure “A” is used principally for high ﬂexural modulus materials that break at comparatively small deﬂections. Procedure “B” is used for those ﬂexible materials that undergo large deﬂections during ﬂexural testing. The basic differences between these two procedures are the strain rates used for these tests. The strain rate for procedure “A” is 0.010 in/min and the strain rate for procedure “B” is 0.100 in/min.

11.3.1

Apparatus

The equipment used for tensile testing is also used for ﬂexural testing. The upper or lower portion of the movable crosshead can be used for ﬂexural testing. The dual purpose load cell that indicates the load applied in tension or in compression facilitates testing of the specimen in either tension or compression. The equipment used for this purpose should operate at a constant strain rate of crosshead motion over the entire range and the error in the load measuring system should not exceed ±1.00% of the maximum load expected to be measured.

Figure 11-12 Flexural test three-point configuration – procedure “A”

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11 Performance Testing of Thermoplastics The loading nose and support must have cylindrical surfaces. The radius of the nose and the nose support should be at least 0.125 in to avoid excessive indentation or failure due to stress concentration directly under the loading nose. A type of strain gauge called a deﬂectometer or compressometer is used to measure the deﬂection of the specimen.

11.3.2

Test Procedures and Equations

The test is initiated by applying the load to the specimen at the speciﬁed crosshead strain rate. The deﬂection is measured either by a strain gauge under the specimen with contact in the middle of the support span, or by measuring the deﬂection of the test specimen. A load-deﬂection curve is plotted to determine the ﬂexural modulus value. When a sample is bent, as in this test, forces of resistance (stress) are set up within the material, which tend to restore it to its original ﬂat position. Also, as a result of the bending, the material is distorted (strained). These stresses and strains are more easily understood from the ﬂexural stress procedure “A” shown in Figure 11-14. The curved line A-B represents the upper curved surface of the bent sample. The curved line E-F represents the lower surface or the outside curved surface and line is C-D an imaginary surface midway between both surfaces. The surface lines are of different lengths; the curved line A-B being the shortest and the curved line E-F the longest. During the bending process, the curved line C-D does not change in length. The outside line E-F becomes longer, while the surface is stretched and the inside curved line A-B becomes shorter as the surface is compressed. In fact, there are increasing degrees of stretching from the mid-line toward the outside of the bend and increasing degrees of compression from the mid-line to the inside. The compression on the inside surface is equivalent to the extension on the outside regarding both stress and strain. Flexural properties are calculated and reported for the maximum stress and strain that occur in the outside surface. The maximum stress (restoring force caused by bending) is related to the load, to the specimen dimensions, and to the deﬂection of the beam. The equations for calculating maximum ﬁber stress and maximum ﬁber strain for procedure “A” and for procedure “B” are shown below. Procedure “A”: σ A =

Compressive forces

P

Loading nose Tensile forces

Specimen

T

Center axis B D

A

C E

F L 2 L Span support

Figure 11-14 Three-point flexural test configuration, procedure “A”

Procedure “B”: σ B

3×W × L 6×δ×t εA = 2 × b × t2 L2

W ×L b × t2

Where: σ = Stress (psi) ε = Strain (in/in) δ = Deﬂection (in) W = Load (lb) L = Length of span (in) b = Width (in) t = Thickness (in) E = Modulus of elasticity (psi)

εB =

2×δ×t L2

δB =

δA =

0.25 × W × L3 e × t × w3

0.425 × W × L3 E × t × b3

11.4 Compressive Strength Testing (ASTM D-695) Flexural strength is equal to the maximum stress in the outer ﬁbers at the moment of break. This value can be calculated by using the above stress equations by setting the load value “W” equal to the load at the moment of break. For materials that do not break at the outer ﬁber of the specimen, strain values up to 5% are recommended; the ﬂexural yield strength is calculated using the same equation. The load value “W” in this case is the maximum load at which there is no longer an increase in load with an increase in deﬂection. The maximum strain in the outer ﬁbers, which also occurs at mid-span, is calculated using the previous equations.

11.3.3

Modulus of Elasticity

Flexural modulus is a measure of the stiffness during the ﬁrst or initial part of the bending process. The ﬂexural modulus should be called “Modulus of Elasticity in Bending”, but other names are also used such as modulus of elasticity, elastic modulus, or simply modulus. In addition, for practical purposes, the ﬂexural modulus, as determined by ASTM D-790, is generally lower in value than the tensile modulus determined by ASTM D-638. The part design should determine, which modulus (tensile or ﬂexural) should be used. The ﬂexural modulus is represented by the slope of the initial straight line portion of the stress-strain curve and is calculated by dividing the change in stress by the corresponding change in strain. The procedure for calculating ﬂexural modulus is similar to the tensile modulus calculations. In the determination of the ﬂexural modulus, several errors can occur. These are largely associated with the fact that stress-strain curves seldom have a truly straight line initial portion and considerable judgment must be used in deciding what line to draw through the data.

11.4

Compressive Strength Testing (ASTM D-695)

Compressive properties describe the behavior of a material when it is subjected to a compressive load at a relatively low and uniform rate of loading, which tends to crush or squeeze the specimen. By choosing proper conditions and by recording sufﬁcient data it is possible to determine the compressive strength and compressive stress with a single test. The compression properties of thermoplastic materials are very important for product design analysis. One of the basic product design rules is the application of compression structure loads, because plastics are stronger in compression than in tension. Molded thermoplastic products seldom deform from compressive loading alone, and the compressive loads are not always applied instantaneously. Compression tests provide properties for product design, data for research and development, quality control, and acceptance or rejection under speciﬁcations. Compressive properties include modulus of elasticity, yield stress, deformation beyond yield point, compressive strength, and compressive strain. However, compressive strength and compressive modulus are the two principal properties that are used in product design and resin speciﬁcation.

733

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11 Performance Testing of Thermoplastics In the case of a polymer that fails in compression by a shattering fracture, the compressive strength has a deﬁnite value. For those polymers that do not fail by a shattering fracture, the compressive strength is not a reliable test method to predict product failure during end use service. To determine the failure mode of the thermoplastic material a dynamic instrumented impact test should be used.

Movable head Specimen Fixed head

11.4.1

Compressive Testing Apparatus

The universal testing machine used for tensile and ﬂexural testing can also be used for testing compressive strength of various materials. A deﬂectometer or a compressometer is used to measure any change in distance between two ﬁxed points on the test specimen at any time during the test. Figure 11-15 shows a typical specimen set-up in compressive testing equipment. Movable head speed 0.05 in/min. Specimen

11.4.2 Direction of load

Fixed head

Figure 11-15 Special specimen holder for compressive testing

Test Specimens and Conditioning

Recommended specimens for this test are either rectangular blocks measuring 0.50 × 0.50 × 1.00 in or cylinders 0.50 in diameter by 1.00 in long. Specimens may be prepared by machining extruded rods or by compression molding thermoset compounds. However, the process of injection molding thermoplastic specimens with this geometry is very difﬁcult, because the product design and molding process principles are violated. The molded specimen will have inferior properties with stress concentration caused by gas and shrinkage voids, a low crystallinity rate, lower density, molded-in stresses caused by the gate, and degating that leaves pits, holes, and imperfections in the gate area full of stress concentration. The specimens are conditioned according to procedure “A” ASTM D-618.

11.4.3

Test Procedures

The test specimens are placed in a testing machine, such as a universal tensile tester equipped with a special specimen holder. It is the purpose of the holder to direct the force of the machine squarely onto the small ends of the specimen. The compressive test starts by lowering the movable crosshead at 0.05 in/min with the specimen being squeezed at that rate and the change in the length of the specimen (deformation, strain, in/in) is measured. The maximum load carried by the specimen during the test is recorded. The stress-strain data is also recorded either by recording the load at corresponding compressive strain or by plotting a complete load deformation curve with an automatic recording device. Compressive strength is calculated by dividing the maximum compressive load carried by the specimen during the test by the original minimum cross sectional area of the specimen. The result is expressed in (psi), either at the rupture of the specimen, or at a given percentage of deformation. Modulus of elasticity or compressive modulus, like tensile and ﬂexural modulus, is also represented by the slope of the initial straight line portion of the stress-strain curve and is calculated by dividing the change in stress by the corresponding change in strain. The method to calculate compressive modulus is the same as the tensile testing procedure. It is common practice to report compressive properties with the “compressive stress at 1.00% and 10.00% deformation”. It is believed that such values give a truer picture of the behavior of thermoplastic materials under compression.

735

11.5 Shear Strength Testing (ASTM D-732)

15

Tension

Stress, (x 103 psi)

10 Cross head speed 0.2 in/min

5 0

Compression 5

Cross head speed 0.05 in/min

10 15 10

8

6

4

2

0

2

4

6

8

10

Strain, (%) Figure 11-16 Acetal homopolymer stress-strain curve in tension and compression at 73 °F

11.4.4

Stress-Strain Tension and Compression Curves

In some design areas it is important to know the stress-strain relationship in tension and compression. In general, thermoplastics are stronger in compression. At high stress levels, the strain in compression is less than in tension. At low stress levels, the tensile and compressive stress-strain curves are similar. Consequently, at low strain the compressive modulus and stress is equal to that in tension. For relatively large strains, the compressive modulus and stress are higher than in tension. In Figure 11-16 stress-strain curves in compression are compared with stressstrain curves in tension.

11.5

Shear Strength Testing (ASTM D-732)

Shear strength of thermoplastic materials is deﬁned as the ability to withstand the maximum load required to shear the specimen so that the moving portion completely clears the stationary portion, which is similar to a paper punch with a female die and a close ﬁtting male punch. Shear strength testing is carried out by forcing a standardized punch at a speciﬁed rate through a sheet of thermoplastic, until the two portions of the specimen completely separate. The shear strength (psi) is determined by dividing the force required to shear the specimen by the area of the sheared edge.

11.5.1

Test Specimen and Apparatus

The specimen for the shear strength test can be either molded or cut from a sheet; it is either a disc with a 2.00 in diameter or 2.00 in2 plate. Such an arrangement is intended to prevent bending of the specimen. The thickness of the specimen may vary from 0.005 to 0.500 in. A hole with a 0.437 in diameter is drilled through the specimen at its center. A universal testing machine with a constant rate crosshead movement, such as the one used for tensile, compressive, and

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11 Performance Testing of Thermoplastics ﬂexural strength testing, can be used. A specially designed punch-type shear tool is used for shear testing. One such ﬁxture with a test specimen already mounted in place is shown in Figure 11-17. Considerable care should be exercised in the use of ASTM shear strength data for design purposes. It is believed that the resulting values are high because the tests use very thin specimens. Very thin samples may stretch down between the male and female portions of the punch rather than be broken off under true shear conditions. Apparatus

11.5.2

Test Procedures

The test is carried out by properly positioning the sample in the shear ﬁxture and pushing the punch onto the specimen, until a complete shear type failure is observed through the sample at a constant rate of speed (0.05 in/min). The maximum load (lb) required to shear (punch) the specimen is determined. The shear strength is calculated as follows:

Loading nose Male punch Clamp

Specimen Clamp Schematic shear testing

Figure 11-17 Shear strength test, apparatus, and schematic

Shear strength (psi) =

Force required to shear the specimen (lb.) Sheared edge area (in2 )

Where: Sheared edge area (in2) = Punch circumference × specimen thickness

11.5.3

Significance and Limitations

Shear strength data is of great importance to a designer of ﬁlm and sheet products, where the products will be subjected to shear loads. For the design of injection molded products, these types of shear loads are rarely considered a design factor. The shear strength data as reported in the material supplier’s literature should be used with extreme care when designing a product, because these shear strength values can be considerably higher than normal. For product design purposes, shear strength is generally considered to be essentially equal to one half the tensile strength, in accordance with common engineering practice.

11.6

Surface Hardness Testing

Hardness is deﬁned as the resistance of a material to deformation, particularly permanent deformation, indentation, or scratching. The hardness of a thermoplastic is related to its resistance to being indented or marked when a tool is pushed down against the surface. Thermoplastic materials vary a great deal regarding surface hardness. Therefore, one type of hardness test is not applicable for the whole range of hardness properties that are encountered. For softer materials, an instrument called a Durometer is used in conjunction with ASTM D-2240. For harder materials, a Rockwell hardness tester is used with procedure ASTM D-785-60T.

11.6 Surface Hardness Testing The Barcol hardness method is not as common as the previous techniques; it is covered by the ASTM D-2583 testing method. The Barcol hardness tester is generally used with reinforced polyesters. Hardness tests are characterized by requiring a specialized testing apparatus or a tool to carry out the tests. The test instrument consists of an indenting point or ball, a means of applying a known force to the indenter, and a scale for reading the penetration of the indenter into the material being tested. The results are obtained directly from a scale (dial); the hardness values do not have units. The hardness tests are not related to abrasion or wear resistance as for metals. Hardness is purely a relative term and should not be confused with wear and abrasion resistance of thermoplastic materials. Polystyrene, for example, has a high Rockwell hardness value but poor abrasion resistance. Hardness testing can distinguish the relative hardness of different grades of a particular thermoplastic. However, it is not valid to compare the hardness of various types of thermoplastics based on this test alone. Several other properties, such as molecular weight, ﬁllers, elastic recovery are also involved. The test is further complicated by creep.

11.6.1

Rockwell Hardness Testing (ASTM D-785-60T)

The Rockwell hardness test measures the net increase in depth impression as the load on an indenter is increased from a ﬁxed minor load to a major load and then returned to the minor load. The hardness numbers derived are unit-less. Rockwell hardness numbers are always quoted with a scale symbol representing the indenter size, load, and dial scale used. The hardness scales in order of increasing hardness are R, L, M, E, and K scales. The higher the number in each scale, the harder the material. There is a slight overlap of hardness scales and, therefore, it is quite possible to obtain two different dial readings on different scales for the same material. For a speciﬁc type of material, correlation in the overlapping regions is possible. However, due to differences in elasticity, creep, and shear characteristics between different thermoplastics, a general correlation cannot be made. 11.6.1.1

Test Apparatus and Specimen

Rockwell hardness is determined with an apparatus called the Rockwell hardness tester as shown in Figure 11-18. A standard specimen of 0.25 in minimum thickness is used. The specimen can either be molded or cut from a sheet. The test specimen also must have parallel and ﬂat surfaces. In addition, the test specimen must be free from sink marks, burrs, or other protrusions. The hardness measurement is carried out by using a heavy instrument called a Rockwell hardness tester. The test specimens are placed on a steel anvil and are brought into contact with a minor load (10 kg), which for testing thermoplastics is a steel ball. This forms the surface indented to “B”. The dial is adjusted to zero under the minor load and a major load is released (60 or 100 kg), which causes the ball to indent into the thermoplastic test specimen, forming the indented surface “D”. After ﬁfteen seconds, the major load is removed and a partial recovery from the indentation takes place. Then, after another 15 seconds (or a total of 30 seconds), the hardness is read on the gauge (dial) of the instrument with the minor load still applied while the surface recovers to “R”. The distance “R-B” shown in Figure 11-18 is used to calculate the Rockwell hardness.

737

738

11 Performance Testing of Thermoplastics Gauge

Gauge Major load

Major load

Gauge Major load

Minor load

Minor load

Minor load Ball

B

B Specimen

Minor load & ball applied surface “B” indented gauge set at zero

B D

D

R D

Major load removed leaving minor load indentation recovers to “R”

Major load added surface indented to “D”

Gauge

Specimen

Gauge

Specimen

Figure 11-18 Rockwell hardness apparatus and schematic sequence

Apparatus

For certain types of materials, particularly those having creep and recovery characteristics, the time periods during the applications of major and minor loads affect the results of the hardness measurements. For thermoplastics, the results are usually reported with R scale or M scale readings. The R scale means a 60.00 kg major load and a 0.50 in indenting ball are used. The M scale uses a 100.00 kg major load and a 0.25 in ball. The E scale, for thermosets, uses a 100.00 kg major load and 0.125 in ball. The Rockwell hardness number is directly related to the indentation hardness of the thermoplastic material, that is, the higher the reading, the harder the material. Readings are reported to be reproducible to ±2 scale (gauge) markings or units of measure for certain hard, homogeneous materials. Softer thermoplastic ﬁlled materials will give a wider range of variation in the surfaces (e.g., molded surfaces will give a higher reading than machined surfaces). At least ﬁve tests should be run and an average value reported. The ASTM test method recommends reporting readings between zero and 100, although readings to 120 are permissible.

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11.6 Surface Hardness Testing 11.6.1.2

Durometer Hardness Testing (ASTM D-2240)

The Durometer hardness test is mostly used for measuring the relative hardness of soft materials. The test method is based on the penetration of a speciﬁed indenter forced into the material under speciﬁed conditions. The Durometer hardness tester consists of a pressure device, an indenter, and an indicating device. The indenter is spring loaded and the point of the indenter protrudes through the hole in the base. The test specimens are at least 0.25 in thick and can be either molded or cut from a sheet. Several thin specimens may be stacked to form a 0.25 in thick specimen; however, one piece specimens are preferred, because the poor contact between the thin specimens may cause results to vary considerably. The test is carried out by ﬁrst placing a specimen on a hard, ﬂat surface. The pressure foot of the instrument is pressed onto the specimen, making sure that it is parallel to the surface of the specimen. The Durometer hardness is read within one second after the pressure device is in ﬁrm contact with the specimen.

Shore “A” durometer hardness tester

Two types of Durometers are most commonly used. Type Shore “A” and type Shore “D”. The basic difference between the two types is the shape and dimension of the indenter. The hardness numbers derived from either scale are unit-less numbers. Type Shore “A” Durometer is used with relatively soft materials, while type Shore “D” Durometer is used with slightly harder materials. Two commercially available Durometer hardness measuring instruments are shown in Figure 11-19. Measurements at scale readings more than 95 or below 5 are not recommended. When the reading registers more than 95 on the type Shore “A” instrument, the type Shore “D” Durometer should be used. For readings less than 5 on the type Shore “D”, the type Shore “A” Durometer should be used. When the readings on the type Shore “D” Durometer register more than 95, the Rockwell hardness tester “R” scale or harder must be used. At least ﬁve Durometer readings must be recorded and an average value reported for each hardness test performed.

11.6.2

Shore “D” durometer hardness tester

Figure 11-19 Shore “A” and Shore “D” Durometer hardness testers

Barcol Hardness Testing (ASTM D-2583)

The Barcol hardness test was devised mainly for measuring hardness of both reinforced and non-reinforced rigid thermoplastics. The tester is a portable instrument that can be carried around to measure hardness of fabricated parts as well as the test specimens. Barcol hardness testers consist of an indenter with a sharp, conical tip and an indicating device, such as a dial with 100 divisions. Each division represents a depth of 0.0003 in penetrations. The test specimens are required to be of a minimum thickness of 0.062 in. The test is carried out by placing the indenter onto the specimen and applying a uniform pressure against the instrument. The pressure is applied until the dial indication reaches maximum. The depth of penetration is automatically converted to a hardness reading in absolute Barcol numbers. When measuring the Barcol hardness of a reinforced thermoplastic material, the variation in hardness reading caused by the difference in hardness between unreinforced resin and reinforced thermoplastic materials should be taken into account. Generally, a larger sample size is recommended for reinforced thermoplastic specimens than the sample size used for unreinforced thermoplastics. A commercially available Barcol hardness tester is shown in Figure 11-20.

Figure 11-20 Barcol hardness tester

740

11 Performance Testing of Thermoplastics Engineering thermoplastic resins Thermoplastic elastomer resins

50 70

Elastomers (rubber)

100

120

140 150

Rockwell hardness "R" 40

55

70 80

95

Shore "D" durometer 20 30

50

70

90 95

Shore "A" durometer

Figure 11-21 Comparison between the three hardness scales

11.6.3

Factors Affecting the Test Results

The hardness of all thermoplastic materials is directly affected by changes in temperature and humidity. Specimens tested at higher than speciﬁed temperatures tend to indicate a lower hardness value. The surface ﬁnish also has signiﬁcant effect on hardness test results. A smooth molded surface yields a higher value than a machined surface. The hardness value may vary, depending on the type and amount of a reinforcement present in the resin. Correlation of hardness data from different apparatuses or operating conditions is discouraged by ASTM, because the correlation between these methods has produced misleading results. Regardless of this advice, Figure 11-21 shows the relationship between these three common hardness scales and the different families of materials using these hardness scales.

11.7

Abrasion Resistance Testing (ASTM D-1044)

Abrasion resistance of thermoplastic polymers is a complex subject. Many theories have been developed to support the claim that abrasion is closely related to frictional force, load, and area of contact. An increase in any one of the three parameters generally results in greater wear or abrasion. The hardness of the thermoplastic polymers affects the abrasion characteristics. For example, a harder material with a rough surface will wear through the surface of a softer material, creating grooves and scratches. The theory is further complicated by the fact that the abrasion process also creates oxidation on the surface from the build-up of localized high temperatures. The resistance to abrasion is also affected by other factors, such as molecular weight, resiliency, and type and amount of additives compounded in the thermoplastic material that inﬂuence the coefﬁcient of friction. This all makes abrasion a difﬁcult mechanical property to deﬁne as well as to measure adequately. Resistance to abrasion is deﬁned as the ability of a material to withstand mechanical forces (such as rubbing, wearing, scraping, or erosion) that progressively remove material from its surface.

11.7 Abrasion Resistance Testing (ASTM D-1044) Resistance to abrasion is signiﬁcantly affected by factors, such as wear test conditions, type of wear (abrasion, adhesion, corrosion, and erosion), and development and dissipation of heat during the test cycle. Many different types of abrasion measuring equipment have been developed. However, the correlation between the abrasion test results obtained from different machines and the actual wear of the thermoplastic polymer test samples is not reliable. Abrasion tests provide a relative comparison among the tested materials when the abrasion tests are performed under a speciﬁed set of conditions.

11.7.1

Taber Abrasion Testing

A material’s ability to resist abrasion is most often measured by its loss in weight when abraded with an abraser. The most widely accepted abraser in industry is called the “Taber Abraser”, which is in accord with the ASTM D-1044 method. A variety of wheels with varying degrees of abrasiveness are available. When the wheel with the grade of abrasiveness designated as CS-17 is loaded with 1,000 g during the abrasion test, it produces satisfactory results with most of the harder thermoplastic materials. For softer thermoplastic polymers, a less abrasive wheel under reduced loads is used. The test specimen is a 4.00 in diameter disc or a 4.00 in2 plate with both surfaces straight and parallel. A 0.25 in diameter hole is drilled in the center of the test specimen. The specimens are conditioned according to Procedure “A” of ASTM D-618 before testing. To commence testing, the test specimen is placed on a revolving turntable. Suitable abrading wheels are placed on the specimen under certain set dead weight loads. The turntable is started and an automatic counter records the number of revolutions. Most tests are carried out to at least 5,000 revolutions. The specimens are weighed to the nearest milligram. The test results are reported as weight loss in milligrams per 1,000 cycles. The grade of abrasive wheel along with an amount of load at which the test was carried out is always reported along with the results. When preparing thermoplastic materials for bearing applications, it is necessary to pay attention to the following factors: • Remove abrasive powder as promptly as possible • Minimize the generation of friction (select low coefficient of friction materials, design wall thicknesses as thin as possible to dissipate heat quickly, use adequate clearances, increase the precision of bores, and provide lubrication) • Increase the hardness of the metal surfaces as much as possible to reduce the surface roughness

11.7.2

Theoretical Analysis of Wear

As for sliding wear with thermoplastic materials, the following equation is proposed: Wear = f × W × v × t Where: Wear = Accumulated wear v = Speed

741

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11 Performance Testing of Thermoplastics

Load Abrader

Rubber wheel

Test specimen

Detail of abrading apparatus Load weight Vacuum nozzle Refacer switch

Specimen holder

Taber abraser tester

Figure 11-22 Taber abraser tester equipment

W T f

= Load = Time = Coefﬁcient of friction

ASTM D-1044 measures the resistance of transparent thermoplastic materials to abrasion. This method estimates the abrasion resistance of transparent thermoplastic materials by measurement of its optical effects. The test is carried out in similar manner to that described above, except that 100 cycles with a 500 g load is normally used. A photoelectric photometer is used to measure light scattered by an abraded track. The percentage of the transmitted light that is diffused by the abraded specimen is reported as a test result. Another method of studying the resistance of thermoplastic materials to abrasion is by measuring the volume loss when a ﬂat specimen is subjected to abrasion with loose abrasive or bonded abrasive on cloth or paper. This method is designated as ASTM D-1242. Figure 11-22 shows a detailed drawing of the Taber abraser testing method and commercially available abrasion testing equipment.

11.8

Coefficient of Friction (ASTM D-1894)

The coefﬁcient of friction is a very important parameter for designing dynamic mechanical components. The coefﬁcient of friction is determined by the sliding method and is a function of the molecular structure of the thermoplastic polymer and the type and amount of self-lubricating additives used to compound the resins. These additives migrate to the surface, lubricating it, and making it more slippery. Because this blooming action may not always be uniform on all areas of the surface, values from these tests may be limited in reproducibility. The low coefﬁcient of friction of semi-crystalline thermoplastic materials also depends on the mating surface, which is also affected by the sliding velocity and temperature. Typical mechanical components such as levers, rolling wheels, axles, wedges, gears, etc. have been used in many engineering thermoplastic applications. The service life of these components is dependent on the friction and wear characteristics of the materials used. Self-lubricating semi-crystalline polymers have the following characteristics: • Self lubricating (natural lubricity)

• Vibration absorbing (damping properties), used in applications where ﬂuctuating loads are applied • High mechanical resistance properties • Endure high end use temperatures

• Hard surface and scratch resistance • Light weight

• Good chemical resistance There are two types of coefﬁcients of friction: The static coefﬁcient of friction is the resistance force on a stationary object, which moves from a standing state to a moving state, and the dynamic coefﬁcient of friction, which is the resistance force during movement.

743

11.8 Coefﬁcient of Friction (ASTM D-1894) Placing an object on a sloped surface (B), as shown in Figure 11-23, with the slope angle increasing gradually, static friction resistance develops when the object starts sliding. At a given time, the ratio between the load applied perpendicular to the sloped surface (W) and the friction resistance (F) is called the coefﬁcient of friction (f) and is determined according to the following equation:

W B

F

A F

φ

F f = = tan φ W The coefﬁcient of friction of two similar thermoplastic resins sliding against each other is signiﬁcantly higher than the coefﬁcient of friction of thermoplastic material sliding against a hard, smooth, polished surface such as steel. Dynamic friction resistance occurs when a material is sliding or in rotary motion, presenting conspicuous variations due to the hardness, the surface roughness of the mating material, the pressure between mating materials, the friction rate, the presence and type of lubricants between mating surfaces, the temperature, contamination, and moisture.

11.8.1

Coefficient of Friction of Thermoplastic Materials

Thermoplastic and elastomeric materials have elastic characteristics, so that they are subject to elastic contact in many cases. Their coefﬁcient of friction is dependent on the shape and the size of the mating surface, the operating speed, the applied load, and the end use temperature. They are also inﬂuenced by lubricating conditions or surface contamination. Among solid materials, PTFE has the lowest coefﬁcient of friction. Semi-crystalline thermoplastics exhibit a smaller coefﬁcient of friction than amorphous thermoplastics. The coefﬁcient of friction of typical general purpose thermoplastics is shown in Table 11-1. Table 11-1 Coefficient of Friction for Selected Materials

Materials

Coefﬁcient of friction Static

Dynamic

Acetal homopolymer / steel

0.23

0.35

Acetal homopolymer (CL) / steel

0.15

0.20

Acetal homopolymer (TFE) / acetal Homopolymer

0.19

0.14

Acetal homopolymer (TFE) / steel

0.07

0.14

Nylon 6/6 / nylon 6/6

0.20

0.26

Nylon 6/6 / steel

0.16

0.21

Nylon 6/6 (TFE)/ steel Glass reinforced nylon 6/6 (TFE) / steel

0.10 0.19

0.18 0.26

Glass reinforced PSU (TFE) / steel

0.12

0.10

Glass reinforced PET / steel

0.17

0.20

Glass reinforced PET / glass reinforced PET

0.24

0.27

Vespel® (Graphite) / steel

0.18

0.20

Vespel® (graphite and TFE) / steel

0.12

0.15

Teﬂon® / steel

0.04

0.05

Figure 11-23 Static coefficient of friction testing example

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11 Performance Testing of Thermoplastics

11.8.3

Effects of Lubricants

A small amount of lubrication applied at assembly timewill remarkably enhance the friction and wear properties of a polymer. Semi-crystalline polymers can frequently be used in operation without lubricants (self lubricated). Excessively alkaline materials or extremely high pressure additives for lubricating metal components are detrimental to thermoplastics. To avoid such problems it is necessary to select the recommended lubricants appropriate for the thermoplastic material. 11.8.3.1

Self Lubricated Additives

Amorphous and semi-crystalline polymers use several low coefﬁcient of friction additives compounded with the matrix to reduce the coefﬁcient of friction and enhance the wear resistance characteristics of the ﬁnal resin. The following additives are used to modify the coefﬁcient of friction properties of polymers: molybdenum disulﬁde, graphite, PTFE powder and ﬁber, polyimide ﬁber, carbon ﬁber, silicone, oil, aluminum stearate, calcium stearate, zinc stearate, wax, and other types of lubricants.

11.9

Mold Shrinkage Test (ASTM D-955)

Injection molded parts made from a thermoplastic material are smaller than the cavity in which they were molded. The reason for the difference in size between the cavity and the molded part is that the mold cavity is quickly ﬁlled with the expanded melt (lower density), using high injection pressure and high temperature. When the polymer melt is cooled-off in the mold cavity, the thermoplastic molecules are crystallized into a smaller solid object of higher density. The difference between the linear dimension of the mold cavity and the corresponding linear dimension of the molded part is called the mold shrinkage. The shrinkage is measured per unit length of a cavity, after the molded part becomes dimensionally stable at room temperature. The mold shrinkage of injection molded thermoplastic materials is inﬂuenced by part thickness, gate dimensions, mold temperature, screw forward time, melt temperature, and injection pressure. For a full understanding of these variables, it is necessary to understand how and why injection molded thermoplastic materials shrink.

11.9.1

Purpose of the Mold Shrinkage Test

This method is intended to measure batch-to-batch uniformity of initial mold shrinkage of thermoplastic materials under speciﬁed conditions. This method does not provide for the measurement of mold shrinkage that may occur as the injection molded part materials age after the ﬁrst 48 hours out of the mold.

11.9 Mold Shrinkage Test (ASTM D-955)

11.9.2

Factors Affecting Mold Shrinkage

Mold shrinkage refers to a change in a linear dimension, never a change in volume. Consider a cavity with a particular dimension “C” (in) from which parts with a corresponding dimension “F” (in) are produced. The difference between C and F is the net shrinkage. The more useful relative value obtained by dividing the net shrinkage by the cavity dimension, i.e., (C – F) / C, is the mold shrinkage. For every inch of mold cavity dimension, the molded part will be smaller by this shrinkage amount. A semi-crystalline thermoplastic part lacking uniform wall thickness and symmetry will suffer molding problems, such as part distortion or warpage, voids, and irregular dimensions because of different resin shrinkages. It is more important to correctly estimate the absolute shrinkage for isolated dimensions within a part than to correctly estimate the relative shrinkage for all dimensions. There are many variables that affect mold shrinkage. Injection pressure, screw forward time, mold temperature, type of gate, location and size, wall thickness, plastifying screw, uniform cycles, types of polymer, melt and mold temperatures, and material handling are among them. While it is not practical to consider all these mold shrinkage variables in sizing a new mold cavity, it is of value to have an appreciation of the more important variables in the molding process.

11.9.3

Injection Molding Effects on Shrinkage

The difference between the dimensions of the mold cavity and of the injection molded product molded from a given thermoplastic material may vary according to the part design and operation of the mold. It may vary with the type and size of molding machine, the wall thickness of the injection molded parts, the degree of ﬂow or movement of polymer melt in the mold, the temperature of the melt, the size of the sprue, runner, gate, and nozzle, the molding cycle, the temperature of the mold, the screw forward time, including the packing pressure maintained through the check valve, and the uniform cushion size. Mold shrinkages are low when the design and the processing conditions are ideal for the polymer melt injected into the mold cavity; the molded product is hardened to a maximum weight while still under packing pressure as a result of the use of a melt cushion, gate, runner, sprue, and nozzle of proper size. Mold shrinkages may be much higher when the melt must ﬂow in the mold cavity, but does not receive and transmit enough melt and pressure to pack ﬁrmly into all the recesses of the mold cavity. The amount of crystallinity of the thermoplastic materials affects the mold shrinkage indirectly; semi-crystalline materials have higher shrinkage rates than amorphous materials. Semi-crystalline materials require a shorter molding cycle, higher melt, and mold temperature.

11.9.4

Requirements for Sampling

Injected molded samples that represent the properties of the material, depending upon its uniformity, shall be kept at room temperature in moisture proof, airtight vacuum-sealed bags until such time as required for mold shrinkage testing.

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11 Performance Testing of Thermoplastics

tion irec ed ers nsv Tra

5.00 in.

Edge gate

Flow direction 3.00 in. Rectangular test specimen flow orientation

Transverse direction

4.00 in. Dia.

Edge gate

Flow direction

Test Specimens To determine the ﬂow and transverse direction of the mold shrinkage, two types of specimens are used. The ﬁrst is an injection molded rectangular plaque having cavity dimensions of 3.00 in × 5.00 in and a wall thickness of either 0.125 in or 0.062 in. The second specimen is a circular plate, 4.00 in in diameter with a wall thickness of 0.125 in. The mold should have a single edge gate for the circular specimen. For the rectangular plaque, a single edge gate located at one end parallel to the 5.00 in length and in the middle of the 3.00 in width is needed. The single edge gate should have a width of 0.25 in by a land depth of 50% of the specimen wall thickness and land length of 0.040 in. Figure 11-24 shows mold shrinkage for the rectangular and the circular plaques, both specimens having a single edge gate. Injection Molding Machine A suitable injection molding machine with a micro-processor control and with proper peripheral equipment suitable for molding the desired polymers is required. The injection molding machine size should range from 1/2 to 3/4 of its rated melt shot capacity for the type of thermoplastic polymer being tested.

Circular specimen flow orientation

Figure 11-24 Mold shrinkage effects caused by flow direction

Conditioning Condition the test specimens at 73° ± 3.6 °F and 50 ± 5% relative humidity for no less than 48 hours in airtight, moisture proof, vacuum sealed bags before testing, according to Procedure “A” of ASTM D-618. In cases of disagreement, the tolerances shall be ±1.8 °F and ±2% relative humidity. Test Conditions Conduct tests in the standard laboratory atmosphere of 73° ± 3.6 °F and 50 ± 5% relative humidity, unless otherwise speciﬁed in the test method. In cases of disagreement, tolerances shall be ±1.8 °F and ±2% relative humidity.

11.9.5

Test Procedures

Measure the length and width of the mold specimen cavity to the nearest 0.001 in. Conduct the measurements at standard laboratory temperature as deﬁned in Section 3.1 of ASTM D-618. Mold at least three good test specimens using the thermoplastic resins to be tested under molding conditions recommended by the resin manufacturer. If no process recommendations are available, the following are suggested as suitable molding procedures for various thermoplastics. Injection Molding Conditions The injection molding machine used should be operated under an injection pressure between 10,000 and 15,000 psi, without exceeding 1/2 to 3/4 of its rated shot capacity for the thermoplastic material being tested. The nozzle and barrel temperature zones, screw back pressure, and screw rotation should be maintained at safe operating conditions (halfway). Under these conditions, low melt temperature causes short shots or premature freeze-off of the molded specimen. High melt temperature causes splay, ﬂashing, discoloration, drooling, dull surfaces, and blisters or bubbles on the molded specimen. The mold temperature should be maintained within ±5 °F of the prescribed temperature,

11.9 Mold Shrinkage Test (ASTM D-955) which depends on the material being molded and ordinarily will range between 120° and 200 °F. Injection Pressure Higher pressures result in less mold shrinkage. Increasing the injection pressure from 10,000 psi to 15,000 psi generally reduces the mold shrinkage, depending on the type of resin. At least two factors are controlled by the injection pressure. First, higher and faster injection pressures result in a lower speciﬁc volume, because the melt is slightly compressed reducing the mold shrinkage. Secondly, a longer and more efﬁcient packing pressure increases the part weight, reducing the mold shrinkage. Mold Temperature Mold temperature affects the mold shrinkage depending on the type of resin. Increasing the mold temperature increases the mold shrinkage. This may be attributed to two primary factors: ﬁrst, the speciﬁc volume of the melt increases with temperature. By increasing the mold temperature, the melt inside the runner, gate, and cavity does not freeze-off, allowing a longer screw forward time. Second, by increasing the mold temperature, the molecular (orientation) crystallinity and consequently, density of the molded part will be increased. Perhaps more important than mold temperature is the uniform cooling capacity of the mold cavity for obtaining uniform shrinkage and the elimination of heat spots that cause dimensional control problems, deformation, poor surface quality, parts sticking in the cavity, and longer molding cycles. Special attention should be given to mold cooling design. Gate Type, Size, and Location The gate is one of the more important variables affecting the mold shrinkage. Proper speciﬁcations for the type of gate, size, and location permit more rapid ﬁlling of the cavity, more efﬁcient screw forward time, and a smaller pressure drop between the cavity and the runner to obtain minimum mold shrinkage. However, a large gate could cause longer gate freeze-off time, because it is a function of the gate land length and depth. A gate with small land length and depth causes high mold shrinkage, because the small gate freezes off quickly, without packing the cavity. Gates with long lands also freeze-off quickly, causing excessive pressure drops and high mold shrinkage. Wall Thickness The part wall thickness is also an important variable affecting mold shrinkage. The thicker the wall, the higher the mold shrinkage. Thicker walled parts require a higher mold temperature for a longer time to reduce mold shrinkage. The melt skin next to the mold surface cools rapidly to a relatively low density solid (amorphous). These layers serve as a thermal insulator for the internal melt, causing it to cool slowly to a higher density solid (crystalline). Thicker walled molded parts experience internal voids, low impact strength, dimensional control problems, long molding cycles, and higher mold shrinkage. Post Molding Testing Allow the specimens to cool to approximately room temperature and then store in the standard laboratory atmosphere prescribed in Section 3.1 of ASTM D-618 before being measured. The period of storage for initial mold shrinkage

747

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11 Performance Testing of Thermoplastics for specimens with a wall thickness of 0.125 in will range from 1.0 to 2.0 hours. Measure the length and the width of each specimen at standard room temperature to the nearest 0.001 in and then return the specimens to storage in the standard laboratory atmosphere. Measure the specimens again no less than 16 and no more than 24 hours after molding to obtain the “24 hour shrinkage”, and again no less than 40 and no more than 48 hours after molding, to determine the “48 hours” or “normal” mold shrinkage. Mold Shrinkage Calculations Calculate the mold shrinkage (in/in) by subtracting the mold cavity dimensions in the melt ﬂow direction and the transverse direction from the corresponding specimen dimensions in both directions. The ﬂow differential dimension is divided by ﬂow cavity dimension to obtain the mold shrinkage in the ﬂow direction. The same procedure is used to calculate the mold shrinkage in the transverse direction by applying the transverse direction dimensions. For injection molding, the temperatures of the nozzle, barrel, and mold, the size of the nozzle, runner, and gate, the injection molding pressure, the molding cycle, the make, type, and the size of the injection molding machine used should be included in the test report. The initial 24 hour and 48 hour mold shrinkages in both directions expressed in in/in or percentages, each representing the mean of values obtained on three or more specimens, should also be included. Precision of the Test The probable limit of precision of this test is ±0.0005 in/in. With this limitation in precision, it is possible for samples to appear to be the same in shrinkage, yet in fact differ by 0.001 in/in, or for samples to appear to differ by 0.001 in/in and actually be identical.

11.10

Specific Gravity Testing (ASTM D-792)

Speciﬁc gravity is the ratio of the weight of a given volume of material at 73 °F (23 °C) to that of an equal volume of water at the same temperature. It is properly expressed as “speciﬁc gravity, 23/23 °C”. Density is the weight per unit volume of material at 23 °C and is expressed as D 23 °C in g/cm3. Speciﬁc gravity values represent the main advantage of thermoplastics over other materials, namely, less weight. All thermoplastics are sold today on a cost per pound basis and not on a cost per unit volume basis. Speciﬁc gravity is a strong element in the price factor and thus has great importance. Beyond the price and volume relationship, however, speciﬁc gravity is used in production control, both in raw material production and injection molding. The discrepancy is caused by the fact that water at 23 °C has a density slightly less than one. To convert density to speciﬁc gravity, the following formula can be used: Speciﬁc gravity (23/23 °C) =

Density, D 23 °C (g/cm3) 0.99756

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11.10 Speciﬁc Gravity Testing (ASTM D-792)

11.10.1 Test Procedures Two basic methods have been developed to determine specific gravity of thermoplastics, depending on the form of thermoplastic material. Method “A” is used for a specimen in the form of sheet, rods, tubes, or molded products. Method “B” is used for materials such as molding powder, ﬂakes, or pellets. Test Method “A” This method requires the use of a precision analytical scale equipped with a stationary support for an immersion vessel above or below the scale platform. Corrosion resistant wires for suspending the specimen and a sinker for lighter specimens with a speciﬁc gravity of less than 1.00 are employed. A sinker is used as an immersion vessel. Typical speciﬁc gravity test equipment is shown in Figure 11-25. The test specimen of any convenient size is weighed in air; this value is referred to as “S”. Next, the specimen is suspended from a ﬁne wire attached to the scale platform and immersed completely in the sinker with distilled water to determine the weight; this value is referred to as “D”. The weight of a specimen in water or sinker, if used, is determined, this value is referred to as “T”. The molded specimens are not tested for speciﬁc gravity until the mold shrinkage test has been accomplished, usually after 24 to 48 hours. The speciﬁc gravity of the specimen is calculated as follows: Speciﬁc gravity =

S (S + T − D)

Where: S = Weight of specimen in air D = Weight of specimen and wire (sinker, if used) suspended in water T = Total weight of specimen, immersed sinker (if used), and partially immersed wire Test Method “B” Method “B” is suitable for thermoplastic pellets, ﬂakes, or powder. It requires the use of an analytical scale, a pycnometer, a vacuum pump, and a vacuum desiccator. The test begins by first weighing the empty pycnometer. The pycnometer is ﬁlled with distilled water and placed in a water bath until an equilibrium in temperatures is attained between the two liquids. The weight of the pycnometer ﬁlled with water is determined; this value is referred to as “P”. After cleaning and drying the pycnometer, the plastic specimen and sinker, referred to as “S”, is added to the pycnometer ﬁlled with water, and the weight of the specimen and sinker plus the pycnometer with water is determined. The pycnometer is ﬁlled with water and placed in a vacuum desiccator. A vacuum is applied until all air has been removed from between the particles of the thermoplastic specimen. Lastly, the weight of the pycnometer ﬁlled with water and the specimen is recorded, this value is referred to as “T”. The speciﬁc gravity is calculated as follows: Speciﬁc gravity =

S (S + T − D)

Figure 11-25 Specific gravity test equipment

750

11 Performance Testing of Thermoplastics Where: S = Weight of the specimen (1.00 to 5.00 g) D = Weight of the pycnometer ﬁlled with water T = Weight of the pycnometer with water and the specimen If the water is replaced with another suitable immersion liquid, the speciﬁc gravity of the immersion liquid must be determined and taken into account in calculating the speciﬁc gravity of the thermoplastic material.

11.11

Density Gradient Testing (ASTM D-1505)

The density of thermoplastic materials is deﬁned as the weight per unit volume and is expressed in g/cm3 or lb/ft3. Density determinations by this method are very accurate and quick, making it a widely used technique, which requires very careful preparation and handling. 58 59 60 61 62 63 64

Figure 11-26 Density gradient column

The test method, developed to accurately determine the density of thermoplastics, is based on observing the level to which a test specimen sinks in a liquid column exhibiting a density gradient in comparison with standard specimens of known density. A number of calibrated glass ﬂoats of precisely known density are introduced into the density gradient and allowed to sink in the column to a point where the glass ﬂoat’s density matches that of the solution. A series of such ﬂoats of differing densities within the range of the column serves as a means of calibrating the column. The ﬂoat position versus ﬂoat density is plotted on a chart large enough to be read accurately to ±0.040 in to obtain a calibration line. When a specimen of unknown density is introduced into the column, its position on reaching equilibrium compared to the calibration line gives an accurate measurement of its density. The whole system is kept at a constant temperature of 73 °F. Figure 11-26 shows a typical commercially available density gradient column.

11.12

Water Absorption Testing (ASTM D-570)

Thermoplastic polymers absorb varying amounts of water and the presence of absorbed water may signiﬁcantly affect some properties. Electrical properties change the most with water absorption and this is one of the reasons that polyethylene, since it absorbs almost no water, is highly favored as a dielectric material. Water absorption characteristics of thermoplastics depend largely on the basic type and ﬁnal composition of a material. For example, thermoplastic polymers containing only hydrogen and carbon, such as polyethylene and polystyrene, are extremely water resistant, whereas thermoplastic polymers, such as cellulose acetate and nylon that contain oxygen groups, are very susceptible to water absorption. Materials containing chlorine, bromine, or ﬂuorine, such as ﬂuorocarbon (PTFE), are water repellent materials. Some thermoplastic polymers absorb very little water at room temperature with negligible effects on properties, but they absorb considerable amounts of water at higher temperatures, losing their properties rapidly. Materials that absorb relatively larger amounts of water tend to change dimensions in the process. When dimensional stability is required in products made of such materials, grades with fewer tendencies to absorb water are chosen.

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11.13 Impact Resistance Testing The water absorption characteristics of thermoplastic polymers are modiﬁed by compounding with different additives and reinforcements, which migrate to the outer surface of the molded product.

11.12.1 Test Specimen For molding materials, the specimens are discs with a diameter of 2.0 in and 0.125 in thickness. For sheet materials, the specimens are bars with 3.0 in × 1.0 in × wall thickness of the sheet material. The specimens are dried 24 hours in an oven at 122 °F, cooled in a desiccator and immediately weighed.

11.12.2 Test Procedure Water absorption data may be obtained by immersion for 24 hours or longer in water at 73.4 °F. Upon removal, the specimens are wiped dry with a cloth and immediately weighed. The increase in weight is reported as percentage gained. Percentage increase in weight during immersion is calculated as follows: Water absorption (%) = 100 ×

Wet Weight − Dried Weight Dried Weight

For materials that lose some soluble matter during immersion, such as cellulosic, the sample must be dried and weighed again after the ﬁrst test, reporting the amount as the “percentage soluble matter lost”. Water absorption (%) = Weight gain (%) + soluble matter lost (%)

11.13

Impact Resistance Testing

Impact resistance is a complex phenomenon that is sensitive to many testing variables. Impact strength is the ability to resist high rate loading. It is probably the most critical mechanical property of thermoplastics, because it relates to the service life of the part and inﬂuences the increasingly important issues of product safety and liability. Resin producers use impact testing as a quality control screening technique. Changes in impact resistance of thermoplastics by rubber modiﬁcation and other alloying methods are determined by impact testing; equally important is assessing the effects of ﬁller and reinforcement modiﬁcations on impact properties. The impact data must be reliable, reproducibl,e and relatable to values for other polymer types, as well as to real part applications in many cases. Product designers pursue a variety of impact tests, some homemade, seeking to simulate the conditions of the impact events for their products. Conventional impact tests often do not duplicate sufﬁcient real life impact parameters. Designing thermoplastic components to resist impact is done by developing geometry, then molding prototype parts, and ﬁnally performing impact tests that simulate some of the more extreme end use situations. Typical tests require striking blows by a variety of devices, or dropping the thermoplastic components

752

11 Performance Testing of Thermoplastics from a predetermined height. The necessity for this type of empirical testing arises from the complex nature of the impact phenomenon itself. The inherent resistance to impact of the thermoplastic material is only one of several major factors that inﬂuence the impact resistance of molded articles. Some of the other characteristics affecting the impact performance of thermoplastic components are the following: • Stress concentration resulting either from molding defects, such as voids, or from design features of the product, such as sharp corners (stress concentrations) • The shape and geometry of the striker. Sharp objects tend to concentrate stresses and increase the severity of the impact • The thickness of the thermoplastic molded part. Thin-walled structures are more brittle than heavy cross sections; they have less volume to absorb and dissipate energy • The geometry of the thermoplastic molded part and the impacted area can interact with other factors reducing the overall resistance to impact • Exposure to environmental stress, temperature, UV exposure, chemicals, and other factors can weaken impact resistance • The impact resistance of thermoplastic molded parts is affected by molding process variables, such as orientation, both in the polymer molecular chains and the type of reinforcements, polymer melt ﬂow, molded-in stresses, degradation, crystallization, morphology, internal voids, and weld lines. Impact strength is also often the ﬁrst property to deteriorate, as a result of environmental assaults. Impact testing can also be used to establish low and high temperature service limits and to evaluate UV resistance. Impact Fracture Mechanism The impact fracture mechanism is complex. High rate impact, before the polymer can fully obtain its viscoelastic strength, pulls apart or ruptures the polymer’s molecular chains. Chain length, packing, alignment, and bonding forces all affect resistance to fracture. Failure occurs in two stages. First, a crack is initiated in the sample or part, and then the crack propagates to failure. A polymeric structure that excels at resisting one of the stages may contribute little to resist the other. Impact tests that focus on one instead of both steps lead to inconsistent test results. Failure will occur in one of two modes, brittle or ductile failure. Brittle failure is characterized by a linear relationship between force and deﬂection from the impact peak force, where the crack is initiated. Little additional energy is required to fracture the sample. By comparison, ductile failure shows thermoplastic yield characteristics. A yield point appears at the maximum impact and considerable additional energy is needed to propagate the crack as the sample deforms. Most thermoplastics offer a choice: they will fail either in a ductile or a brittle manner, depending on a host of variables including temperature, sample geometry, and strain rate. High impact resistant thermoplastics, such as polycarbonate and “supertough Nylon 6/6”, are ductile, while a conventional polystyrene is brittle.

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11.13 Impact Resistance Testing Also to be considered is the ductile/brittle transition; for instance, when impact at low temperature causes a brittle fracture of a resin, while higher temperatures cause ductile failure. One objective of careful impact testing is to identify these transitions, not just with regard to temperature but with regard to strain rate as well. In addition, the geometry of the test specimen can shift the transition zone. The material that excels over another in impact resistance at a certain temperature or rate or thickness may not maintain that edge when the geometry of the part is changed. Several types of impact tests for thermoplastic molded products have been developed over the years for comparing the inherent impact resistance of different thermoplastics. Several impact test methods hace been adopted for quality control and speciﬁcations of the product. An impact test can be divided into ﬁve major classes and subdivided into many different types with slight variations. The ﬁve major classes are pendulum impact tests, falling weight impact tests, instrumented pendulum impact tests, high rate instrumented impact tests, and miscellaneous impact tests. These major tests are subdivided into: 1) Pendulum impact tests • Izod impact test • Charpy impact test • Chip impact test • Tensile impact test

5) Miscellaneous impact tests

0.006 inches

Specimen

4) High rate instrumented impact tests • Dynatup®, vertical high speed impact test • Horizontal high speed plunger impact test

0.50

1.25 inches

3) Instrumented pendulum impact tests • Izod impact test • Charpy impact test • Tensile impact test

Figure 11-27 Izod impact specimen “V” notch cutter

2.50 inches

2) Falling weight impact tests • Drop weight (tup) impact test • Gardner impact test

45˚

0.5˚

Radius 0.010 0.002 inches

11.13.1 Pendulum Impact Tests

Izod was a metallurgist who devised the test memorializing his name as a method for screening metals for cutting tools; it was subsequently applied to thermoplastics by the plastics industry. The notched Izod test is the ﬁrst of ﬁve procedures, Method “A”, of ASTM Standard D-256, “Impact Resistance of Plastics and Electrically Insulating Materials”. All D-256 methods use a pendulum arm to deliver the impact to the sample. In notched Izod, the specimens are injection molded bars, 0.50 in wide by 3.50 in long, and either 0.125 or 0.250 in thick, gripped over half their length by machining the notch and length of the specimen.

0.10

0.002 inches

Striking radius 0.031 0.005 inches

Specimen

Izod Impact Testing (ASTM D-256)

Vise

Pendulum

“V” notch at front

Vise

Figure 11-27 shows this machining procedure. The notched specimen is positioned as a vertical cantilevered beam to be impacted; the notch faces the pendulum. The dimensions of the notch are ﬁxed, as shown in Figure 11-28.

Figure 11-28 Notched Izod impact test (Method “A”)

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11 Performance Testing of Thermoplastics The pendulum is released from a ﬁxed height; from the height it reaches after impacting the sample, energy expended can be calculated in foot-pounds per inch of notch. One sample produces one piece of data.

Pendulum ready to impact Mounted specimen

The velocity of the pendulum at impact, a ﬁxed parameter, is approx. 8,300 in/ min. The test cannot identify the ductile/brittle transition rate unless the velocity is varied.

“V” notch at front Vise

Mounted izod specimen before being impacted by the pendulum Pendulum after the impact Broken specimen

Vise

The use of a notch implies quite rigid conditions. The intention of the notch is to approximate end use conditions; the notch serves as a stress concentrator (or stress riser). Concentrations of stress occur in parts, not just from nicks or other surface irregularities, but from part design, sharp corners, and moldedin stress caused by the processing conditions. The notch also dictates that the fracture be essentially unidirectional and not multi-axial like most real life impacts. This is in fact an artiﬁcial crack; consequently, the Izod test is primarily measuring the energy to propagate a crack. Impact resistance mainly recognizes the energy needed to propagate the crack. Figure 11-29 shows ﬁrst the notched specimen before impact, then the specimen fractured after impact. The Izod test has many shortcomings, the important ones among them:

Broken notched specimen by the impact of the pendulum

Figure 11-29 Notched Izod impact tester (Method “A”)

Thermo chamber

Izod impact tester

Figure 11-30 Izod impact tester with thermo chamber

• It is a single speed test. It is known from the viscoelastic nature of thermoplastics that impact resistance is likely to vary with a complex speed that will be different for each thermoplastic. Therefore, to characterize and compare different thermoplastics, testing should take into consideration the effect of speed rather than testing with a standard speed. • Different thermoplastics can have very different patterns of failure in this test, involving combinations of one or more of the following mechanisms: tnitiation of a crack, bending of the specimen, propagating the crack through the specimen, various kinds of tearing, rending, deforming, and separating the broken halves. Each mechanism makes a separate contribution to the energy extracted from the pendulum, with tearing and propagating being the largest. This method of testing favors soft, ductile materials, which bend and tear and ﬁber reinforced materials, in which much of the energy is used for tearing out ﬁbers after the specimen has failed. It penalizes stiff, hard, homogeneous materials, which break cleanly and elastically. The unnotched Izod test is Method “E” of ASTM D-256, in which the sample is simply turned around so that the notch is opposite the impact side and thus eliminated as a stress concentrator. The test, as run with engineering resins, usually produces the notation NB (no break) – a very limited piece of technical data. This unnotched Izod test is useful for reinforced and ﬁlled materials, especially for ﬁber glass reinforced compounds, which require little energy to crack. Figure 11-30 shows an impact pendulum thermo chamber used for performing impact tests at different temperatures.

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11.13 Impact Resistance Testing

Specimen

Notch at front Pendulum Anvil

Method “B” of ASTM D-256 enlists a pendulum tester to break a beam-type sample supported at the ends and notched on the side opposite impact. In this test, the specimen is mounted horizontally and supported at both ends. The dimensions of the notch are the same as in Method “A” for Izod impact tests. Again, the impact fracture is unidirectional. Only the specimens that break completely are considered acceptable.

Anvil

11.13.2 Charpy Impact Testing (ASTM D-256)

Like notched Izod, Charpy ﬁxes notch radius and depth and will show variations in values if those parameters, as well as thickness, are changed. It is not often used, because ductile materials, bending at impact, occasionally interfere with pendulum clearance. The sample is twice as thick as the typical Izod impact specimen. The Charpy impact strength is calculated by dividing the indicator reading by the thickness of the specimen. The results are reported in ft-lb/in of notch for notched specimens and ft-lb/in for unnotched specimens. The Charpy impact system is shown in Figure 11-31.

Schematic top view

Notch at front Specimen Anvil

11.13.3 Chip Impact Testing This test is considered to be particularly valuable for measuring the effect of surface micro cracking caused by weathering on impact strength retention. The material’s toughness is measured by this test as opposed to the material’s notch sensitivity as measured by the notched Izod impact test. The test also allows determining the orientation, ﬂow effects, and the weld line strength.

Pendulum Charpy impact equipment

Figure 11-31 Charpy impact test

If the test is used exclusively to study the effect of weathering on a polymer, the sample must be struck with a pendulum hammer on the side exposed to weather. The chip impact test is also useful in measuring the relative toughness of a rather large, complex shaped part that is difﬁcult to hold in a conventional ﬁxture. The chip impact test is somewhat similar to the Izod impact test. The chip impact test requires the use of a pendulum type hammer device and a holding ﬁxture for the specimen. The test specimens are usually 1.00 in long × 2.00 in wide and 0.065 in thick. They can be prepared either by injection or compression molding or by cutting them from a sheet. The test is carried out by mounting the test specimen on the chip pendulum impact tester as shown in Figure 11-32. The pendulum hammer is released, allowed to strike, and swing through the specimen. The retained toughness is proportional to the energy absorbed during impact that is measured by the angle of travel of the pendulum after impact. The value is expressed in in-lb/ in2.

11.13.4 Tensile Impact Testing (ASTM D-1822) The tensile impact test is similar to the Izod test, with some adjustments to the ﬁxturing. The specimen geometry is of the “dog bone” type. One end is clamped in the pendulum head and the other in a crosshead clamp. When the pendulum swings down, the crosshead clamp holding the trailing end of the sample is halted by an anvil, the sample breaks and the pendulum head, now bearing half the specimen, swings up to a height that is translated into energy to break. The tensile impact test is shown in Figure 11-33.

Specimen Pendulum Vise

Figure 11-32 Chip pendulum impact tester

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11 Performance Testing of Thermoplastics

0.375

Type “S” = 0.00 Type “L” = 0.38

0.50 R.

0.125

Specimen (inches)

Specimen

Pendulum head

Anvil

Anvil

Specimen

Schematic top view

Anvil

Although velocity at impact is ﬁxed, the strain rate varies with the choice of the specimen’s lengths. The short necked type “S” promotes little extension under impact, so brittle fracture is more likely. The test gives the energy per area to fracture the sample, but does not differentiate between type “S” and type “L”. For instance, two materials can produce the same value, one requiring a large force and little elongation to fail for type “S”, and the other just the opposite for type “L”. The test is uniaxial and does not relate directly to the mechanics of impact events. As the standard states: The data variations may be due to different failure mechanisms within a group of specimens. Some materials may exhibit a transition between different failure mechanisms.

Crosshead clamp

Pendulum

The tensile impact test is simply a tensile test at a high rate of speed as compared to normal tensile testing (velocity at impact is approximately 11.3 ft/s) and can be performed on high rate tensile (or universal) testers as well. It was devised because of the limitations of Izod and Charpy tests, eliminating the notch as a factor, and because those tests do not break some ductile materials, especially thin sheets and ﬁlms.

Crosshead clamp Test equipment

Figure 11-33 Tensile impact test

11.13.5 Drop Weight Impact Testing (ASTM D-3029) This drop weight test is often referred to as Gardner testing; a weight is dropped on a specimen positioned over a circular opening. The testing parameters are established in ASTM D-3029: “Impact Resistance of Rigid Plastic Sheeting or Parts by Means of a Tup (Falling Weight)”. Its utility for both samples and parts indicates why many prefer this test. Impact is multi-axial, as in many real life events. While the notch directs crack initiation in Izod and Charpy specimens, this test relies on the anisotropic nature of thermoplastics to divulge the easiest route to failure. Figure 11-34 shows a typical Gardner drop weight impact tester. The test does not seek to break the specimen completely. Crack initiation or a speciﬁc deformation may be enough to deﬁne failure. The problem is that many samples must be consumed to establish a value. That number represents the ft-lb required to cause 50% of the specimens tested to fail, whatever has been deﬁned as failure. The Bruceton Staircase statistical exercise is used to zero in at the right impact level. For quality control screening one could establish the height and weight of the tup for a pass/fail determination.

Tower

Falling weight tup

Specimen Base with hole

Figure 11-34 Gardner drop weight impact tester

In drop weight testing one can vary height and hold weight constant, or the opposite, or both can vary. A 100.00 lb weight dropped from 1.00 ft, or a 10.00 lb weight dropped from 10.00 ft both deliver 100.00 ft-lb at impact, but the velocities are quite different, which may be a possible source of misdirection. At the expense of many specimens, one can identify the ductile/brittle transition rate. Another source of variability is found in the tup (falling weight) and the opening dimensions. An early version of the test had both almost equal, so that there was little clearance for the tup in the opening and the specimen often suffered shear fracture, hardly a representative parallel to actual impacts. The latest Gardner impact probes and circular dimensions of the bases are shown in Figure 11-35.

757

11.13 Impact Resistance Testing The dimensional relationship between probe and opening does affect the strain rate in the material. However, there are limitations. Drop weight tests usually give a larger data scatter because of mechanical difﬁculties in controlling the attitude of the striker without impeding free fall. The failure criterion is usually visual (ﬁrst appearance of cracks). Inherently, it is a “go” or “no-go” type of test that does not yield insight into material performance.

0.625 dia. Impact tup Specimen Base

Impact strength measured by dropping a weight pays more attention to the energy required to initiate a crack, especially for ductile materials.

0.640 dia. 0.625 dia.

11.13.6 Falling Weight Impact Testing The falling weight impact test, also known as the drop impact test or the variable height impact test, employs a falling weight tower. This falling weight may be a tup with a conical nose, a ball, or a ball-end dart. The energy required for the specimen to fail is measured by dropping a known weight from a known height onto a test specimen. The impact energy is normally expressed in ft-lb and is calculated by multiplying the weight of the projectile by the drop height. Figure 11-36 shows a typical falling weight impact tester.

1.50 dia.

Figure 11-35 Gardner drop weight impact test configurations (in)

The biggest advantage of the falling weight impact test over the pendulum impact test is its ability to duplicate the multi directional impact stresses that an injection molded thermoplastic product would be subjected to in actual service. The other obvious advantage is the ability to use specimens of different sizes and shapes, including an actual product. The falling weight impact tests introduce polyaxial stresses to the specimen and measure its toughness. The test is carried out by raising the weight to a desired height manually or automatically with the use of a motor driven mechanism and allowing it to fall freely onto the other side of the round nosed punch. The punch transfers the impact energy to the ﬂat test specimen positioned on a cylindrical die or an actual product lying on the base of the machine. The kinetic energy possessed by the falling weight at the instant of impact is equal to the energy used to raise the weight to the height of drop and is the potential energy possessed by the weight as it is released.

Tower

Since the potential energy is expressed as the product of weight and height, the guide tube can be marked with a linear scale representing the impact range of the instrument in in-lb. Consequently, the toughness or the impact strength of a specimen can be read directly off the calibrated scale in in-lb. A number of different techniques are used to determine the impact strength value of a specimen. One of the most common methods used is called the Bruceton Staircase Method in which testing is concentrated near the mean to reduce the number of specimens required. An alternative method, known as the ultimate non-failure level (UNF), requires testing the specimen in successive groups of ten (10). One missile weight is employed for each group and the weight is varied in uniform increments from group to group. The variations in the test results due to the ﬁllers and reinforcements, clamping pressure, and material orientation are virtually eliminated in the falling weight impact test. This type of test is also very suitable for determining the impact resistance of injection molded thermoplastic products, ﬁlms, and sheets.

Variable weight

Chuck for falling tup

Place for specimen

Figure 11-36 Falling weight impact tester

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11 Performance Testing of Thermoplastics

11.13.7 Instrumented Impact Testing The previous impact tests produce a single impact value for each specimen under standardized conditions. The instrumented impact tester will provide more detailed information, allowing variables to be manipulated to obtain the desired analysis. Instrumented impact procedures seek to deﬁne the whole impact event by plotting load against time or deformation. Computerized data acquisition and reporting can be speciﬁed with this instrumented impact tester. 11.13.7.1 Instrumented Pendulum Testers Figure 11-37 shows an instrumented impact pendulum tester from Ceast, with instrumentation from Effects Technology (Dynatup®). The striking tup is ﬁtted with a strain gage. During testing, a light sensor triggers an oscilloscope just before impact. The output of the strain gauge bridge circuit is recorded by a dual beam oscilloscope; the signal can be directly analyzed as an analog of load time history of the specimen. That history is short. The interval of contact between tup and specimen is typically 0.10 to 1.0 ms. Figure 11-37 Instrumented Izod impact pendulum tester (Courtesy: Ceast)

In comparison, the instrumented Charpy testing generates large shear and tensile forces and is particularly effective for the characterization of inter-laminar shear and bonding forces in high strength composite materials. 11.13.7.2 Instrumented Impact Testing (Dynatup®)

Small dynatup®

Large dynatup®

Two sizes of dynatup® testers

Tup

One of the biggest drawbacks of the conventional impact test method is that it provides only one value: the total impact energy and nothing else. The conventional tests cannot provide additional information on the ductility, dynamic toughness, fracture, and yield loads or the behavior of the specimen during the entire impact event. The instrumented impact test is a reliable and fast method for obtaining accurate data on the dynamic properties of thermoplastic materials. Specimens can range in size from substandard impact specimens to entire assemblies and structures. By instrumenting the vertical tower impact tester, a complete load and energy history of the specimen become available. An instrumented impact system monitors and precisely records the entire impact event, from initial impact and acceleration from rest to thermoplastic bending to fracture initiation and propagation to failure. Typically, a specimen is in physical contact with a striking tup from 0.10 to 1.00 ms. With the load-time and energy-time curves produced by this equipment, a wealth of information on the dynamic properties of the specimen is readily available. This includes yield strength, fracture load, ductility, fracture initiation, propagation energies, shear lip energy, total impact energy, crack arrest, critical ﬂaw size, ductility transition temperature, and loading and unloading characteristics. Figure 11-38 shows two sizes of Dynatup® instrumented impact testers and a close-up view of a Dynatup® impacting a complete thermoplastic housing.

Specimen loaded Dynatup® impacting a thermoplastic housing

Figure 11-38 Instrumented impact (Dynatup®) testing equipment

An instrumented impact test system can measure and record the load-time and the energy-time behavior of a specimen during the impact event. The striking bit or tup of the impact machine is instrumented with strain gauges and properly calibrated. During the actual test, a ﬁber optic device triggers an oscilloscope just before the instrumented tup strikes the specimen. The output of the strain gauge

759

11.13 Impact Resistance Testing

The apparent total energy absorbed by the specimen can be calculated and plotted against time. The specimen displacement can be calculated by the double integration of the load-time curve and the load-displacement curve, which can be plotted. With the advent of microprocessor technology, some manufacturers are now capable of offering a unit that automatically calculates the sample displacement and provides a load-displacement curve, eliminating the need for calculations. Many other useful data, such as the impact rate, force and displacement at yield and at break, failure energies, as well as the modulus of elasticity are calculated and printed. An instrumented drop weight test system that provides versatility in the size and shape of the specimens can be accommodated. Also, these systems permit dynamic tests to be made at different impact energies and velocities. The test system’s ability to easily and inﬁnitely vary velocities from 2.00 ft/s to 28.00 ft/s and energy from 2.00 ft-lb to 5,700 ft-lb is signiﬁcant because such variations now are being incorporated into many standard test procedures. The drop weight impact test consists of a drop tower and instrumentation package. The drop tower is a gravity driven impact machine equipped with remote controls for release of the hammer and tup assembly and a motorized lift mechanism for easy return of the hammer to a predetermined drop position. An automatic rebound brake is standard. The hammer and tup assembly are designed so that the weight of the hammer can be varied easily and the tups interchanged quickly and easily. The tup design is based on actual test requirements. The base plate permits easy interchangeability of anvil supports to ﬁt a range of specimen geometries. Maximum energies and impact velocities are achieved by combinations of three basic options: • Increased weights for increasing maximum energy • Elastic propulsion acceleration kit for increasing impact velocity by accelerating the hammer/tup assembly at rates greater than normal gravitational values • Extended drop height for boosting maximum impact velocity Instrumented impact tests are used for testing automotive components, safety equipment, threaded fasteners, furniture, sports equipment, and others.

Load

1.00 0

Energy

2.00

0.50

0 0

1

2

Time (sec.) PBT-30% glass reinforced load = 0.44 & energy = 1.80

2.00

0.50 1.00 0

Energy

The trace is retained on the oscilloscope and photographically recorded to provide a permanent record of the load-time and energy-time behavior of the specimen, or the velocity of the striking head. In a sense, it is like a stress-strain curve obtained from a tensile test, but it also yields additional information on the dynamic performance of materials. Figure 11-39 shows the Dynatup® graphs with the results from the dynamic impact testing of 0.125 in tensile bar specimens molded of 30% glass reinforced PET and 30% glass reinforced PBT. This test was performed to compare the impact resistance of these two thermoplastic materials.

PET-30% glass reinforced load = 0.66 & energy = 2.30

Load

bridge circuit is then recorded by the dual beam storage oscilloscope during the small interval of time that the tup is in physical contact with the specimen. The signal produced during this contact can be analyzed directly as an analog of the load-time history of the specimen. The testing procedure does not interfere in any way with the measurement of impact energy.

0 0

1

2

Time (sec.) Figure 11-39 Dynatup® comparison using 0.125 in tensile bar

760

11 Performance Testing of Thermoplastics For example, the instrumented impact test is used to check the durability and quality of glass reinforced nylon automotive cooling fans. Testing time for the instrumented system is less than an hour, compared to over 80 hours for previously used fatigue testing methods.Values from the instrumented test are compared to values from the fatigue test and a linear relationship is found with little data scatter. Data provided by the test include a complete load-time and energy time record, including fracture load, time to fracture, and total impact energy. In another automotive application, an automotive component’s manufacturer uses an instrumented impact test system to evaluate the relative toughness of the threaded section of a heat-treated fastener. Overheating, which adversely softens the smooth shank section, produces excessive toughness in the threaded section. Similarly, underheating results in insufﬁcient toughness. As a result of the testing program, maximum and minimum limits on the allowable toughness were set. 11.13.7.3 Instrumented High Speed Plunger Impact Tester This versatile high-speed impact testing machine is capable of testing everything from a thin ﬁlm that may require as low an impact rate as 30.00 in/min to a thermoplastic automotive bumper that may require a high impact rate up to 60,000 in/min. The specimen or product can be tested under a controlled environment of temperature and humidity. A commercially available high speed plunger impact tester is shown in Figure 11-40. The equipment basically consists of an instrumented horizontal test plunger attached to a hydraulically powered actuator. The hydraulic power is delivered by an accumulator that is charged by a pump. The force is detected with a quick responding quartz load cell mounted directly on the actuator. The velocity can be set digitally from 30 to 60,000 in/min. The actuator is capable of moving in either direction to make both tensile and impact tests possible. The starting and stopping point of the actuator can be set for the repeated test and for the control of penetration depth in thick samples. The horizontal actuator design permits virtually unlimited variations in sample geometry, size, or orientation. The tester is equipped with a “CRT” and a plotter that automatically displays load versus displacement data. A built-in microprocessor provides a more useful data handling system. High rate impact testers provide a great deal of information: force, displacement, energy, and velocity are retrieved; load/deﬂection curves are plotted; and data obtained includes apparent modulus, yield and ultimate strength, and energy at yield and break. This information is also useful in deriving parameters, such as fracture toughness, and for use in elastic and plastic stress analyses. High rate impact testers have been used for material evaluation.At low strain rates, some polymers fail in a ductile manner. The same polymers appear to show brittle failure at high strain rates. The point at which this ductile-to-brittle transition takes place is of particular importance. A high rate impact test can be provided in a graphical form. Tests can also be carried out at different temperatures to ﬁnd ductile-to-brittle transition points at various temperatures. 11.13.7.4 Miscellaneous Impact Testing

Figure 11-40 Instrumented high speed plunger impact tester (Courtesy: Du Pont)

Depending on the product end use requirements, many different types of impact tests have been devised to simulate actual impact conditions. Underwriter’s Laboratory requires a fully assembled, working electronic housing to withstand 5.00 ft-lb of impact from a swinging ball at the most vulnerable external spot to check the damage to the electronics. Small electrical units susceptible to being

761

11.14 Creep, Rupture, Relaxation, and Fatigue knocked off a desk must pass a 3.00 ft drop while the units are plugged in the electrical outlets. Drop impact resistance of a blow molded thermoplastic container is a standard ASTM D-2463 test. The drop impact resistance is determined by dropping conditioned blow molded containers ﬁlled with water from a platform onto a prescribed surface. Another type of impact test, called the air cannon impact test (ACIT), is used to determine the toughness of rigid thermoplastic exterior building components. The air cannon impact tester consists of a compressed air gun that propels spherical thermoplastic projectiles at a test specimen. Interchangeable barrels allow projectiles of varying sizes and weights to be used. Projectile velocity is controlled by varying air pressure. Polypropylene and polyethylene molded balls of different sizes are used as projectiles to simulate the effect of different sized hail stones. The weight and the velocity of the projectile are used to calculate the impact force absorbed by the test specimen. This test more realistically simulates end use environmental conditions than conventional impact tests and provides important information for evaluating a product and establishing product quality.

Creep, Rupture, Relaxation, and Fatigue

Thermoplastics are viscoelastic materials; their strength and deformation under stress are dependent on load duration. Therefore, to predict the performance and anticipate the useful life of a material, it must be tested for properties that include time as an independent variable. The test for creep is to predict deformation. The strength and the related stress relaxation tests are to measure the loss of stress at constant strain.

11.14.1 Tensile Creep Testing Tensile creep measurements are made by applying a constant load to a tensile test specimen and measuring its extension as a function of time. The extension measurement can be carried out in several different ways. The simplest way is to make two gauge marks on the tensile specimen and measure the distance between the marks at speciﬁed time intervals. The percentage of creep strain is determined by dividing the extension by initial gauge length and multiplying by 100. The percentage of creep strain is plotted against time to obtain a tensile creep curve. Figure 11-41 shows the long term behavior of acetal homopolymer

3.0

2.0 1.5 1.0

at 73° F. Air

2.0

ur

10

1.0

ho

ur

2.5

2.0

1.

1.5 1.0

10.0

2.5 s

1.5

ur

1.0

s

0.5

ho

0

0

0

10

0.5

ho

Stress, (psi)

2.5

Stress, (psi)

When a thermoplastic is subjected to a constant load, it tends to deform gradually with time beyond the initial deformation predicted by its modulus. This characteristic is known as creep or cold ﬂow. If the load is removed after a time, the thermoplastic may or may not recover the total deformation, depending on the magnitude of the stress. Creep and stress relaxation are the most important mechanical properties of thermoplastics. They are more important than stressstrain properties, because they more accurately reﬂect the actual behavior of thermoplastics in use. At the same time, they can be substituted numerically in engineering formulas. In the design of thermoplastic injection molded products that are intended to support a load over a long time, creep should always be considered as one of the important design variables. These creep properties can be obtained in tension, compression, or ﬂexure.

1 .0 h ou 10 h r ours 100 hou rs 1.0 00 5.0 hour s 00 ho urs 10 .00 0h ou rs

11.14

0

00

ho

hou 00 5.0 s r u o 00 h

ur

s

rs

0.5

at 185° F. Air 0

0

0.5

1.0

1.5

2.0

2.5

Strain, (%) Figure 11-41 Long-term behavior of acetal homopolymer under load

762

11 Performance Testing of Thermoplastics under load at temperatures of 73° and 185 °F. The tensile stress values are also determined at speciﬁed feed time intervals to facilitate plotting a stress-rupture curve. The more accurate measurements require the use of a strain gauge, which is capable of measuring and amplifying small changes in length with time and directly plotting them on a chart paper. The test is also carried out at different stress levels and temperatures to study the effect on tensile creep properties.

11.14.2 Flexural Creep Testing Flexural creep measurements are also made by applying a constant load to a standard ﬂexural test specimen and measuring its deﬂection as a function of time. A typical test set-up for measuring creep in ﬂexure is shown in Figure 11-42. The deﬂection of the specimen at mid-span is measured using a dial indicator gauge. The electrical resistance gauge may also be used in place of a dial indicator. The deﬂections of the specimen are measured at a predetermined time interval. The percentage of ﬂexural creep strain is calculated using: εC = 6 × t × δ ×

100 L2

Where: εC = Maximum creep strain percentage (in/in) δ = Maximum deﬂection at mid span (in) t = Specimen thickness (in) L = Span support length (in) The percentage of creep strain is plotted against time to obtain a ﬂexural creep curve. The test is carried out at various stress levels and temperatures and similar ﬂexural creep curves are plotted. If necessary, the maximum ﬁber stress for each specimen in psi can also be calculated as follows: σ =

3×P×L 2 × b × t2

Where: σ = Stress (psi) W = Load (lb) L = Span support length (in) b = Specimen width (in) t = Specimen thickness (in) There is no established method for determining creep. ASTM test D-674 is actually not a test method but “A Recommended Practice for Creep Tests”. It discusses the complications of measuring creep and the precautions to be taken in using creep data. Complications arise largely because the creep measurements are made over a long time, ranging from months to several years.

Figure 11-42 Laboratory equipment for measuring flexural creep (Courtesy: Du Pont)

Creep tests are carried out by placing test specimens in a constant temperature/ humidity chamber and attaching a load to the test specimens either directly or through a level system for multiplying the load.Very careful measurements of the deﬂection or elongation of the test specimens are made over time, starting with hourly, then daily, and ﬁnally weekly measurements. Specially built micrometers are used for measuring the small changes that are of interest.

763

11.14 Creep, Rupture, Relaxation, and Fatigue

Creep measurements are carried out under constant stress (load) conditions and the deformation (strain) continues to increase with time. Data concerning the initial portions of stress-strain diagrams are often summarized by calculating the modulus of elasticity (E) by dividing the stress (σ) by the strain (ε): σ (Stress) E (Modulus of Elasticity) = ε (Strain)

High load

Deformation, (inch)

The original data from creep experiments often are plotted on a deformation versus time graph as shown in Figure 11-43.

Medium load Low load

Time, (hours)

Similarly, creep data are put into a useful design engineering form by calculating a characteristic called an “apparent modulus” (EA). It is calculated by dividing the stress by the strain. As strain continues to increase with time, the apparent modulus decreases with time. Therefore, a value of the modulus applies only at a particular time. However, this method makes it easier to use creep data in engineering formulas and when plotted on logarithmic coordinates, this curve is often linear. This permits a comparison between creep curves for different materials. Creep data are usually summarized in a graph of apparent modulus versus time for several levels of constant stress. An example of this graph is shown in Figure 11-44. With this data it is possible to calculate the deformation expected at a given time.

Apparent modulus, (psi)

Figure 11-43 Flexural creep test data – deformation versus time

Low stress High stress

Time, (log scale, hours)

Figure 11-44 Apparent modulus versus time for high and low stresses

The short term stress-strain data is of little practical value in actually designing the product, because such data does not take the effect of long term loading on thermoplastics into account. Creep behavior varies considerably among types of thermoplastics; however, under proper stress and temperature conditions, all thermoplastics will exhibit a characteristic type of creep behavior. Figure 11-45 shows the creep behavior of thermoplastics in two different graphs to provide a better understanding of creep characteristics. The total creep graph is divided into four continuous stages. The ﬁrst stage (A-B) represents the instantaneous elastic deformation. This initial strain is the sum of the elastic and plastic strain. The ﬁrst stage is followed by the second stage (B-C), in which strain occurs rapidly but at a decreasing rate. This stage, where creep rate decreases with time, is sometimes referred to as creep or primary creep. The straight portion of the curve (C-D) is characterized by a constant rate of creep. This process is called “cold ﬂow”. The ﬁnal stage (D-F) is marked by an increase in creep rate until creep fracture occurs.

Second creep

First creep

Third creep

Breakage

F Breakage

D C Sta

t0 t1

Strain, (%)

C A

App

0

B

B

Removing load t2

t 0’

t 1’

t 3’

Time, (t)

Figure 11-45 Thermoplastic short-term creep behavior graphs

2 ge

aren t mo dulu s

1

A = Recovery B = Instantaneous recovery C = Full recovery

e3 Stag

Strain, (%)

2

Stage 1

A

Time, (t)

Sta

ge 4

764

11 Performance Testing of Thermoplastics 4

Isochronous Stress-Strain Curves It is necessary to compare various thermoplastic materials to select the best resin for a product design application. The basic creep curves, such as creep strain versus time or apparent modulus versus time, are not completely satisfactory. The comparison is extremely difﬁcult to make, especially when different stress levels are used for different materials.

Strain, (%)

3

00 4.0

2

i) (ps si) 0 (p 3.00

1

2.000 (psi)

(psi) 1.000

0 0

10

1.000

100

10.000

Time, (hours)

For example, a design requires a thermoplastic shelf to withstand a continuous load for 1,000 hours. If after 1,000 hours of continuous loading, the deformation is not to exceed 2.0%, what should be the maximum allowable stress? To solve this problem, the designer needs to erect an ordinate on the basic creep strain versus time curve at 1,000 hours, as shown in Figure 11-46. From Figure 11-46, the designer can determine the strain value at different stress levels. These values allow the designer to plot another curve of stress versus strain at 1,000 hours (Figure 11-48).

Figure 11-46 Creep strain vs. time at different stress levels

Figure 11-47 shows the stress-strain creep curve at different times.

O

O

Creep time between T and TO = εT – ε0 (%). The creep modulus ET (psi) at any time, temperature, and stress-strain condition is identical to the corresponding “apparent modulus” (EA). ET for design in creep applications at stress and time T is the slope of the secant from the origin to the point σO εT (psi).

T

O

Stress, (psi)

O

A wide variety of polymers can easily be compared by studying isochronous stress-strain curves at 1,000 hours, such as those shown in Figure 11-48. O O

T

11.14.3 Procedure for Applying Creep Modulus

Strain, (%)

Stress, (psi)

Figure 11-47 Stress-strain creep curves at any given time

4.000

The following procedure is recommended for the use of creep modulus data: • Select the required service or operational life for the thermoplastic molded product.

PVC (pipe)

3.000

AB

S

2.000

er polym PP co ac t PP-impmer copoly

PP homo 1.000

HDP

E LDPE

0 0

1

2

3

4

5

6

Strain, (%)

Figure 11-48 Isochronous stress-strain curves at 1,000 hours for common thermoplastic polymers

7

• Consult or plot (using the stress-strain creep curves) the creep modulus curve of the selected material for the temperature at which the product will be exposed, extrapolating where necessary. Like creep rupture, creep modulus varies greatly with temperature. In addition, it is also subject to stress level. For each thermoplastic material grade, creep modulus data are tabulated in a creep graph, ﬁrst by test temperature and second by applied (test) stress. For very rigid materials, such as thermosets, reinforced thermoplastics, and amorphous thermoplastics at room temperature and below, the creep modulus curves show minor variations with applied stress. However, for the more ﬂexible and ductile thermoplastics, the creep modulus curves will vary signiﬁcantly and systematically with stress level. The higher the stress, the lower the creep modulus. This is a consequence of viscoelasticity. To cope with the effects of stress levels, the designer has two alternatives. If the stress level can be predetermined, the designer should select the creep modulus curve with the closest stress level. If the stress level is not known, the designer should choose a creep modulus curve at a conservative stress level and check the choice after calculating a stress level. • Read from the selected creep modulus curve the modulus value corresponding to the design life selected in the ﬁrst step. This is the design modulus. • Apply a safety factor to the design modulus to calculate a working modulus for any uncertainties arising from extrapolations or other compromises that may have been made. Safety factors of 1.50 to 1.75 are typical.

765

11.14 Creep, Rupture, Relaxation, and Fatigue • Substitute the calculated working modulus in the part design equation. For example, to calculate the width of a simple rectangular beam required to limit the maximum deﬂection to a speciﬁed value when the span and wall thickness of the beam are ﬁxed, the working modulus would be substituted for E in the following design equation. b=

W × L3 E × δ × 4 × t3

Where: b = Width of beam (in) t = Wall thickness of beam (in) L = Span (in) W = Load (lb) δ = Maximum deﬂection allowed for the beam (in) E = Working modulus (psi)

If a thermoplastic is loaded in tension and permitted to creep indeﬁnitely, it will tend to rupture at stresses well below its stress-strain yield or tensile strength. This characteristic is known as creep rupture. The time to failure depends on the characteristic of each thermoplastic, stress, end use temperature, and environmental conditions. Ductile types of thermoplastics often go through a third stage of creep, which involves necking, drawing, and drastic elongation before actual rupturing. Figure 11-49 shows the apparent modulus versus time for various stress levels at 150 °F for a high molecular weight polyethylene (HMWPE). Creep rupture tests are generally conducted by starting with stresses that cause rupture in minutes and gradually lowering the stress until very long failure times are reached. A stress level with a safety factor selected for the application can serve as a basis for design with a safe expectation that failure will not occur within the useful life of a thermoplastic injection molded product. The data obtained using this method forms the basis for designing thermoplastic pipe to withstand internal pressure at 73 °F. The tests are made by pressurizing samples of pipe, because the stresses in a pipe wall are polyaxial and complex and cannot be related numerically to tensile or other simple stresses. The essence of this method consists of conducting time to failure versus hoop stress tests for 10,000 hours (more than a year), plotting the data on logarithmic coordinates, and extrapolating the data to 100,000 hours or 10 years. 100,000 hours of stress divided by a safety factor is approximately the design stress, which will be used to calculate pipe wall thicknesses. If the same types of pipe tests are conducted at elevated temperatures, it will be found that thermoplastics may be subject to premature failure that substantially reduce their life expectancy at these temperatures. The shape of stress versus time curves in ductile thermoplastics usually shows a change in failure mechanism from ductile to brittle. 11.14.3.2 Stress Relaxation If a thermoplastic application is subjected to constant deformation rather than constant load (gaskets, seals, and valve seats), it tends to undergo stress relaxation;

Apparent modulus, (psi)

11.14.3.1 Creep Rupture 5.000 4.000

600 (

800 (

psi)

psi)

3.000

900

2.000

70

0(

(psi

)

ps

i)

1.000 0 0.01

0.1

1

10

100

1.000

Time, (hours) Figure 11-49 HMWPE apparent modulus vs. time curves, various stress levels at 150 °F

766

11 Performance Testing of Thermoplastics i.e., the stress induced in the thermoplastic by the initial deformation tends to cause the loss of mechanical properties with time. This can affect the ability of a thermoplastic to maintain a seal and stay tight in an assembly over a long time and is also important in mechanical joining of thermoplastics, such as bolting. Although most thermoplastic applications that involve stress-relaxation are compressive and are tested in compression, unfortunately, most of the data available is in tension. The test is conducted by straining a specimen to a predetermined deformation and measuring the resultant stress as it loses the mechanical properties with time. Like creep, to which it is theoretically related, stress relaxation varies with strain level and temperature and the data is more convenient to plot and use as apparent modulus (stress divided by original strain). Creep, creep rupture, and stress relaxation data do more than provide realistic values of strength and deformation. They describe the way most thermoplastics really behave under stress. The curve presents a more realistic view than can be obtained from any amount of stress-strain or data sheet properties. 11.14.3.3 Fatigue Fatigue describes the failure of materials after being subjected to repeated loads. These repeated loads may be bending, stretching, compressing, or twisting. After being subjected to such forces for many cycles (hundreds, thousands, or millions), the material may break under a considerably lower load than the one that would cause failure on a single loading. Therefore, fatigue relates to the material becoming tired and weaker after cyclic distortion. At high loads, thermoplastic materials become weaker and weaker under all conditions of loading; but at low loads, thermoplastic materials can be loaded repeatedly forever without the loss of strength. This is the purpose of fatigue tests: to determine the highest level of repeated loading that a material can withstand in long-term use. Such a level of loading is called a “fatigue endurance limit”. The term has been borrowed from metallurgical and mechanical engineers, as this property was ﬁrst studied in metals. Basically, fatigue tests consist of evaluating a series of test specimens under different loads. Test specimens subjected to a high load fail in a short time (few cycles) and those under lower loads after a long time (many cycles). Fatigue testing is an extremely complex subject because of the number of possibilities concerning test conditions. Considering tension and compression only, there are still a number of general conditions. The test specimen may be pulled in tension only or the load may cycle from tension to compression. The test specimen may be preloaded, for example, with a low load in tension and cycled to a high load in tension. The test must also be deﬁned whether it is a constant load or a constant deformation test. In the ﬁrst case, with the constant load, the deformation would continue to increase. In the second case, when constant deformation is desired, the load would continuously decrease. These complications are mentioned to caution the reader that there is not a single standard fatigue test and that the test conditions should be fully understood before applying fatigue data to a new application.

767

11.16 Vicat Softening Point (ASTM D-1525) ASTM test D-671 is entitled, “Repeated Flexural Stress (Fatigue) of Plastics”. It describes a testing machine and gives a procedure for a constant deﬂection test. This test procedure presents a good discussion of fatigue testing and speciﬁcally points out that it is a research type test. This ASTM test has not been used for a number of years. Rather, fatigue tests have been carried out with Sonntag fatigue machines. Such testing machines are widely used and test the samples under constant load conditions. Molded dumb bells about 2 in long, with a cylindrical cross section, are frequently used as test samples. Both ends of the specimen are fastened in the testing machine with Jacob chucks and the apparatus repeatedly loads the sample at 1,800 cycles per minute. It is common practice to test with a completely reversed tension and compression cycle. It turns out that the samples are deformed (strained) about 1.0 to 2.0% during each part of the cycle. Routine practice has been to test for 2,000,000 cycles, unless the test specimen fails before this.

11.15

Melting Point Test (ASTM D-795)

The basic method used for melting point determination is described in ASTM D-795. A Fisher-Johns melting point apparatus, as shown in Figure 11-50, is most commonly used. The apparatus consists of a rheostatically controlled heated block, a thermometer, and a viewing magniﬁer. A small pellet or a sliver of the thermoplastic material to be tested is placed on the electrically heated block along with a few drops of silicone oil. A cover glass is placed over the material and the heat is gradually increased until the sample material melts or softens enough to deform. The meniscus formed by the oil is viewed through the magniﬁer. The temperature at which the meniscus moves is considered the melting point. The expected accuracy of the test is within ±5 °F of the published literature value. This method can be used for both semi-crystalline and amorphous thermoplastics. All semi-crystalline thermoplastics have a sharp melting point, which is much easier to detect. Amorphous thermoplastics melt over a wide range of temperatures and an exact melting point is difﬁcult to determine.

11.16

Vicat Softening Point (ASTM D-1525)

The Vicat softening temperature is the temperature at which a ﬂat ended needle of 1.0 mm2 circular cross section will penetrate a specimen to a depth of 1.0 mm under a speciﬁed load, using a selected uniform rate of temperature rise. This test is very similar to the deﬂection temperature under load test and its usefulness is limited to quality control, development, and characterization of materials. The Vicat softening temperature is a good way to compare the heat softening characteristics of thermoplastics. However, the test is not recommended for ﬂexible PVC or ethyl cellulose or other materials with a wide Vicat softening range. The test apparatus designed for deﬂection temperature under load test can be used for the Vicat softening temperature test with modiﬁcations. Flat specimens must be at least 0.50 in wide and 0.125 in thick. Two specimens may be stacked, if necessary, to obtain the necessary the thickness. Specimens may be compression or injection molded.

Figure 11-50 Fisher-Johns melting point tester

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11 Performance Testing of Thermoplastics Figure 11-51 shows the apparatus for testing Vicat softening point; it consists of a temperature regulated oil bath with a ﬂat ended needle penetrator mounted to register the degree of penetration on a gauge. The needle with a 1,000 g load is placed on the specimen. The temperature of the bath is raised at the rate of either 50 °C/h or 120 °C/h. The temperature at which the needle penetrates 1.0 mm is the Vicat softening point.

Schematic configuration of test

11.16.1 Melting Point, Glass Transition Temperature Polymers having a stereo regular structure are readily crystallized and become crystalline polymers with a melting point. The melting point (Tm) can be determined by the following relation between melt enthalpy (∆H) and melt entropy (∆S): Tm =

Apparatus

Figure 11-51 Vicat softening point tester

∆H ∆S

H is an intermolecular force, which increases with hydrogen bonds. S decreases when main chains have symmetrical structure, are cross linked, or have higher rigidity. On the other hand, polymers having twisted phenylene groups become amorphous and have a glass transition temperature (Tg). Generally, below the glass transition temperature, the molecular movement is frozen with fewer changes in properties, but when the glass transition temperature is exceeded, brisk molecular movement takes place, resulting in a decrease in elastic modulus and volume expansion. Therefore, the melting point with semi-crystalline polymers and the glass transition temperature with amorphous polymers become criteria to determine heat resistance.

11.17

Brittleness Temperature (ASTM D-746)

At low temperatures, all thermoplastics tend to become rigid and brittle, mainly because the mobility of polymer chains is greatly reduced. Brittleness temperature is deﬁned as the temperature at which plastics exhibit brittle failure under impact conditions. Another way to deﬁne brittleness temperature is the temperature at which 50% of the specimens tested exhibit brittle failure under speciﬁed impact conditions.

11.17.1 Test Apparatus and Procedures The test apparatus consists of a specimen clamp and a striking member. The specimen clamp is designed so that it holds the specimen ﬁrmly as a cantilever beam. The test apparatus most commonly used is a motor driven brittleness temperature tester as shown in Figure 11-52. The tester has a rotating striking tool that rotates at a constant linear speed of 6.5 ± 0.5 ft/s. An insulated refrigerant tank with a built-in stirrer to circulate the heat transfer medium is used. The stirrer maintains the temperature equilibrium. Any type of heat transfer medium can be used as long as it remains liquid at test temperature.

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11.17 Brittleness Temperature (ASTM D-746) Motor Flywheel

Cover

Striking arm Clutch Stirring motor

Operating lever Safety devices

Adjustment hand wheel

Specimen clamp in loading jig

s en cim e Sp

Thermometer well Insulated tank

Specimens clamp (in loading position)

Figure 11-52 Brittleness temperature tester

The test specimens are usually die-punched from a sheet. The specimen are 1.0 in long, 0.25 in wide, and 0.075 in thick. Specimen conditioning is carried out according to the standard conditioning procedures. To perform the test, the specimens are securely mounted in a specimen clamp. The entire assembly is then submerged in the refrigerant. After immersion for a speciﬁed time at the test temperature, the striking tool is rotated to deliver a single impact blow to the specimens. Each specimen is carefully examined for failure. Failure is deﬁned as the division of a specimen into two or more completely separated pieces or as any crack in the specimen that is visible to the unaided eye. The temperature is raised by uniform increments of 2° or 5 °C per test and the test is repeated. This procedure is followed until both the no failure and all failure temperatures are determined. The percentage of failures at each temperature is calculated by using the number of specimens that failed. Brittleness temperature is calculated as follows: Tb = Th + ∆T [(S / 100) – (0.50)] Where: Tb = Brittleness temperature (°C) Th = Highest temperature at which all specimens fail (°C) ∆T = Temperature increment (°C) S = Sum of the percentage of breaks at each temperature Brittleness temperature has very little practical value, because the data obtained by such a method can only be used in applications in which the conditions of deformation are similar to those speciﬁed in the test. The method is, however, useful for speciﬁcation, quality control, and research and development purposes.

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11 Performance Testing of Thermoplastics

11.18

UL – Temperature Index

The service temperature of a material indicates its ability to retain a certain property, whether electrical or physical, when exposed to elevated temperatures for an extended period of time. Service temperature is therefore an important property when considering the end use applications of a thermoplastic. The use of thermoplastic materials in electrical applications such as appliances, portable electrically operated tools, and electronic encapsulation require withstanding mechanical abuse and high temperatures. Serious personal injury, electric shock, or ﬁre may occur if the product does not perform in its intended function. Realistically, thermoplastic components must retain their electrical properties and physical integrity after many hours of exposure to elevated temperatures. The service temperature, or relative thermal index of a thermoplastic, is critical to its proper selection. There are generally three service temperatures that are used to characterize the properties of thermoplastic materials: electrical, mechanical with impact, and mechanical without impact. The electrical service temperature is assigned based on destructive testing of the thermoplastic material using a dielectric strength test. The mechanical with impact service temperature is assigned based on the test results of monitoring the degradation of the tensile impact or Izod impact tests. Lastly, the mechanical without impact service temperature is assigned based on the test results of the tensile strength or ﬂexural strength tests. Many techniques are available for estimating the thermal life expectancy of thermoplastics. The method used by Underwriter’s Laboratories is outlined in IEEE Standard 101-1972. Underwriter’s Laboratories, an independent, nonproﬁt organization concerned with consumer safety, has developed a temperature index to assist UL engineers in judging the acceptability of individual thermoplastics in speciﬁc applications involving long-term exposure to elevated temperatures. The UL temperature index correlates numerically with the temperature rating or maximum temperature in °C above which a material may degrade prematurely and therefore be unsafe.

11.18.1 Relative Thermal Indices The relative thermal index of a polymeric material indicates the material’s ability to retain a particular property (physical, electrical, etc.) when exposed to elevated temperatures for an extended period of time. It is a measure of the material’s thermal endurance. For each material, a number of relative thermal indices can be established, each index related to a speciﬁc thickness of the material. The relative index of a material is determined by comparing the thermal aging characteristic of one material that has a proven ﬁeld service at a particular temperature level with the thermal aging characteristics of another material with no ﬁeld service history. A great deal of consideration is given to the properties that are evaluated to determine the relative thermal index. For the relative thermal index to be valid, the properties being stressed in the end product must be included in the thermal aging program. If, for any reason, the speciﬁc property under stress in the end product is not part of the long-term aging program, the relative thermal index may not be applicable to the use of the material in that particular application.

11.18 UL – Temperature Index The service temperatures are based on an aging study, from which the test performance of the material at lower temperatures is predicted, based on results at higher temperatures. A minimum of 5,000 h of thermal aging is necessary before a service temperature can be assigned. The ﬁnal temperature ratings that result from these investigations are critically dependent on the control material selected, specimen thickness, and type of property being evaluated. Test specimens should be in a stressed position to ensure maximum deterioration. 11.18.1.1 Control Material The service temperature depends on the comparison of the thermal aging characteristics of one material, which has a proven ﬁeld service history at a given temperature level, with those of another material with no ﬁeld service history. Therefore, one of the most important steps in the analysis is to select a suitable control material that is as similar as possible to the new/candidate material. The control material should already have been assigned a service temperature as a result of the same procedure. Any reformulation of a thermoplastic should require service temperature requaliﬁcation. Because only a small quantity of thermoplastic resin is used in its raw form, even small changes in the amount of ﬂame retardants, molding process additives, and reinforcements can create major changes in property characteristics. 11.18.1.2 Selection of Aging Temperatures A minimum of four aging temperatures should be selected for the thermal aging analysis. The temperatures may be different for each of the three properties under investigation, namely the dielectric, impact, and mechanical strengths. It may be useful to review the aging data of the control material to estimate the appropriate oven conditioning temperatures for the candidate material. 11.18.1.3 Test Procedure For each property and thickness being evaluated under the test, one set of at least ﬁve test specimens is subjected to the tests to establish the starting value, or 100% property retention value. For each oven aging temperature, ﬁve sets of test specimens are placed in the air circulating ovens. At the end of the ﬁrst, second, and third cycle, an additional set is added. The specimens are conditioned for a speciﬁed test cycle, with the highest temperature being assigned a test cycle of 3 days. The second highest oven temperature is assigned a test cycle of 7 days, and the third a test cycle of 14 days. The lowest test temperature is assigned a test cycle of 28 days. Some of the original specimens are removed from the oven and subjected to the applicable tests only at the end of the third cycle. If these specimens do not show the end of life value, namely 50% of original property retention, the test is to be repeated after every third cycle until 50% retention is reached. When this 50% retention point is achieved, the groups of specimens that were placed in the oven at delayed times are removed from the oven and tested. A performance analysis provides a more accurate determination of the time to 50% property retention. It is important to note that at least one additional data point should be obtained that shows less than 50% of the initial property value to conﬁrm the end of life value.

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11 Performance Testing of Thermoplastics

11.18.2 Long Term Thermal Aging Index The long term thermal aging program consists of exposing polymeric materials to heat for a predetermined time and observing the effect of thermal degradation. To carry out the test, an electrically heated mechanical convection oven is preferred; however, with some provisions, a noncirculating static oven may be employed. The speciﬁc properties to be evaluated in the thermal aging program are to be closely representative of the properties required in the end application. The most common mechanical properties include tensile strength, ﬂexural strength, and Izod impact strength. The electrical properties of concern are dielectric strength, surface or volume resistivity, arc resistance, and arc tracking. The test specimens are standard ASTM test bars, depending on the type of test. UL publication, “Polymeric Materials, Short Term Property Evaluations, UL 746A”, describes the specimen and test procedures to determine mechanical and electrical properties. To determine the relative thermal index, a control material with a record of good ﬁeld service at its rated temperature is selected. The control material of the same generic type and the same thickness as the candidate material is preferred. At least four different oven temperatures are selected. The highest temperature is selected so that it will take no more than two months to produce end of life of the material. The next two lower temperatures must produce the anticipated end of life by 3 and 6 months, respectively. The lowest temperature selected will take 9 to 12 months for the anticipated results. The end of life of a material is based on the assumption that at least a 2 : 1 factor of safety exists in the applicable physical and electrical property requirements. The end of life of a material is the time at each aging temperature, at which a property value has decreased by 50% of its unaged level. A 50% loss of properties due to thermal degradation is not expected to result in premature, unsafe failure. The use of the Arrhenius equation to represent the dependence of the life of the material on temperature is assumed as the functional basis for analyzing the life test data. The Arrhenius equation for a chemical reaction rate is given by the following analysis: K = A exp

E R×T

(11-1)

Where: K = Speciﬁc reaction rate E = Activation energy of the reaction R = Gas constant T = Absolute temperature (K) A = Frequency factor (assumed constant) An adaptation of Eq. 11-2 to represent insulation life, y, which is assumed to be inversely proportional to the chemical reaction rate: log10 (life) = log10 y = Constant +

E 2.303 × R × T

(11-2)

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11.18 UL – Temperature Index Equation 11-2 has the algebraic form: Y = a + (b × X )

(11-3)

Where: Y = log10 y X =l/T a = Constant b = E / (2.303 × R), another constant The constants a and b can be estimated by ﬁtting the experimental data as log10 y = Y versus 1 / T to Eq. 11-3. This can be done by graphing the data on semi-log paper and visually ﬁtting the best straight line through the points. It can also be done more precisely by using the method of least squares. To avoid underrating the material’s relative temperature index, UL publishes the ratings in three categories: applications involving electrical properties only, applications involving both electrical and mechanical properties, and applications involving both electrical and mechanical properties without impact resistance. UL publishes such data on its widely recognized “Yellow Card” and in its annual recognized component index. A Mylar® yellow card is shown in Figure 11-53. Since the long term heat aging resistance of thermoplastics is dependent on the thickness of the part or test specimen, Underwriters Laboratories requires that the thermal index testing be carried out over a wide range of thicknesses. Finally, it is important to understand that the UL temperature index program recognizes that the upper temperature limits of thermoplastics are dependent upon the stresses applied on the end product in use. Consequently, the temperature index of a particular thermoplastic qualiﬁes it only for those UL applications that UL has speciﬁcally approved.

Figure 11-53 UL yellow card for Mylar®

11.18.3 Creep Modulus/Creep Rupture Tests Both creep modulus and creep rupture strengths decrease signiﬁcantly as the temperature is increased. Before selecting a material for a load bearing application at an elevated temperature, carefully evaluate the published creep modulus and creep rupture strength test data. This is accomplished by studying the bar chart

774 Temperature, (˚C.)

11 Performance Testing of Thermoplastics

L CP

180

PPS PE T PE T

150

100

PB T Nylon 6/6 PET PBT Nylon 6/6 PC Modified PPO Acetal homopolymer

0 200 400 600 700 Creep modulus, x 103 (psi) Unfilled 30% Glass reinforced 20% Glass reinforced High Tg, unfilled

Figure 11-54 Creep modulus vs. UL temperature for various thermoplastic polymers

such as the one shown in Figure 11-54. In one chart, creep modulus of various thermoplastics at four different temperatures is compared. A similar comparison is made using creep rupture strength for the same set of thermoplastic materials. A study of the charts reveals that, while most of these thermoplastics have substantial rigidity and strength at room temperature, only a few retain enough of these properties at elevated temperatures to bear signiﬁcant loads.

11.19

Heat Deflection Temperature (ASTM D-648)

Heat deﬂection temperature is deﬁned as the temperature at which a standard test specimen 5.0 in long × 0.50 in wide × 0.25 in thick deﬂects 0.010 in under a determined stress of either 66 or 264 psi. The heat deﬂection temperature test, also referred to as the heat distortion temperature test, is commonly used for quality control and for screening and ranking materials for short term heat resistance. The data obtained by this method cannot be used to predict the behavior of thermoplastic materials at elevated temperatures, nor can it be used in designing a part or selecting a thermoplastic material. Heat deﬂection temperature is a single point measurement and does not indicate long-term heat resistance of thermoplastic materials. Heat distortion temperature, however, does distinguish between those materials that lose their rigidity and those materials that are able to sustain light loads over a narrow temperature range.

11.19.1 Apparatus and Test Specimens The apparatus for measuring heat deﬂection temperature consists of an enclosed oil bath ﬁtted with a heating chamber and automatic heating controls that raise the temperature of the heat transfer ﬂuid at a uniform rate. A cooling system is also incorporated to quickly cool the heat transfer medium for conducting repeated tests. The specimens are supported on steel supports, 4.00 in apart with the load applied on top of the specimen vertically and midway between the supports. The contact edges of the support and piece, by which pressure is applied, is rounded to a radius of 0.25 in. A suitable deﬂection measurement device, such as a dial indicator, is normally used. A mercury thermometer is used for measuring temperature. The unit is capable of applying a stress of either 66 or 264 psi on specimens by means of a dead weight. Figure 11-55 shows a commercially available ﬁve-station heat deﬂection temperature measuring system and a close up view of a loaded specimen, a holder, a dial indicator, and oil bath. More recently, automatic heat deﬂection temperature testers have been developed. These newer testers typically replace conventional temperature and deﬂection measuring devices with more sophisticated electronic measuring devices with digital read out systems and a chart recorder that prints out the test results. An automatic apparatus eliminates the need for continuous presence of an operator and minimizes operator related errors.

Figure 11-55 Heat deflection temperature tester

The test specimens consist of test bars 5.00 in long × 0.50 in wide × a wall thickness from 0.125 to 0.50 in. The test bars may be molded or cut from extruded sheets as long as they have smooth, ﬂat surfaces and are free from excessive sink marks or ﬂash. The specimens are conditioned employing standard conditioning procedures.

11.20 Soldering Heat Resistance

11.19.2 Test Procedure The specimen is positioned in the apparatus along with the temperature sensor, the deﬂection measuring devices, and the entire assembly is submerged into the oil bath kept at room temperature. The load is applied to a desired value (66 psi or 264 psi stress). Five minutes after applying the load, the pointer is adjusted to zero and the oil is heated at the rate of 2.0 ± 0.2 °C per minute. The temperature of the oil at which the specimen has deﬂected 0.010 in is recorded as the heat deﬂection temperature at the speciﬁed stress.

11.19.3 Test Variables and Limitations 11.19.3.1 Residual Stress Specimens containing any amount of “molded-in” stresses or a high degree of melt ﬂow orientation have a tendency to stress-relieve as the temperature is increased. The specimen warpage occurs in a downward direction because of the external load from the test. The combination of stress relaxation and deﬂection on the specimen yields a false heat deﬂection temperature. However, if the specimens are annealed before testing, the molded-in stresses can be relieved and the warpage can be reduced, causing a higher heat deﬂection temperature of the specimen. 11.19.3.2 Specimen Thickness Thicker specimens tend to exhibit a higher heat deﬂection temperature. This is because of the inherently low thermal conductivity of the thermoplastic materials. The thicker specimen requires a longer time to heat through, yielding a higher heat deﬂection temperature. 11.19.3.3 Applied Stress The higher the stress applied to the specimen, the lower the heat deﬂection temperature. The difference in heat deﬂection temperature values resulting from different stress varies, depending on the type of polymer. 11.19.3.4 Specimen Molding Injection molded specimens tend to have a lower heat deﬂection temperature than compression molded specimens, because the compression molded specimens are molded relatively stress-free.

11.20

Soldering Heat Resistance

Packaging of electric/electronic parts applied to printed circuit boards was formerly carried out by directly immersing the thermoplastic injection molded products into a molten soldering bath; now this method is being replaced by the vapor reﬂow soldering method. The original soldering bath method reached a temperature of 500 °F, identifying the thermoplastic injection molded products that did not exhibit the required heat resistance properties to support these temperatures. The soldering terminals of the parts became loosened or the resin materials were molten during the soldering process, causing manufacturing problems.

775

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11 Performance Testing of Thermoplastics

SOLDERING TEMPERATURES, (˚F.) TYPES OF POLYMERS

430

460

500

540

580

610

640

LCP type I LCP type II PEEK 30% glass filled PPS - R-4 PPS - R-10 PAS 30% glass filled PET 30% glass filled

Figure 11-56 Soldering heat resistance temperatures for high temperature thermoplastic resins

The heat resistance soldering temperature of 500 °F can be handled by some thermoplastic materials, when there is an extremely short time for the soldering process. The larger size parts have a greater heat capacity and before the temperature reaches 500 °F, the soldering process is ﬁnished. Materials having a heat resistant soldering temperature under load of 420 °F are able to withstand the soldering bath temperature of 500 °F. However, higher soldering temperatures are used to reduce the soldering time. To secure the soldering process without leaving excess molten solder on the printed circuits, the thermoplastic materials must withstand the temperatures under load above 500 °F. Vapor reﬂow soldering is a method brought to practice as a result of miniaturization and mass production of electric/electronic parts and components. It is often used for the production of chip parts. These parts are soldered under temperature conditions exceeding 420 °F for more than one minute. The vapor reﬂow soldering method requires lower temperatures than the soldering bath method. Since the thermoplastic parts are very small, they require a longer soldering time (60 s) under load for this process, thus demanding stricter heat resistance levels for the thermoplastic materials. The resin materials having soldering heat resistance are limited to only a few. Figure 11-56 shows the results of evaluation for soldering heat resistance of various high heat resins.

11.21

Coefficient of Linear Thermal Expansion Testing

The coefﬁcient of linear thermal expansion is deﬁned as the fractional change in length or volume of a material for a unit change in temperature. The coefﬁcients of linear thermal expansion values for different thermoplastics are of considerable interest to design engineers. Thermoplastics tend to expand and contract anywhere from six to nine times more than metals. This difference in the coefﬁcient of linear thermal expansion develops internal stresses and stress concentrations in the polymer and premature failures may occur. Special expansion joints, which generally require the use of rubber gaskets to overcome the expansion of thermoplastics, are commonly used. The use of reinforcements

777

11.22 Thermal Conductivity Testing (ASTM C-177) such as ﬁber glass lowers the linear coefﬁcient of thermal expansion considerably and brings the value closer to that of metals. Two basically similar methods have been developed by ASTM to measure the coefﬁcient of linear thermal expansion (ASTM D-696) and the coefﬁcient of cubical thermal expansion of thermoplastics (ASTM D-864), respectively. The values reported are the coefﬁcients of linear thermal expansion.

11.21.1 Test Procedure The test method requires the use of a fused quartz-tube dilatometer, a device for measuring changes in length (dial gauge or LVDT), and a liquid bath to control temperature. Figure 11-57 shows a schematic conﬁguration of a quartz-tube dilatometer. LVDT transformer LVDT core

Temperature controller zone

Specimen Quartz rod Quartz tube

Low-force spring

Figure 11-57 Schematic of a quartz-tube dilatometer

The test is started by mounting a preconditioned specimen, usually between 2.0 and 5.0 in long, into the dilatometer. The dilatometer, along with the measuring device, is then placed below the liquid level of the bath. The temperature of the bath is varied as speciﬁed. The change in length is recorded. The coefﬁcient of linear thermal expansion is calculated as follows: X =

∆L × ∆T Lo × ∆ L

Where: X = Coefﬁcient of linear thermal expansion per °C ∆L = Change in specimen length due to heating or cooling Lo = Length of the specimen at room temperature ∆T = Temperature difference in °C, over which the change in the length of the specimen is measured

11.22

Thermal Conductivity Testing (ASTM C-177)

Thermal conductivity is deﬁned as the rate at which heat is transferred by conduction through a unit cross sectional area of a material with a temperature gradient perpendicular to the area. The coefﬁcient of thermal conductivity, called the “K” factor, is expressed as quantity of heat that passes through a unit cube of the specimen in a given unit of time when the difference in temperature of the two faces is 1.0 °C.

778

11 Performance Testing of Thermoplastics One of the major advantages of thermoplastic materials is their low thermal conductivity and excellent thermal insulation. The particularly low thermal conductivity of foamed thermoplastics is due to the entrapped gasses and not to the polymeric material that serves merely as an enclosure for entrapment of gasses. As the density of the foamed thermoplastic decreases, the conductivity also decreases up to a minimum value and rises again due to increased convection effects caused by a higher proportion of open cells. The quantity of heat ﬂow depends upon the thermal conductivity of the material and on the distance the heat must ﬂow. This thermal conductivity relationship is expressed as: Q=

K X

Where: Q = Quantity of heat ﬂow K = Thermal conductivity X = Distance the heat must ﬂow Closed cell structures provide the lowest thermal conductivity because of the reduced convection of gas in the cells. The primary technique for measuring thermal conductivity of insulating materials is the guarded hot plate method. The equipment used for this test is fairly complex and expensive. The guarded heaters are used to prevent the heat ﬂow in all directions except in the axial direction towards the specimens. The test is carried out by placing the specimen between the main heater and the cooling plate. The time required to stabilize the input and the output temperature is determined.

Thermocouple to measure sample surface temperature Sample test area Top heat sink Insulation

Top Sample Q = 0 Bottom sample

Sample thickness

Q /2

Main heater Q /2

Sample thickness

Bottom heat sink

Q = power to main heater Outer shroud Apparatus

Schematic of the guarded hot plate

Figure 11-58 Thermal conductivity testing configuration

11.23 Melt Flow Testing Thermal conductivity is calculated as follows: K =

Q×t A (T1 − T2 )

Where: K = Thermal conductivity (BTU/in/h/ft2/°F) Q = Time rate of heat ﬂow (BTU/h) t = Thickness of specimen (in) A = Area under test (in2) T1 = Temperature of hot surface (°F) T2 = Temperature of cold surface (°F) Figure 11-58 shows a schematic assembly of the guarded hot plate and commercially available guarded hot plate test equipment.

11.23

Melt Flow Testing

An increasing number of processors are looking into various techniques for characterizing the incoming thermoplastic resin to prevent batch-to-batch variations. To understand the phenomenon of batch-to-batch variation, one must understand the polymerization process. The polymerization mechanism can often be very complex. Basically, it involves adding a monomer or a mixture of monomers to a reaction along with several other additives, such as stabilizers, ﬂame retardants, lubricants, reinforcements, and other ingredients, depending on the type of polymerization process. Since it is difﬁcult to manufacture a mono-dispersed polymer (a polymer, in which all molecules have the same size) commercially, variations in the size and weight of molecules in a polymer are unavoidable. Such variations in the size and the weight of the polymer’s molecules are also extremely difﬁcult to control. The relative proportions of molecules of different weights within a polymer comprise its molecular weight distribution. Depending on the range of distribution (narrow or broad), the processibility and the properties of a polymer vary signiﬁcantly. Thermoplastic materials differ widely in their viscosity (ease of ﬂow) and the problem of testing such materials is made more difﬁcult by the fact that each material is available in a range of grades, each of which also has a different ﬂow behavior. To complicate matters, the ﬂow properties of thermoplastics are nonNewtonian. In most cases, thermoplastics are pseudo thermoplastic materials meaning they become less viscous (easier ﬂowing) when they are moved more quickly. Consequently, there is not a linear relationship between pressure and ﬂow. This basic nature of the polymerization process creates a need for material characterization tests that can be used as an insurance against variations in the characteristics of the polymer. There are numerous ways of characterizing a polymer. Some are very basic and simple; others are more sophisticated and complex. The ﬁve most common and widely accepted viscosity tests are:

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11 Performance Testing of Thermoplastics • Melt index or melt ﬂow rate test • Capillary rheometer test • Viscosity tests • Gel permeation chromatography test • Analytical test (TGA, TMA, DSC)

11.23.1 Moisture Content With several thermoplastic materials, the moisture level in the material fed to the processing equipment must be kept below very small values. For example, the material fed to an injection molding machine must have a moisture level below 0.02%. This is usually necessary to prevent the production of molded parts with a poor surface ﬁnish and lower impact strength. However, the moisture can act as a ﬂow promoter, meaning that if the ﬂow properties of a production material are being assessed, the sample used for ﬂow testing must have the same moisture content as the production material.

11.24

Melt Index Testing (ASTM D-1238)

The melt index test results do not provide sufﬁcient information because it is a single point test. The ﬂow rates are measured at a single shear stress and shear rate performed at one set of temperatures and geometric conditions. Since the melt index measurement takes into account the behavior of the polymer at only one point, it is quite possible for two materials with the same melt index values to behave completely differently when exposed to shear stresses that are different from the ones used during the melt index measurements. The melt index tests measure the rate of extrusion of a thermoplastic material through an oriﬁce of speciﬁc length and diameter under prescribed conditions of temperature and pressure. This test is primarily used as a means of measuring the uniformity of the ﬂow rate of the material. The reported melt index values help to distinguish between the different grades of a polymer. A high molecular weight material is more resistant to ﬂow than a low molecular weight material. However, the data obtained from this test do not necessarily correlate with the processibility of the polymer, because thermoplastic materials are rarely manufactured without incorporating additives that affect their processing characteristics, such as stability and ﬂow ability. The effect of these additives is not readily observed via the melt index test. The rheological characteristics of polymer melts depend on a number of variables. Since the values of these variables may differ substantially from those in large-scale processes, the test results may not correlate directly to processing behavior. Flow rate testing is governed by various international standards, such as ASTM Method D-1238 and ISO R1133. Such standards specify oriﬁce size, melt temperature, heat chamber size, piston tip diameter, and the method of conducting the melt index test; the objective being to obtain consistent results from different melt indexes. Two basic methods have been developed for melt index testing, Method “A” and Method “B”. Method “A” is the traditional manual testing method, whereas Method “B” uses electronic sensing of plunger

11.24 Melt Index Testing (ASTM D-1238) displacement and calculates the ﬂow data from such measurements. Once set up, Method “B” is easier to run and more accurate for routine testing.

11.24.1 Melt Flow Rate Usually referred to as melt ﬂow rate (MFR), or melt ﬂow index (MFI), this test is widely used, particularly for PE, because the test is easy to run and to understand. A heated thermoplastic material (e.g., PE) is forced through a hole of a certain size (a die) at a speciﬁed temperature with a speciﬁed weight. The amount of PE extruded in 10.0 min is called the MFR, and the results are reported as MFR (190, 2.16) = 2.3. This means that the temperature was 190 °C and a load including the piston of 2.16 kg was used. Under these conditions, 2.3 g of the thermoplastic material was extruded in 10 minutes. It is therefore important to specify in any report, or table, the test procedure used, the nature and physical form of the material fed to the cylinder, the temperature and load used, details of any material conditioning (for example, drying the resin), the procedure used (Method “A” or “B”), and any unusual behavior of the thermoplastic material noted during the test. If more thermoplastic material had been extruded in the 10 minute period, the material would be a more easily ﬂowing material. If less thermoplastic material had been extruded in the 10 minute period, the material would be a harder ﬂowing material. Alternative Settings For acetal homopolymers, a total weight including the piston of 2.16 kg (4.75 lb) may be employed and the temperature suggested in ASTM D-1238 is 190 °C (374 °F). Larger weights than 2.16 kg may be used and different temperatures may also be employed, depending on the type of polymer and the grade of material. When the MFR test for some types of PE is run with a high load of 21.6 kg (47.52 lb), it is referred to as the high load melt index or the HLMI. Factors Affecting Flow Rate Test Results • Preheat time. If the cylinder is not preheated for a speciﬁed time, there is usually some nonuniformity in temperature along the walls of the cylinder, even though the temperature indicated on the thermometer is close to the set point. This causes the ﬂow rate to vary considerably. • Moisture. Moisture in the material, especially a highly pigmented one, causes bubbles to appear in the extrusion that may not be seen with the naked eye. Frequent weighing of short cuts of the extrusion during the experiments reveals the presence of moisture. The weight of the extrusion is signiﬁcantly inﬂuenced by the presence of the moisture bubbles. • Loading the polymer. The sample resin in the cylinder must be packed properly by pushing the rod with substantial force to allow the air trapped between the resin pellets to escape. Once the piston is lowered, the cylinder is sealed off and no air can escape. This causes variations in the test results. • Volume of sample. To achieve the same response curve repeatedly, the volume of the sample in the cylinder must be kept constant. Any change in sample volume causes the heat input from the cylinder to the material to vary signiﬁcantly.

781

782

11 Performance Testing of Thermoplastics Weight Piston

Insulation bushing Heater

Insulation 6.37 inches

0.125 inches plate

Orifice 2.00 inches

Schematic plastometer

Interpretation of Test Results The melt index values obtained from the test can be interpreted in several different ways. A slight variation (up to 10% in the case of polyethylene) in the melt index value should be interpreted as an off-speciﬁcation material. The material supplier should be consulted to determine the viscosity variation of the thermoplastic material. A signiﬁcantly different melt index value than the control standard may indicate several different things. The resin may be of a different grade with different ﬂow characteristics. It could also mean that the average molecular weight or the molecular weight distribution of the material is different than the control standard and that the material may have different properties. This test is traditionally associated with the testing of polyethylene materials for quality control purposes; for example, to determine lot-to-lot consistency of resin lots or batches. It can also be used for other purposes. For example, to test other materials or to determine barrel residence times within thermoplastic processing equipment or to assess reground content within materials for molding. Melt index is an inverse measure of molecular weight, because ﬂow characteristics are inversely proportional to the molecular weight. Figure 11-59 shows a schematic and commercially available Plastometer.

11.25

Apparatus

Figure 11-59 Melt index, melt flow rate testing (plastometer)

Capillary Rheometer Melt Viscosity Testing (ASTM D-1703)

The capillary rheometer measures apparent viscosity over an entire range of shear stresses and shear rates encountered in extrusion, injection molding, and other polymer melt processing operations. A capillary rheometer is a precision instrument that provides the accuracy and reproducibility necessary for polymer characterization tests. The quality control test to qualify the incoming thermoplastic material using the capillary rheometer method is fast, simple, and accurate. This equipment allows the establishment of at least three different load and speed settings. During the test, the force required to move the plunger at each speed is detected and plotted on a strip chart recorder against time. A viscosity history is developed for each type of material. The batch-to-batch uniformity of incoming materials can be veriﬁed by conducting a test and comparing the viscosity history. A visual comparison check for smoothness of extrusion is possible to detect gross differences in material viscosity. Figure 11-60 shows a capillary rheometer cross section and commercial test unit, consisting of an electrically heated cylinder, a pressure ram, temperature controllers, timers, and interchangeable capillaries. The plunger can be moved at a constant velocity causing a constant shear rate condition. The force to move the plunger at this speed is recorded determining the shear stress. Polymer characterization to predict the behavior of the material during processing is accomplished by measuring apparent viscosities at the range of shear rates that may be encountered during processing. Rheological analysis of the molecular chain, the melt strength characteristics of the polymers, and basic research and development work are possible with this equipment.

783

11.25 Capillary Rheometer Melt Viscosity Testing (ASTM D-1703) The temperatures used in this test are those suggested for injection molding grades of those particular materials; the most useful data is obtained at the temperatures used in production. The shear stress and shear rate conditions used should also closely approximate those used in the production process so that the melt ﬂow properties are measured at several shear rates that are representative of those used at different points in the production process. Such characteristic shear rates could, for example, represent the speed of shearing at various points within an injection molding system; at a low shear point, a medium shear point, and a high shear point. High shear rate ﬂow testing is commonly carried out by forcing the molten thermoplastic material through a speciﬁed die, using a ram at a known ram speed and material temperature. The pressure opposing ﬂow or the force needed to maintain the speciﬁed ﬂow rate is measured. The ram speed is then changed successively and the new force needed to maintain this speed is measured and recorded. For each ram speed, a force is recorded. In this case, the weight of material extruded is noted. Using the melt density, the volume extruded per unit time is calculated and from this the shear rate is determined.

Heater Extrusion barrel Barrel jacket

Insulated jacket

Barrel heater

Thermocouple

Capillary ring seal

Heater clamp

Temperature override

Clamping nut

Changeable capillary Rheometer extrusion cross section

Shear stress, shear rate, and apparent melt viscosity are calculated as follows: τ=

F ×r 2 × π × R2 × L

Where: τ = Shear stress (psi) F = Force on the plunger (lb) r = Radius of capillary oriﬁce (in) R = Radius of the barrel (in) L = Length of capillary oriﬁce (in) γ=

4×Q π × r3

Apparatus

Figure 11-60 Capillary rheometer melt viscosity test configuration

Where: γ = Shear rate (s–1) Q = Flow rate (in3/s) r = Radius of capillary oriﬁce (in) Apparent melt viscosity (poise) η =

Shear Stress (τ) Shear Rate ( γ)

11.25.1 Melt Viscosity vs. Shear Rate Curves By dividing the shear stress by the apparent wall shear rate, a viscosity at a particular shear rate may be obtained. This is an apparent viscosity; in practice the word “apparent” is often omitted. As ﬂow testing is performed over a range of conditions (for example, over a range of temperatures and ram speeds), the variation of viscosity with temperature and rate of ﬂow can be easily obtained by simple calculations. On some capillary rheometers, these calculations can be performed automatically by a built-in computer and displayed in tabular form. The information obtained may be used to construct viscosity versus shear rate curves at various temperatures, as shown in Figure 11-61.

784

11 Performance Testing of Thermoplastics

Viscosity, lb-sec/in.2 (Poise)

10

1

0.1

536° F. 554° F. 0.01

572° F. 0.001

0.0001 0.01

0.1

1

10

1.000

100

10.000

100.000

-1

Shear rate, (sec. ) Figure 11-61 Nylon 6/6 melt viscosity vs. shear rate at various temperatures

11.26

Electrical Properties Testing

The advent of new high performance engineering thermoplastic materials has assisted in the development and applications of electrical/electronic components. The combination of characteristics, such as ease of fabrication, low cost, light weight, and excellent insulation properties has made thermoplastic one of the most desirable materials for electrical applications. Thermoplastics are now specified in a majority of applications requiring resistance to extreme temperatures, chemicals, moisture, and stress. In electrical/electronic applications, thermoplastic materials usually serve some selective insulation or signal transmission function, which is frequently combined with some structural purpose. The important properties and characteristics of insulating materials can be grouped into two basic categories: insulation or dielectric characteristics and thermal classiﬁcation. Thermal classiﬁcations are particularly important in today’s electronic systems, which often are required to operate at elevated temperatures. The most important insulation characteristics are dielectric strength, resistance and resistivity, dielectric constant, power factor, dissipation factor, and arc resistance. A product designer making a ﬁrst design analysis of an electrical/electronic thermoplastic product will encounter a lot of unfamiliar terminology. Data sheets describing properties, such as “UL Temperature Index”,“Dielectric Constant”, and “Dissipation Factor” will be confusing and may require interpretation. Methods of testing and rating thermoplastics for electrical products may constitute another unknown.

785

11.26 Electrical Properties Testing To assist those who are just becoming involved with designing electrical/electronic products, some practical technical information concerning the primary electrical insulation properties and the electrical test methods that are most commonly associated with each property will be provided in the following.

11.26.1 Underwriter’s Laboratories (UL) Yellow Cards For a thermoplastic material to obtain a recognized status by UL, it must pass a variety of UL tests, including the UL 94 ﬂammability test, the UL temperature test, the UL electrical test, and the UL electrical insulation system components tests. When a thermoplastic material is granted recognized component status, a yellow card is issued. Figure 11-62 shows a yellow card for Rynite®, which contains precise identiﬁcation of the material including supplier, product designation, color, and its UL 94 ﬂammability classiﬁcation at one or more thicknesses. Also included are many of the property values, such as temperature index, hot wire ignition, high current arc ignition, and arc resistance. These data also appear in the recognized component directory. Upon successful completion of the investigation and after agreement of the terms and conditions of the listing and follow-up service, UL publishes the names of the companies that have demonstrated the ability to provide a product conforming to the established requirements. The listing signiﬁes that production samples of the product have been found to comply with the requirements and that the manufacturer is authorized to use the UL’s listing mark on the listed products that comply with the requirements.

1 – Material designation 2 – Colors 3 – Wall thickness 4 – Flammability UL rating 5, 6, 7 – Temperature index 5 – Electrical 6 – Mechanical with impact 7 – Mechanical without impact 8 – Hot wire ignition (s) 9 – High amperage ignition 10 – High voltage track rate 11 – Arc resistance 12 – Comparative tracking index Figure 11-62 UL yellow card

786

11 Performance Testing of Thermoplastics

11.26.2 How to Read and Interpret the “UL Yellow Card” The UL classiﬁcation system for thermoplastic materials is the most widely used, even outside the USA, because it is the only classiﬁcation system listing thermoplastic materials. The listing is according to temperature, ﬂammability, and electrical properties. Very important is that the famous V-0, V-1, V-2, HB, 5VA and 5VB ﬂammability classiﬁcation according to UL 94 (column 4) is only one of a total of nine properties. The remaining eight properties can be just as important. Column 1 Material designation identiﬁes the resin grade Note that quite a number of different grades can be listed together. Column 2 Colors Refer to colors with “black” and “all”, meaning as pigmented, cube blends included. Column 3 Wall thickness Shows the minimum wall thickness in mm, for which a given rating was obtained. The thickness usually ranges from 0.79 up to 6.35 mm. Column 4 Flammability classiﬁcation according to UL 94 This is the best known of all UL ratings. UL 94 rates different thermoplastics according to the ease of extinguishment after the ignition ﬂame has been removed. Columns 5, 6, and 7 Relative temperature index (RTI) UL 746 B (°C) These values indicate the long-term behavior of a thermoplastic resin concerning selected properties. Three different values are given: • Electrical properties • Mechanical properties with impact • Mechanical properties without impact Column 5 Electrical properties This column shows the upper use temperature in °C related to electrical material properties. The criterion is the temperature at which after 60,000 hours (7.0 years) the most sensitive electrical property drops to 50% of its initial value. Normally, only dielectric strength is tested.

787

11.26 Electrical Properties Testing Column 6 Mechanical properties with impact This column shows the upper use temperature in °C for impact related mechanical properties. The criterion is the temperature at which after 60,000 hours (7.0 years) the most sensitive impact property drops to 50% of its initial value. Normally measured are: • Tensile impact (tested on unreinforced resins only) • Izod impact (tested on reinforced resins only) Column 7 Mechanical properties without impact This column shows the upper use temperature in °C for nonimpact related mechanical properties. The criterion is the temperature at which after 60,000 hours (7.0 years) the most sensitive nonimpact related mechanical property drops to 50% of its initial value. Normally only tensile strength is measured. Column 8 Hot wire ignition (HWI), UL 746 A (s) This indicates how easily the ignition of a thermoplastic part in contact with a heat source (not an open ﬂame) takes place. The test simulates the case where the thermoplastic part is in contact with an overheated electrical wire. A wire is wound around a test bar (length = 5.0 in, width = 0.50 in, wall thickness as shown on yellow card), then the wire is heated to 930 °C (6.7 A leading to 0.26 W/mm heat generation) recording the time (s) until the sample ignites. At least ﬁve test bars are tested. The test bars are conditioned for 40 hours, at 23 °C, 50% relative humidity. Hot wire ignition (HWI) performance is expressed as the mean number of seconds needed either to ignite standard specimens or to burn through the specimens without ignition. The specimens are wrapped with resistant wire that dissipates a speciﬁed level of electrical energy. In order to avoid an excessive level of implied precision and bias, material performances for several tests are recorded as performance level categories (PLC), based on the mean test results (rather than recording the exact numerical results), as indicated in Table 11-2. Table 11-2 Hot Wire Ignition Time and Associated PLC Rating

HWI range – ignition time (s) 120 ≤ IT

Assigned PLC 0

60 ≤ IT < 120 30 ≤ IT < 60

15 ≤ IT < 30

1 2 3

7 ≤ IT < 15

4

7

5

0 ≤ IT <

788

11 Performance Testing of Thermoplastics Column 9 High amperage arc ignition (HAI), UL 746 A (number of arcs) This test simulates an arc between two electrodes under low voltage but high current, e.g., the two connector pins of a plug. The arc is created on the surface of the thermoplastic resin sample. The test specimen is a test bar (length = 5.0 in, width = 0.50 in, wall thickness as speciﬁed on the yellow card). Two copper electrodes are placed on the sample, where an arc is developed on the surface with a short circuit current of 32.5 A (at 240 V, 60 Hz) and a power factor of 0.50. Forty complete arcs per minute are created. A minimum of three test specimens is subjected to the test. No test sample condition is speciﬁed. High current arc ignition (HAI) performance is expressed as the number of arc rupture exposures (standardized as to electrode type and shape and electrical circuit) which are necessary to ignite the material when they are applied at a standard rate, either on the surface of the material or at a speciﬁed distance from it. Table 11-3 High Amperage Arc Ignition and Asociated PLC Rating

HAI range – number of arcs to cause ignition (NA)

Assigned PLC on UL card

120 ≤ NA

0

60 ≤ NA < 120

1

30 ≤ NA < 60

2

0 ≤ NA < 15

4

15 ≤ NA < 30

3

Column 10 High voltage tracking rate (HVTR), UL 746 A This test is designed to determine the ability of a material to withstand repeated high-voltage/low-current arcing at its surface without forming a conductive path. Two electrodes are attached to the test specimen (length = 5.0 in, width = 0.50 in, wall thickness normally 0.125 in or as shown on the yellow card) at a distance of 0.157 in apart and 5,200 V at 60 Hz are applied to the electrodes. When an arc occurs (maximum current = 2.36 mA), the electrodes are separated until the arc extinguishes. Then the electrodes are again moved closer together until the arc is reestablished. This procedure is carried out for a total testing time of 2.0 minutes, except if the tracking length (= distance between electrodes) is 2.0 in. Three test bars are conditioned for 40 hours at 73 °F, 50% relative humidity. The tracking rate number shown on the yellow card is the classiﬁcation according to the tracking length in mm/min. High voltage arc tracking rate (HVTR) is denoted as the rate, mm/min, at which a tracking path can be produced on the surface of the material under standardized test conditions. Note is made if ignition of the material takes place. The results of testing the nominal (0.125 in) thickness is considered representative of the material’s performance with any wall thickness.

789

11.26 Electrical Properties Testing Table 11-4 High Voltage tracking Rate and Associated PLC Rating

HVTR range – tracking rate (mm/min)

Assigned PLC

0 < TR ≤ 10

0

10 < TR ≤ 25

1

80 < TR ≤ 150

3

25 < TR ≤ 80

2

150 < TR

4

Column 11 High voltage, low current, dry arc resistance ASTM D-495/UL 746A This test simulates the creation of a conductive path on the resin surface when subjected to high voltage between two electrodes. This may happen if two high voltage conductors are separated by a thermoplastic insulator. Two electrodes are separated by a distance of 0.25 in and 15,000 V are applied to the electrodes, which will create an arc on the test sample surface (test specimen: length = 5.0 in, width = 0.50 in, wall thickness normally 0.125 in or as shown on yellow card). Table 11-5 Sequence of One Minute Current Steps in High Voltage, Low Current, Dry Arc Resistance Test

Step

Current (mA)

Time cycle

Total time (s)

1/8 10

10

0.25 s “ON”, 1.75 s “OFF”

60

1/4 10

10

0.25 s “ON”, 0.75 s “OFF”

120

1/2 10

10

0.25 s “ON”, 0.25 s “OFF”

180

10

10

Continuous

240

20

20

Continuous

300

30

30

Continuous

360

40

40

Continuous

420

Therefore, from 0.0 to 180.0 s, the arc is lit and extinguished with an increasing frequency. Failure of the part occurs • When a current occurs between the two electrodes • When the test sample ignites Arc resistance according to ASTM D-495 is expressed as the number of seconds that a material resists the formation of a surface conducting path when subjected to an intermittently occurring arc of high voltage, low current characteristics. The results of testing the nominal 0.125 in wall thickness is considered representative of the material’s performance with any wall thickness.

790

11 Performance Testing of Thermoplastics Table 11-6 Arc Resistance and Associated PLC Rating

Time of arc resistance (s)

Assigned PLC

420 ≤ TAR

0

360 ≤ TAR < 420

300 ≤ TAR < 360

240 ≤ TAR < 300

180 ≤ TAR < 240

120 ≤ TAR < 180 60 ≤ TAR < 120 0 ≤ TAR < 60

1 2 3 4 5 6 7

Column 12 Comparative track index (CTI) ASTM D-3638 /UL 746 A (V) This test simulates a current developing due to a surface contamination of the thermoplastic insulator between two conductors. Two electrodes are placed at a distance of 0.157 in apart on a test specimen (length = 5.0 in, width = 0.50 in, wall thickness normally 0.125 in, or as shown on the yellow card). The following test procedure is run on ﬁve (5) specimens each: • A given voltage is applied to the two electrodes. • A 0.1% ammonium chloride solution is dropped midway between both electrodes, at a rate of 2 drops/min until tracking occurs. • The tracking is deﬁned by a current increase from almost 0 A to 1 A, together with a voltage decrease. • The average number of drops per voltage relation is used for this analysis. The selection of the voltage should be done in a way that at least two test values (= voltages at which tracking occurs) need less than 50 drops of the solution, and two test values need more than 50 drops. A curve is then drawn based on at least four test values. The shape of the curve reﬂects the UL experience with the (number of drops per voltage) relationship of thermoplastics. The CTI value is deﬁned as the voltage that leads to tracking at 50 drops. Comparative tracking index (CTI) is deﬁned as the voltage that causes tracking on a material, after 50 drops of 0.1% ammonium chloride solution have fallen. Results of the 0.125 in thick test specimen is considered representative of the material’s performance with any wall thickness. Table 11-7 Comparative Tracking Index and Associated PLC Rating

CTI range – tracking index (V) 600 ≤ TI

400 ≤ TI < 600

250 ≤ TI < 400

175 ≤ TI < 250

100 ≤ TI < 175 0 ≤ TI < 100

Assigned PLC 0 1 2 3 4 5

791

11.26 Electrical Properties Testing

11.26.3 “UL Insulation Systems Recognition” Underwriter’s Laboratories Recognition is a fundamental requirement in electrical/electronics applications. There are no shortcuts for getting UL recognition. For electrical insulation systems, UL recognition requires extensive testing (UL Standard 1446) that includes heat aging, cold shock, vibration stress, moisture exposure, dielectric, and other tests. Components are evaluated according to IEC Publication 85, to meet US and most European standards. The testing is an important market requirement, but it can take 12 to 18 months to complete the tests to obtain UL recognition. Because of this time requirement, major companies pretested a “Total Package” electrical insulation system that meets UL requirements. The idea of UL recognized electrical insulation systems is not new. Companies have been getting systems listed for years. But most companies keep their systems private, for their use only, in their own proprietary product lines. A “Total Package” insulation system includes more than just a ground insulation, or coil form. It also includes magnetic wire and typically dip varnish, tape, and other major and minor elements of construction. Today, there are several hundred “Total Package” system listings available, ranging from UL Class “B” (130 °C) to “R” (220 °C), in open and closed systems. Figure 11-63 shows a typical “UL Component – System Components – Electrical Insulation Recognition Card” for Rynite®.

Replaces E69939F dated July 30, 1996. Underwriters Laboratories Inc. 324299182 N3162

(Cont. on G card) D1I/0165692

331 OBJS2 January 12, 1995 Component - System Components, Electrical Insulation

E I DUPONT DE NEMOURS & CO INC

E69939 (M) G - Continue. from F card)

Rynite FR530, 530L, RE9005

R400-N2

200(N)

600 v

Rynite FR530, 530, FR543

R402G, R402G1

200(N)

600 v

Rynite FR530, 530, FR330

R405REM, R407REM

200(N)

600 v

Rynite 530, 530L, FR530, FR530L, FR330

R408CR

200(N)

600 v

Reports: April 17, 1986; July 5, 1989; May 23, 1991; March 4,1992

Figure 11-63 “UL component – system components, electrical insulation recognition card” for Rynite® (Courtesy: Du Pont)

792

11 Performance Testing of Thermoplastics

11.27

Electrical Insulation Properties

When using thermoplastic injection molded components as parts of electric/ electronics equipment, electrical insulation properties play an important role and almost all thermoplastics have a high degree of insulation properties. Electrical insulation properties, as the words imply, are those properties that represent the degree of not conducting electricity, generally expressed in two ways: with volume resistivity and surface resistivity. Volume resistivity and surface resistivity refer to the insulation properties through the thickness and the surface, respectively. Thermoplastics have volume resistivity exceeding 1014 Ohm-cm. With a higher water absorption rate, electrical insulation properties of the materials decline. Depending on reinforcements, stabilizers, and other additives, the insulation properties change. Engineering thermoplastics are frequently specified for a wide variety of applications requiring electrical insulation. While each of the applications has its own set of requirements, hazards regarding electrical shock and ﬁre must be considered. Safety concerns are addressed by regulatory agencies through safety standards that test the performance and check suitability of the material selected. In trying to understand the behavior of the different materials as electrical insulators and their resistance to ignition, engineers need to understand what electrical properties are and what their signiﬁcance is.

11.28

Electrical Resistance Properties

The primary function of an insulator is to insulate current-carrying conductors from each other as well as from ground and to provide mechanical support for components. Naturally, the most desirable characteristic of an insulator is its ability to resist the leakage of electrical current. The higher the insulation resistance, the better the insulator. Failure to recognize the importance of insulation resistance values while designing products, such as appliances and power tools, could lead to ﬁre, electrical shock, and personal injury. The insulating materials are used to isolate the components of an electrical system from each other and from the ground. Because insulation resistance or conductance combines both volume and surface resistance or conductance, its measured value is most useful when the test specimen and electrodes have the same form as the one required in actual use. Surface resistance or conductance changes rapidly with humidity, but volume resistance or conductance changes slowly, although the ﬁnal change may eventually be greater. Along with the usual environmental variables, the dielectric resistance or conductance depends on the time of electriﬁcation and on the value of applied voltage. These parameters must be known to make the measured value of resistance or conductance meaningful. Thermoplastics not only act as effective insulators but also provide mechanical support for ﬁeld carrying conductors. For this very reason, the mechanical properties of thermoplastic materials used as insulators become very important. Typical electrical applications of thermoplastic materials include thermoplastic

793

11.28 Electrical Resistance Properties coated wires, terminals, connectors, industrial and household plugs, switches, and printed circuit boards. The following are the typical requirements of an insulator: • An insulator must have a high enough dielectric strength to withstand an electrical ﬁeld between the conductors. • An insulator must possess good arc resistance to prevent damage in case of arcing. • An insulator must maintain integrity under a wide variety of environmental conditions, such as humidity, temperature, and UV. • Insulating materials must be mechanically strong enough to resist vibration shocks and mechanical forces. • An insulator must have high insulation resistance to prevent leakage of current across the conductors.

11.28.1 Volume Resistivity Testing (ASTM D-257) Volume resistivity is the internal resistance of an insulating material at a given temperature to current ﬂow that is distributed through the volume of the specimen, expressed in Ohm-cm. Volume resistivity is related to the nature of the insulator, moisture in the material, and temperature. Volume resistance is the ratio of the direct voltage applied to two electrodes that are in contact with, or embedded in, a specimen to that portion of the current between them that is distributed through the volume of the specimen. For thermoplastics, the term “Volume Resistivity” is the resistance to current leakage through the body of an insulator. Volume resistivity or conductivity can be used as an aid in designing an insulator for a speciﬁc application. The change in resistivity or conductivity with temperature and humidity may be signiﬁcant and must be known when designing for operating conditions. Volume resistivity or conductivity determinations are often used in checking the uniformity of an insulating material, either concerning processing or to detect the conductive impurities that affect the quality of the material and that may not be readily detectable by other methods. Volume resistivity above 1019 Ohms-cm obtained for specimens under usual laboratory conditions is of doubtful validity, considering the limitations of commonly used measuring equipment. Surface resistance or conductance cannot be measured accurately, but only approximated, because approximate volume resistance or conductance is nearly always involved in the measurement. The measured value is largely a property of the contamination that happens to be on the specimen at the time. However, the dielectric constant of the specimen inﬂuences the deposition of contaminants and its surface characteristics affect the conductance of the contaminants. Figure 11-64 shows a schematic of volume resistivity test: Specimen: Electrodes: Read-off time: Applied voltage:

Plate, tape, tube Guarded or unguarded rings with metal electrodes 1.0 min 500 ± 5 V

Top electrode Power supply

Guard electrode Specimen Bottom electrode

Figure 11-64 Schematic of volume resistivity test

794

11 Performance Testing of Thermoplastics

11.28.2 Surface Resistivity Testing (ASTM D-257) The ratio of the direct voltage applied to the electrodes to the portion of the current between them that is in a thin layer of moisture or other semi-conducting material that may be deposited on the surface deﬁnes the parameters for testing. Surface resistivity is the resistance between two opposite edges of a surface ﬁlm 1.00 cm2, correctly expressed in Ohm because the centimeter terms cancel; however, the term is frequently expressed as Ohm/cm2, to avoid confusion with usual resistance values. Surface resistivity is the resistance to current leakage along the surface of an insulator, is very sensitive to humidity, surface cleanliness, and surface contour. The surface resistivity is used to check purity of an insulating material during development and its uniformity during processing. The specimens can be any practical form, such as ﬂat plates, sheets, or tubes. Measuring the surface resistivity is accomplished by using typical equipment, as shown in Figure 11-65. In measuring the resistance or conductance of insulating materials, the electrodes should be of a type of material that is readily applied, allows intimate contact with the specimen surface, and introduces no appreciable error due to electrode resistance or contamination of the specimen. The electrode should be corrosion resistant under the conditions of the test. Typically, a test report should contain the following information so that engineering decisions regarding manufacturing quality control, material acceptance, or screening can be made quicker and easier: • Description and identiﬁcation of the materials, such as name, grade, color, and manufacturer • Shape and dimensions of the test specimens • Type and dimensions of the electrodes • Conditioning of the specimens, such as cleaning, predrying, hours at humidity and temperature • Test conditions, such as specimen temperature and relative humidity at time of measurements • Method of measurement • Applied voltage • Time of electriﬁcation of measurement • Measured values of the appropriate resistance in Ohm • Computed values when required, for example, surface conductivity in siemens (per square) • Statement as to whether the reported values are apparent or steady state.

Figure 11-65 Surface resistivity test equipment

The precision and accuracy of this type of test are inherently affected by the choice of method, apparatus, and specimen, because of the variability of the resistance of a given specimen under similar test conditions and the nonuniformity of the same material from specimen to specimen. The test results are usually not reproducible to closer than 10% and are often even more widely divergent.

795

11.28 Electrical Resistance Properties Specimen: Electrodes: Read-off time: Applied voltage:

Plate, tape, tube Flat metal ring, mercury, evaporated metal 1.0 min 500 ± 5 V

11.28.3 Dielectric Strength Testing (ASTM D-149) Dielectric strength is considered the single best indicator of the insulating capabilities of a thermoplastic material. All insulating materials will fail at some level of applied voltage for a given set of operating conditions. The dielectric strength of an insulating material is deﬁned as the maximum voltage required to produce a dielectric breakdown. Failure is the passage of an arc through the test specimen. The voltage gradient is obtained by dividing the voltage at breakdown by the thickness of the insulation at the point of failure. All insulators allow a small amount of current to leak through or around themselves. Only a perfect insulator can be completely free from small current leakage. The small leakage generates heat, providing an easier access to more current. The process slowly accelerates with time and the amount of voltage applied until a failure with dielectric breakdown or a puncture occurs. Obviously, dielectric strength, which indicates electrical strength of a material as an insulator, is a very important characteristic of an insulating material. The higher the dielectric strength, the better the quality of an insulator. Figure 11-66 shows the basic dielectric strength test equipment. A variable transformer and a pair of electrodes are normally employed. Specimens of any desirable thickness prepared from the material to be tested are used. Specimen thickness of 0.062 in is fairly common. The specimen is placed between heavy cylindrical brass electrodes that carry electrical current during the test. Two dielectric strength test methods are used: Short Time Test In this method, the voltage is increased from zero to breakdown at a uniform rate. The rate of rise is generally 100, 500, 1,000, or 3,000 V/s until failure occurs. The failure is made evident by actual rupture or decomposition of the specimen. Sometimes a circuit breaker or other similar devices are employed to signal the voltage breakdown. This is not considered a positive indication of voltage breakdown, because other factors such as ﬂash over, leakage current, corona current, or equipment magnetizing current can inﬂuence such indicating devices. The initial voltage applied is 50% of breakdown voltage shown by the short time test. It is increased at rates speciﬁed for each type of material and the breakdown level noted. Breakdown by these tests means passage of sudden excessive current through the specimen and can be veriﬁed by instruments and visible damage to the specimen. Step by Step Test The step by step test method requires applying initial voltage equal to 50% of the breakdown voltage, as determined by the short time test. The voltage is then increased in equal increments and held for speciﬁed time periods until the

Figure 11-66 Dielectric strength test equipment

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11 Performance Testing of Thermoplastics specimen breaks down. In almost all cases, the dielectric strength values obtained by a step by step method correspond better with actual use conditions. However, the service failures are generally at a voltage below the rated dielectric strength because of the time factor involved. For some thermoplastics, the dielectric strength values are the same regardless of which of the two test procedures is followed, but others show a marked difference. For example, a 10% ﬁber glass reinforced polycarbonate shows a dielectric strength of 450 V/mil (for 0.125 in thick specimen) by either test method. Rigid PVC, on the other hand, registers a dielectric strength of 1,500 V/mil in short term tests, but only 375 to 750 V/mil in step by step tests (both based on 0.125 in thick specimens). The relationship of specimen thickness and dielectric strength is inverse; a ﬁlm (0.001 in) of polycarbonate, for example, might show a dielectric strength in the range of 3,500 V/mil, as compared to about 380 V/mil for a 0.125 in specimen (both at room temperature). Moreover, the dielectric strength of most thermoplastics increases with increasing temperature. Consequently, a 0.125 in thick polycarbonate specimen that has a dielectric strength of 380 V/mil at room temperature would have a dielectric strength of about 575 V/mil at 212 °F. Unfortunately, the dielectric strengths of materials vary greatly with conditions, such as humidity, specimen geometry, and the presence of contaminants in the material. Therefore, the dielectric test results are of relative rather than absolute value as speciﬁcation guides. Factors Affecting Dielectric Strength Test Results Specimen Wall Thickness: The dielectric strength of an insulator varies inversely with the fractional power (generally 0.015 in) of the thickness. The thicker specimen requires higher voltage to achieve the same voltage gradient. At higher voltages, a reduction in inter-molecular bonds is observed resulting from thermal expansion created by heat generation. Thicker sections also have internal voids, ﬂashes, moisture, nonuniformity, and greater current leakage causing early failure of the specimen. Thin ﬁlms with higher dielectric strength value have been used successfully in critical space saving applications, such as capacitors. Temperature: Dielectric strength decreases with the increase in the temperature of the specimen. If the design calls for the use of the product at various temperatures, the dielectric strength values at end use temperatures should be carefully evaluated. The dielectric strength of the material below room temperature is constant and independent of temperature change. Relative Humidity: Humidity affects the dielectric strength of the material. Surface moisture as well as moisture absorbed by hygroscopic materials affects the test results. Electrode Conﬁgurations: Dielectric strength of a material is affected by the electrode geometry, the electrode area, and the electrode material composition. Generally, the breakdown voltage decreases with increasing electrode area. The effect is more pronounced with thinner test specimens. Time Procedure: The rate of voltage application signiﬁcantly alters the test results. The dielectric strength values obtained with the step by step method are lower than those obtained by the short time test method.

11.28 Electrical Resistance Properties Mechanical Stress: Mechanical stress tends to reduce the dielectric strength values substantially. Injection Molding Process Defects: Injection molding processing defects, such as poor weld lines, gas voids, contamination, splay, and ﬂow marks tend to reduce the dielectric strength anywhere from 30 to 60%, depending on the severity of the molded defects. Test Limitations and Interpretations Dielectric strength values must be studied in detail during design stages of thermoplastic products for electrical applications. If the parts are designed based on the values derived from general published literature, without considering the effects of thickness, moisture, temperature, mechanical stress, and actual end use conditions, the product development results could be disastrous. The actual end use conditions of the insulating materials are quite different than the conditions of the dielectric strength test. Material performance in actual service end use conditions must be tested. Specimen: Specimens are thin sheets or plaques having parallel plane surfaces of a size sufﬁcient to prevent ﬂashing. This test is sometimes carried out with the sample immersed in transformer oil. Dielectric strength varies with thickness and therefore specimen thickness must be reported. Since temperature and humidity affect test results, it is necessary to condition each type of material as prescribed in the speciﬁcations for that material. The test should be run in the conditioning chamber or just after removal of the specimen from the chamber. Also, since the dielectric strength of thermoplastic varies inversely with thickness, specimen thickness must be taken into account. Procedure: Specimen: Sheets or ﬁlms Electrodes: 2.00 in diameter Rise of Voltage: – Short time 100; 200; 500; 1,000; 2,000; 5,000 V/s – Step by Step equal increments working with expected breakdown voltage Medium: Air, gas, or oil Temperature: 73 ± 9 °F

11.28.4 Dielectric Constant Testing (ASTM D-150) The dielectric constant is the most fundamental property of an insulating material. Generally, high values for the dielectric constant signify that the material is particularly well suited for use in a capacitor and low values mean the material is well suited for other electrical applications. Like other electrical properties, dielectric constant values are affected by AC frequency, temperature, and humidity. The loss factor is the product of the dielectric constant and the power factor, and is a measure of total losses in the dielectric material. The dielectric constant of an insulating material is deﬁned as the ratio of the charge stored in an insulating material placed between two metallic plates to the charge that can be stored when the insulating material is replaced by air (or vacuum). Or, the dielectric constant is deﬁned as the ratio of capacitance induced

797

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11 Performance Testing of Thermoplastics by two plates with an insulator placed between them and the capacitance of the same plates with a vacuum between them. Dielectric constant ε =

Capacitance, material as dielectric Capacitance, air or vacuum as dielectric

When voltage is applied to the insulating materials, especially in the high frequency region, thermal energy develops within the insulating materials (because power loss is converted to thermal energy). The degree of the evolving thermal energy is proportional to the product of dielectric constant ε and the dielectric loss factor (tan δ) inherent in each material. Selecting insulating materials for high frequency equipment used at higher voltage and in the region exceeding 1.0 MHz requires employing those materials with smaller ε and tan δ.

The dielectric properties, especially tan δ, of thermoplastic materials are closely related to the molecular structure of the material. The smaller the dipole moment (or polarity), the smaller the value of tan δ of the material. Representative examples with smaller tan δ include PTFE, polyphenylene ether, and others.

In many applications, insulating materials are required to perform as capacitors. Such applications are best served by thermoplastic materials having a high dielectric constant. Materials with a high dielectric constant have also helped in reducing the physical size of the capacitors. In addition, the thinner the insulating material, the higher the capacitance. Because of this fact, thermoplastic ﬁlms, such as metallized Mylar® and Kepton®, are extensively used in applications requiring high capacitance. One of the main functions of an insulator is to insulate the current carrying conductors from each other and from the ground. If the insulator is used strictly for this purpose, the capacitance of the insulating material should be as small as possible. Insulating materials are generally used in two distinct ways: to support and insulate components of an electrical network from each other and from ground, and as the dielectric of a capacitor. For the ﬁrst use, it is generally desirable to have the capacitance of the support as small as possible, consistent with acceptable mechanical, chemical, and heat resisting properties. A low value of dielectric constant (relative permittivity) is therefore desirable. For the second use, it is desirable to have a high value of dielectric constant so that the capacitor dimensions can be as small as possible. Concerning AC losses (that is dissipation factor, power factor, and so on), materials that are used to provide both insulation and capacitor dielectrics should have small losses to reduce the heating of the material and to minimize its effect on the rest of the network. In high frequency applications, a low value of loss index is particularly desirable, because, for a given value of loss index, the dielectric loss increases directly with frequency. The comprehensive test method for determining dielectric constant, dissipation factor, loss index, power factor, phase angle, and loss angle of solid electrical insulating materials developed and published by ASTM Committee D-9 is ASTM D-150. This standard test method is very popular among manufacturers and end users of thermoplastic materials. Dielectric constant generally increases with temperature, humidity, exposure to weather, and deterioration. For most materials, dielectric constant varies considerably with frequency and less with voltage as a result of polarization.

11.28 Electrical Resistance Properties Factors Affecting the Dielectric Constant Frequency: The changes in dielectric constant and loss index with frequency are produced by the dielectric polarization within the material. The two most important are dipole polarization due to polar molecules and inter-facial polarization caused by inhomogeneity in the materials. Dielectric constant and loss index vary with frequency. Starting at the highest frequency at which the dielectric constant is determined by electronic polarization, each succeeding polarization, either dipole or interfacial, contributes to the dielectric constant and the result is that the dielectric constant has its maximum value at zero frequency. Each polarization furnishes a maximum of both loss index and dissipation factor. The frequency at which the loss index is a maximum is called the relaxation frequency for that polarization. It is also the frequency at which the dielectric constant is increasing at the greatest rate and at which half its change for that polarization has occurred. Knowing the effects of these polarizations is often helpful in determining the frequencies at which measurements should be made. Any DC conductance in the dielectric caused by free ions or electrons, while affecting the dielectric constant, will also produce a dissipation factor that varies inversely with frequency and becomes inﬁnite at zero frequency. Temperature: The major electrical effect of increased temperature on an insulating material is an increase in the relaxation frequencies of its polarization. The temperature coefﬁcient of the dielectric constant at lower frequencies would always be positive. The temperature coefﬁcient will be negative at high frequencies, zero at some intermediate frequencies, and positive as the relaxation frequency of the dipole or inter-facial polarization is approached. The temperature coefﬁcient of loss index and dissipation factor may be either positive or negative, depending on the relationship of the measurement to the relaxation frequency. It will be positive for frequencies higher than the relaxation frequency and negative for lower frequencies. Voltage: All dielectric polarizations, except inter-facial, are nearly independent of the existing potential gradient until such a value is reached that ionization occurs in voids in the material or on its surface, or breakdown occurs. Humidity: The major electrical effect of elevated humidity on an insulating material is to greatly increase the magnitude of its inter-facial polarization, increasing conductance. These humidity effects are caused by the absorption of water into the volume of the material and by the formation of an ionized water ﬁlm on its surface. The water ﬁlm forms in a matter of minutes, while the increased conductance may require days and sometimes months to attain equilibrium, particularly for thick and relatively impervious materials. Weathering: By varying temperatures and humidity, falling rain, severe winds, impurities in the atmosphere, the ultraviolet light, and heat of the sun, the surface of an insulating material may be permanently changed, either physically, by roughening and cracking, and/or chemically, by the loss of relatively soluble components and by the reactions of the salts, acids, and other impurities deposited on the surface. Any water ﬁlm formed on the surface will be thicker and more conductive and water will penetrate more easily into the volume of the material. When adequate correlating data are available, the dissipation factor or power factor can be used to indicate the characteristics of a material in other respects,

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11 Performance Testing of Thermoplastics such as dielectric breakdown, moisture content, and deterioration from any cause. However, deterioration due to thermal aging may not affect the dissipation factor unless the material is subsequently exposed to moisture. Although the initial value of the dissipation factor is important, the change in dissipation factor with aging may be much more signiﬁcant. Specimen: The specimen may be a sheet of any size with uniform thickness. The test should be run at standard room temperatures and humidity, or with special sets of conditions as required. In either case, preconditioning of the specimens is needed. Procedure Metallic plates Power supply Insulating specimen

Figure 11-67 Schematic of dielectric constant test

t

The dielectric constant test is fairly simple. The electrodes are applied to opposite faces of the test specimen, as shown in Figure 11-67. The capacitance and dielectric loss are then measured by comparison or substitution methods in an electric bridge circuit. From these measurements and the dimensions of the specimen, dielectric constant and loss factor are computed. The dielectric constant value is determined from the ratio of the two measurements. Dielectric constant values are affected by factors such as frequency, voltage, temperature, moisture, and weathering. The test methods cover the determination of relative permittivity, dissipation factor, loss index, power factor, phase angle, and loss angle of the specimens of solid electrical insulating materials. The frequency range that can be covered extends from less than 1.0 Hz to several hundred MHz. Specimen: Dimension: Electrodes:

Sheet or disc can be used 0.062 in thickness ≥ Guarded or unguarded plate or cylinder, metal foils, silver, mercury, sprayed, or evaporated metal Temperature: 73 °F Voltage and frequency: Volts not deﬁned; 1.0 Hz to 108 Hz

11.28.5 Dissipation Factor Testing (ASTM D-150) In all electrical applications, it is desirable to keep the electrical losses to a minimum. Electrical losses indicate the inefﬁciency of an insulator. The dissipation factor is a measure of such electrical inefﬁciency of the insulating material. The dissipation factor indicates the amount of energy dissipated by the insulating material when the voltage is applied to the circuit. The dissipation factor is deﬁned as the ratio of the conductance of a capacitor in which the material is the dielectric to its susceptance or the ratio of its parallel reactance to its parallel resistance. Most thermoplastics have a relatively low dissipation factor at room temperature. However, at high temperatures, the dissipation factor is quite high, resulting in greater overall inefﬁciency in electrical systems. The loss factor, the product of dielectric constant and the dissipation factor, is a frequently used term, which relates to the total loss of power occurring in insulating materials. When an alternating power source is applied to an ideal dielectric material, current will ﬂow so that it is 90° out of phase with the voltage. Real insulators result in the current leading the voltage by something less than 90°. The

11.28 Electrical Resistance Properties dissipation factor is the tangent of this small loss angle. Lower values correspond to ideal dielectric materials. Conditions that affect dissipation factors are increases in temperature, frequency, humidity, and voltage. 11.28.5.1 Definition The dissipation factor or tan δ is the tangent of the loss angle. The loss angle for an insulator is the angular change in the current (I) – voltage (V) relation induced by the insulator in a capacitor versus an ideal capacitor. The dissipation factor is a dimensionless number used to calculate power losses in an insulator.

11.28.6 Arc Resistance Testing (ASTM D-495) Arc resistance is the ability of a thermoplastic material to resist the action of a high voltage, low current arc (create a leakage or fault path) under carefully controlled laboratory conditions. Arc resistance is usually stated with time required to make material electrically conductive. Arc resistance is measured in seconds, the longer the time to form a conducting path the better the arc resistance of the material. Resistance to arcing or tracking depends on the type of thermoplastic materials. Thermoset phenolics tend to carbonize easily and therefore have relatively poor arc resistance. Plastics, such as alkyds, melamine, and ﬂuorocarbons, are excellent arc resistant materials. Arc resistance of thermoplastics can be improved substantially by the addition of reinforcements, such as ﬁber glass, minerals, wood ﬂour, and other inorganic ﬁllers. Even with materials having a higher dielectric breakdown voltage, when corona discharge (arc) takes place on the surface, deterioration could be accelerated, which causes the arc resistance to become an important factor in selecting insulating materials used at higher voltages. Many points remain unclear in the deterioration mechanism of thermoplastics due to the corona discharge, e.g., its relation to the molecular structure of the material has not been explained. Failure of the test specimen may be caused by heating to incandescence, burning, carbonization of the surface, or tracking (the formation of a thin wiry line between the electrodes). In all applications, in which conducting elements are brought into contact, arcing is inevitable. Switches, circuit breakers, and automotive distributor caps are a few good examples of applications where arcing is known to cause failure. Tracking is accelerated by the presence of surface contaminants, such as dirt, oil, and moisture. This method is not used for material specifications and does not permit conclusions to be drawn concerning the relative arc resistance ranking of materials that may be subjected to other types of arcs, such as low voltage arcs at low or high currents (caused by surges or by conducting contaminants). Because of its convenience and the short time required for testing, the dry arc resistance test is intended for the preliminary screening of materials.

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11 Performance Testing of Thermoplastics The high voltage, low current dry arc resistance test is intended to simulate such service conditions as those existing in AC circuits operating at high voltage and currents limited to tens of milliamperes. To distinguish more easily between materials that have low arc resistance, the early stages of the test are mild, while later stages are successively more severe. The arc occurs intermittently between two electrodes resting on the surface of the specimen, in regular or inverted orientation. The severity is increased in the early stages by successively decreasing the time interval between ﬂashes of uniform duration to zero, and in later stages by increasing the current. The arc resistance of a material is determined by this method by measuring the total elapsed time of operation of the test until failure occurs. Four general types of failure have been observed: • Many inorganic dielectrics become incandescent, at which point they are capable of conducting the current. However, when cooled, they return to their earlier insulating condition • Some organic compounds burst into ﬂame without the formation of a visible conducting path in the substance • Some organic compounds fail by tracking, that is, a thin wiry line is formed between the electrodes • Some compounds experience carbonization of the surface until sufﬁcient carbon is present to carry the current ASTM D-495, which tests high voltage, low current, dry arc resistance of solid electrical insulation, has been the most widely used and accepted test. This