241 39 34MB
English Pages 849 [850] Year 2023
Koichi Hashiguchi
Foundations of Elastoplasticity: Subloading Surface Model Fourth Edition
Foundations of Elastoplasticity: Subloading Surface Model
Koichi Hashiguchi
Foundations of Elastoplasticity: Subloading Surface Model Fourth Edition
123
Koichi Hashiguchi MSC Software Japan Ltd. Shinjuku-Ku, Tokyo, Japan
ISBN 978-3-030-93137-7 ISBN 978-3-030-93138-4 https://doi.org/10.1007/978-3-030-93138-4
(eBook)
1st–3rd editions: © Springer International Publishing AG 2009, 2014, 2017 4th edition: © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 This work is subject to copyright. All rights are solely and exclusively licensed by the Publisher, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publisher, the authors, and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publisher nor the authors or the editors give a warranty, expressed or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publisher remains neutral with regard to jurisdictional claims in published maps and institutional affiliations. This Springer imprint is published by the registered company Springer Nature Switzerland AG The registered company address is: Gewerbestrasse 11, 6330 Cham, Switzerland
Preface
The elastoplasticity theory has developed drastically during the last half century. Then, the elastoplastic constitutive equation was formulated so as to describe not only the monotonic but also the cyclic loading behaviors accurately. Further, the rigorous formulations of the viscoplastic constitutive equation for the rate-dependent deformation behavior, the multiplicative hyperelastic-based plastic constitutive equation for the finite deformation, the multiplicative hyperelastic-based crystal plasticity and the constitutive equation for friction have been attained. This book is the standard text book for elastoplasticity which is explained comprehensively covering the monotonic to the cyclic loading behaviors. Further, the viscoplastic constitutive equation, the constitutive equation of elastoplastic-damage, the multiplicative hyperelastic-based plastic constitutive equation, the crystal plasticity and the constitutive equation for friction are explained thoroughly. Prior to these descriptions, the comprehensive explanations on vector-tensor analysis and continuum mechanics will be provided as a foundation for elastoplasticity theory, covering the underlying physical concepts, mathematical operations and theories, e.g. convective time-derivative, integration of convective rate variable, pull-back and push-forward operations, objectivity, work-conjugacy and multiplicative decomposition. The following theories/models, which provide the fundamental framework of the constitutive equations for irreversible (plastic) deformation/sliding of solids, have been proposed by the author and thus they are explained concisely and comprehensively in this book, which can never be described by the other plasticity models, e.g. the Chaboche, the Mroz and the Dafalias models assuming the purely-elastic domain within which the stress lies always. 1. The loading criterion, i.e. the judgment whether the plastic strain rate is induced is formulated based on the rational physical background (Chap. 8 Sect. 8.5), which holds in the general material exhibiting not only the hardening but also the softening behaviors. 2. The continuity condition and the smoothness condition are formulated (Chap. 9 Sect. 9.1), which must be satisfied in constitutive equations.
v
vi
Preface
3. The subloading surface concept underling the cyclic plasticity is introduced (Chap. 9 Sect. 9.2), which insists that the plastic deformation is not induced suddenly at the moment when the stress reaches the yield surface as has been postulated in the other models but it develops continuously as the stress approaches the yield surface, renamed the normal-yield surface. Here, the approaching-degree of the stress to the normal-yield surface is represented by the ratio (called the normal-yield ratio) of the size of the subloading surface to that of the normal-yield surface, while the subloading surface passes through the current stress point and is similar to the normal-yield surface. Therefore, the smooth elastic-plastic transition leading to the continuous variation of the tangent stiffness modulus is described always. Consequently, the yield judgment whether the stress reaches the yield surface is not required in constitutive equations based on the subloading surface concept. This concept is inevitable to describe the cyclic loading behavior under a general stress amplitude. 4. The explicit constitutive equation based on the subloading surface concept is formulated (Chaps. 9, 11 and 12), by which the cyclic loading behavior under the general stress amplitude is described accurately. On the other hand, the other models, i.e. the cylindrical yield surface model (Chaboche and Ohno-Wang), the multi-surface model (Mroz) and the two(bounding)-surface model (Dafalias-Yoshida) which adopt the yield surface enclosing a purely-elastic domain inside which the stress lies always are incapable of describing the cyclic loading behavior under the stress amplitude inside the yield surface (Chap. 10). Therefore, these models other than the subloading surface model are inapplicable to the mechanical design of solids and structures subjected to the general cyclic loading. Here, it should be noted that there exists the limitation in the downsizing of the elastic domain in the multi-surface and the two-surface models, because the outward-normal direction of the yield surface fluctuates busily by the downsizing of the yield surface, so that the direction of the plastic strain rate based on the normality-rule becomes unstable. Worst of all, the yield surface is limited to the large Mises yield surface in the cylindrical yield surface (Chaboche) model falling within the framework of the conventional plasticity. Soon or later these models will have to disappear from the history of plasticity. 5. The subloading-overstress model is formulated (Chap. 14), which is capable of describing the elastoplastic deformation in the quasi-static loading and the viscoplastic deformation in the general rate ranging from the quasi-static to the impact loading during the monotonic and the cyclic loading with a general stress amplitude. On the other hand, the past overstress (Bingham-Prager-Perzyna) model is incapable of describing the cyclic loading behavior induced by the variation of stress inside the yield surface and the loading behavior at high rate, predicting the elastic response under the impact loading. 6. The tangential-inelastic strain rate induced by the rate of stress tangential to the subloading surface was formulated (Chap. 9 Sect. 9.7 and Chap. 11 Sect. 11.10), which is inevitable in the analysis of the non-proportional deformation resulting in the instability and localization. On the other hand, the continuity
Preface
vii
condition is violated if the tangential-inelastic strain rate is incorporated into the other models because they assume the purely-elastic domain. 7. The constitutive equations of various materials, i.e. metals, soils, polymers, etc. are formulated (Chaps. 12, 13 and 18) by virtue of the high generality of the subloading surface model. In contrast, the Chaboche (Ohno) model published and praised mainly in Int. J. Plasticity for the last nearly half a century is limited to the description of only Mises metal without the inherent (e.g. orthotropic) anisotropy and the pressure-dependence, etc. and thus it is the typical ad hoc model only for the limited particular deformation and material, i.e. metals. 8. The subloading-damage model is formulated (Chap. 15), while the formulation by the subloading surface model is of crucial importance because the damage progresses by the slight plastic deformation induced by the variation of stress inside the yield surface. 9. The complete multiplicative-subloading hyperelastic-based (visco)plasticity for the exact descriptions of the finite elastic and (visco)plastic deformations under the monotonic/cyclic loadings is formulated (Chap. 17), which cannot be formulated by the other plasticity models. 10. The subloading-multiplicative hyperelastic-based crystal (visco)plasticity model is formulated (Chap. 21), by which the finite elastic and visco(plastic) deformations of crystalline solids can be described exactly. On the other hand, the other crystal plasticity models fall within the framework of the hypoelastic-based plasticity so that their formulations are of quite complicated forms, requiring the cumbersome stress integration procedure, and the elastic deformation described by them is limited to be infinitesimal. 11. The subloading-friction model is formulated (Chap. 22), which is capable of describing the smooth elastic-plastic transition, the reduction of friction-coefficient from the static to the kinetic friction, the recovery to the static friction during the cease of sliding, the saturation of the tangential contact stress with the increase of the normal contact stress and the dry and the fluid (lubricated) frictions at the general rate of sliding from the static to the impact sliding. Consequently, the subloading-overstress model is to be the only model which is capable of describing the rate-independent (elastoplastic) and rate-dependent (viscoplastic) deformation/sliding behaviors of solids under the monotonic and the cyclic loading processes in the general stress/strain amplitude and the general rate ranging from the static to the impact loading. Then, all the constitutive models for the irreversible deformation/sliding behaviors of solids will have to be unified to this model in the near future. The subloading surface model will be engraved as the governing law of the irreversible phenomena in the history of solid mechanics. However, what a lot of effort and time have been wasted repeatedly in order to be converged to the subloading surface model. The Chaboche model, the multi surface (Mroz) model, the two surface (Dafalias) model, the creep model, the two scale damage model, the cap soil model, the rate-and-state model, etc. are the typical huge wastes.
viii
Preface
The subloading surface model for metals has been implemented to the commercial FEM software “Marc” marketed from MSC Software Ltd. as the standard uploaded (ready-made) program by the name “Hashiguchi model” after the 2017 version, so that Marc users can apply these models to their deformation analyses. Further, the performance that the material parameters can be determined automatically by inputting the stress-strain curve is installed after the 2019 version, so that even the users unfamiliar to the elastoplasticity theory can receive the high benefits of the subloading surface model. These implementations have been highly supported by Dr. Motohatu Tateishi under the active encouragement by President Takehiko Kato, MSC Software Ltd., Japan. The subloading surface model for soils and the subloading-friction model will be also installed to the Marc as the standard uploaded programs soon. The computer programs of the subloading surface model for metals and soils and the subloading-friction models are given in Appendix L so as to use these models, capturing clearly the formulations and the calculation processes. The detail of the implicit numerical calculation method will not be provided, since it has been described in detail in the former monographs (Hashiguchi 2017, 2020). The author wishes to express cordial thanks to his colleagues at Kyushu University, who have discussed and collaborated with the author over a long period of time: Prof. Masami Ueno (currently Emeritus Prof., Univ. Ryukyus) in particular, and Dr. Takashi Okayasu (currently Prof. Kyushu Univ.), Dr. Seiichiro Tsutsumi (currently Associate Prof., Osaka Univ.), Dr. Toshiyuki Ozaki, Kyushu Electric Eng. Consult. Inc., Dr. Shingo Ozaki (currently Prof. Yokohama National Univ.) and Dr. Tatsuya Mase, Tokyo Electric Power Services Co., Ltd (currently Prof. Tezukayama Univ.). In addition, Emeritus Prof. Tadatsugu Tanaka, Univ. of Tokyo, Dr. Jnnji Yoshida, Yamanashi Univ. and Dr. Shogo Sannomaru, Mazuda Motor Corporation are appreciated for their valuable discussions and advices. Further, the author would like to express his sincere gratitude to Emeritus Prof. Akira Asaoka and his colleagues at Nagoya University: Prof. Masaki Nakano and Prof. Toshihiro Noda who have appreciated and used the author’s model widely in their analyses and who have offered discussion continuously on deformation of geomaterials. In addition, the author thanks Emeritus Prof. Teruo Nakai and Prof. Feng Zhang, the Nagoya Inst. Tech., for their valuable comments. Furthermore, the author is thankful to Dr. Kazuo Okamura, Dr. Noriyuki Suzuki, Dr. Ryoichi Higuchi and Dr. Masahiro Ogawa, Nippon Steel Corporation, Dr. Atsuro Iga, Dr. Masanori Oka and Mr. Takuya Anjiki, Yanmar Co. Ltd. for the collaborations on constitutive relations of metals and the numerical calculations. He is also grateful to Mr. Takehiko Kato (President) and Dr. Motoharu Tateishi, MSC Software, Ltd., Japan for the standard implementation of the Hashiguchi (subloading surface) model to the commercial FEM software Marc.
Preface
ix
The heartfelt gratitude of the author is dedicated to Prof. Yuki Yamakawa, Tohoku University, for a lot of valuable advices for the improvements of the descriptions through the intensive reading the manuscript and a close collaboration with detailed discussions on elastoplasticity theory, particularly on the finite strain theory based on the multiplicative elastoplasticity. In particular, the author expresses his sincere gratitude to emeritus Prof. Genki Yagawa, University of Tokyo, for encouraging always the author with undeserved high appreciation of research contributions, and thus the author was stimulated to the publication of this book. Finally, the author would like to acknowledge the enthusiastic supports by Dr. Thomas Ditzinger (Editorial director), Mr. Holger Schöpe (Book editorial service team), Springer, Heidelberg and Ms. Sindhu Sundararajan (Project manager at Scientific Publishing Services, Chennai) for the publication of this book. Tokyo, Japan Feb 23, 2022
Koichi Hashiguchi Technical Adviser, MSC Software Ltd. (Emeritus Professor, Kyushu University, Japan)
Contents
1
Mathematical Preliminaries: Vector and Tensor Analysis . . . . . 1.1 Conventions and Symbols . . . . . . . . . . . . . . . . . . . . . . . . 1.1.1 Summation Convention . . . . . . . . . . . . . . . . . . . 1.1.2 Kronecker’s Delta and Permutation Symbol . . . . 1.1.3 Matrix and Determinant . . . . . . . . . . . . . . . . . . 1.2 Vector . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.2.1 Definition of Vector . . . . . . . . . . . . . . . . . . . . . 1.2.2 Operations of Vectors . . . . . . . . . . . . . . . . . . . . 1.2.3 Coordinate Transformation of Vector . . . . . . . . . 1.3 Tensor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.1 Definition of Tensor . . . . . . . . . . . . . . . . . . . . . 1.3.2 Quotient Law . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.3 Notations of Tensors . . . . . . . . . . . . . . . . . . . . . 1.3.4 Orthogonal Tensor . . . . . . . . . . . . . . . . . . . . . . 1.4 Operations of Tensors . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.4.1 Notations in Tensor Operations . . . . . . . . . . . . . 1.4.2 Trace . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.4.3 Various Tensors . . . . . . . . . . . . . . . . . . . . . . . . 1.5 Eigenvalues and Eigenvectors . . . . . . . . . . . . . . . . . . . . . 1.6 Calculations of Eigenvalues and Eigenvectors . . . . . . . . . . 1.6.1 Eigenvalues . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.6.2 Eigenvectors . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.7 Eigenvalues and Eigenvectors of Skew-Symmetric Tensor . 1.8 Cayley-Hamilton Theorem . . . . . . . . . . . . . . . . . . . . . . . . 1.9 Scalar Triple Products with Invariants . . . . . . . . . . . . . . . 1.10 Positive Definite Tensor . . . . . . . . . . . . . . . . . . . . . . . . . 1.11 Polar Decomposition . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.12 Isotropic Tensor-Valued Tensor Function . . . . . . . . . . . . . 1.13 Representation of Tensor in Principal Space . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . . . . . . . . . . .
1 1 1 2 2 7 7 7 15 17 18 19 20 22 24 24 25 25 35 43 44 45 46 47 49 51 51 54 58
xi
xii
Contents
1.14 1.15 1.16 1.17 1.18
2
3
4
5
Two-Dimensional State . . . . . . . . . . . . . . . . . . . . Tensor Functions . . . . . . . . . . . . . . . . . . . . . . . . . Partial Differential Calculi . . . . . . . . . . . . . . . . . . Differentiation and Integration in Tensor Field . . . Representation in General Coordinate System . . . . 1.18.1 Primary and Reciprocal Base Vectors . . 1.18.2 Metric Tensor and Base Vector Algebra . 1.18.3 Tensor Representations . . . . . . . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
61 65 66 70 74 74 76 79
Description of Motion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1 Motion of Material Point . . . . . . . . . . . . . . . . . . . . . . . 2.2 Time-Derivatives . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3 Variations and Rates of Geometrical Elements . . . . . . . 2.3.1 Deformation Gradient and Variations of Line, Surface and Volume Elements . . . . . . . . . . . . 2.3.2 Velocity Gradient and Rates of Line, Surface and Volume Elements . . . . . . . . . . . . . . . . . . 2.4 Material-Time Derivative of Volume Integration . . . . . .
. . . .
. . . .
. . . .
. . . .
. . . .
83 83 85 86
.....
86
..... .....
89 92
. . . . . . . .
. . . . . . . .
. . . . . . . .
Description of Tensor (Rate) in Convected Coordinate System . 3.1 Reference and Current Primary and Reciprocal Base Vectors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2 Description of Deformation Gradient Tensor by Embedded Base Vectors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3 Pull-Back and Push-Forward Operations . . . . . . . . . . . . . . 3.4 Convected Time-Derivative . . . . . . . . . . . . . . . . . . . . . . . 3.4.1 General Convected Derivative . . . . . . . . . . . . . . 3.4.2 Corotational Rate . . . . . . . . . . . . . . . . . . . . . . . 3.4.3 On Adoption of Convected Rate Tensor in Hypoelastic Constitutive Equation . . . . . . . . . 3.4.4 Time-Integration of Convected Rate Tensor . . . .
...
95
...
95
. . . . .
. . . . .
. 97 . 99 . 104 . 105 . 106
. . . 108 . . . 109
. . . . 115 . . . . 115 . . . . 120
Deformation/Rotation Tensors . . . . . . . . . . . . . . . . . . . . . . . . 4.1 Deformation Tensors . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2 Strain Tensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3 Volumetric and Isochoric Parts of Deformation Gradient Tensor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4 Strain Rate and Spin Tensors . . . . . . . . . . . . . . . . . . . . . 4.5 Logarithmic (True) and Infinitesimal (Nominal) Strains . .
. . . . 126 . . . . 129 . . . . 141
Stress 5.1 5.2 5.3 5.4
. . . . .
Tensors and Conservation Laws . . . . . . . Stress Tensor . . . . . . . . . . . . . . . . . . . . . Conservation Law of Mass . . . . . . . . . . . Conservation Law of Linear Momentum . . Conservation Law of Angular Momentum
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
145 145 151 151 152
Contents
5.5 5.6 5.7 5.8 5.9 5.10
6
7
Equilibrium Equation . . . . . . . . . . . . . . . . . . . . Equilibrium Equation of Angular Moment . . . . . Virtual Work Principle . . . . . . . . . . . . . . . . . . . Conservation Law of Energy . . . . . . . . . . . . . . . Work Conjugacy . . . . . . . . . . . . . . . . . . . . . . . . Various Simple Deformations . . . . . . . . . . . . . . 5.10.1 Uniaxial Loading . . . . . . . . . . . . . . . . 5.10.2 Simple Shear . . . . . . . . . . . . . . . . . . . 5.10.3 Combination of Tension and Distortion
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
Objectivity and Objective (Rate) Tensors . . . . . . . . . . . . . . . . 6.1 Objectivity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Influence of Rigid-Body Rotation on Various Mechanical Quantities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3 Material-Time Derivative of Tensor . . . . . . . . . . . . . . . . 6.4 Objectivity of Convective Time-Derivative and Corotational Rate . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5 Various Objective Stress Rate Tensors . . . . . . . . . . . . . . 6.6 Jaumann Rate with Plastic Spin . . . . . . . . . . . . . . . . . . . 6.7 Time-Derivative of Scalar-Valued Tensor Function . . . . . Elastic 7.1 7.2 7.3
7.4 7.5 7.6 7.7 8
xiii
Constitutive Equations . . . . . . . . . . . . . . . Definition of Hyperelasticity . . . . . . . . . . . Hyperelastic Equations . . . . . . . . . . . . . . . Explicit Hyperelastic Models . . . . . . . . . . . 7.3.1 St.Venant-Kirchhoff Model . . . . . 7.3.2 Neo-Hookean Model . . . . . . . . . . 7.3.3 Mooney Model . . . . . . . . . . . . . . 7.3.4 Ogden Model . . . . . . . . . . . . . . . Rate Forms of Hyperelastic Equation . . . . . Infinitesimal Strain-Based Elastic Equation . Cauchy Elasticity . . . . . . . . . . . . . . . . . . . Hypoelasticity . . . . . . . . . . . . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
Elastoplastic Constitutive Equations . . . . . . . . . . . . . . . . . . . . 8.1 Fundamental Requirements for Elastoplastic Constitutive Equations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2 Classification of Elastoplastic Constitutive Equations . . . . 8.2.1 Infinitesimal Hyperelastic-Based Plasticity . . . . 8.2.2 Hypoelastic-Based Plasticity . . . . . . . . . . . . . . 8.2.3 Multiplicative Hyperelastic-Based Plasticity . . . 8.3 Conventional Plastic Constitutive Equation . . . . . . . . . . . 8.4 Constitutive Equation of Metals . . . . . . . . . . . . . . . . . . . 8.5 Formulation of General Loading Criterion . . . . . . . . . . . 8.6 Physical Backgrounds of Associated Flow Rule . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
152 155 156 157 157 162 162 164 167
. . . . 171 . . . . 171 . . . . 172 . . . . 176 . . . .
. . . .
. . . .
. . . .
178 180 183 183
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
. . . . . . . . . . . .
189 189 190 196 196 197 200 201 201 202 208 209
. . . . 211 . . . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
212 213 213 215 216 217 221 223 227
xiv
Contents
8.6.1
8.7
8.8 8.9 8.10 9
Positiveness of Second-Order Plastic Work Rate: Prager’s Interpretation . . . . . . . . . . . . . . . . . . . . . 8.6.2 Principle of Maximum Plastic Work . . . . . . . . . . 8.6.3 Positiveness of Work Done During Stress Cycle: Drucker’s Interpretation . . . . . . . . . . . . . . . . . . . 8.6.4 Positiveness of Second-Order Plastic Relaxation Work Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.5 Comparison of Interpretations for Associated Flow Rule . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Anisotropy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.7.1 Definition of Isotropy . . . . . . . . . . . . . . . . . . . . . 8.7.2 Elastoplastic Constitutive Equation with Kinematic Hardening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.7.3 Kinematic Hardening Rules . . . . . . . . . . . . . . . . . Plastic Spin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Physical Interpretation of Nonlinear Kinematic Hardening Rule . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Limitations of Conventional Elastoplasticity . . . . . . . . . . . .
Unconventional Elastoplasticity Model: Subloading Surface Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1 Mechanical Requirements . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.1 Continuity Condition in the Small . . . . . . . . . . . . 9.1.2 Continuity Condition in the Large: Smoothness Condition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2 Subloading Surface (Hashiguchi) Model . . . . . . . . . . . . . . . 9.3 Distinguished Advantages of Subloading Surface Model . . . 9.4 Numerical Performance of Subloading Surface Model . . . . . 9.5 On Bounding Surface Model with Radial-Mapping: Misuse of Subloading Surface Concept . . . . . . . . . . . . . . . . . . . . . 9.6 Incorporation of Kinematic Hardening . . . . . . . . . . . . . . . . 9.7 Incorporation of Tangential-Inelastic Strain Rate . . . . . . . . . 9.8 Limitation of Initial Subloading Surface Model . . . . . . . . . .
10 Classification of Plasticity Models: Critical Reviews and Assessments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1 Cyclic Loading Behavior . . . . . . . . . . . . . . . . . . . . . . . . . 10.2 Classification and Assessment of Plasticity Models . . . . . . 10.3 Plasticity Models with Elastic Domain . . . . . . . . . . . . . . . 10.3.1 Common Drawbacks in Models with ElasticDomain . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.2 Cylindrical Yield Surface (Chaboche) Model: Ad Hoc. Primitive Conventional Model Limited to Simple Metal Behavior . . . . . . . . . . . . . . . . .
. . . .
. . 227 . . 228 . . 228 . . 229 . . 232 . . 233 . . 233 . . 234 . . 237 . . 240 . . 240 . . 241 . . 243 . . 243 . . 243 . . . .
. . . .
245 247 253 256
. . . .
. . . .
259 260 262 270
. . . .
. . . .
271 272 273 276
. . . 276
. . . 277
Contents
xv
10.3.3
10.4 10.5
Multi-surface (Mroz) Model: Incapable of Describing Mechanical Ratchetting . . . . . . . 10.3.4 Two Surface (Dafalias) Model: Incapable of Describing Plastic Strain Rate in Unloading Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Extended Subloading Surface (Hashiguchi) Model: Capable of Describing General Loading Behavior . . . . . . Overall Assessment of Plasticity Models . . . . . . . . . . . .
11 Extended Subloading Surface Model . . . . . . . . . . . . . . . . . . . . 11.1 Normal-Yield and Subloading Surfaces . . . . . . . . . . . . . 11.2 Evolution Rule of Elastic-Core . . . . . . . . . . . . . . . . . . . 11.3 Plastic Strain Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.4 Stain Rate Versus Stress Rate Relations . . . . . . . . . . . . . 11.5 Calculation of Normal-Yield Ratio in Unloading Process . 11.6 Improvement of Inverse and Reloading Responses . . . . . 11.7 Loading Criterion for Large Loading Increment . . . . . . . 11.7.1 Exact Judgment of Loading . . . . . . . . . . . . . . . 11.7.2 Initial Value of Normal-Yield Ratio in Plastic Corrector Step . . . . . . . . . . . . . . . . . . . . . . . . 11.8 Plastic Spin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.9 Incorporation of Tangential-Inelastic Strain Rate . . . . . . .
. . . . 283
. . . . 284 . . . . 288 . . . . 290 . . . . . . . . .
. . . . . . . . .
. . . . . . . . .
. . . . . . . . .
293 293 296 302 303 304 305 308 309
. . . . 312 . . . . 315 . . . . 315
12 Constitutive Equations of Metals . . . . . . . . . . . . . . . . . . . . . . . 12.1 Yield Surface, Isotropic, Kinematic Hardening and Elastic-Core . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2 Cyclic Stagnation of Isotropic Hardening . . . . . . . . . . . . . 12.3 Calculation of Normal-Yield Ratio in Unloading Process . . 12.4 Implicit Stress-Integration . . . . . . . . . . . . . . . . . . . . . . . . 12.5 Material Parameters and Comparisons with Test Data . . . . 12.5.1 Material Parameters . . . . . . . . . . . . . . . . . . . . . 12.5.2 Comparisons with Test Data . . . . . . . . . . . . . . . 12.6 Analyses of Engineering Phenomena . . . . . . . . . . . . . . . . 12.7 Orthotropic Anisotropy . . . . . . . . . . . . . . . . . . . . . . . . . . 12.7.1 Representation of Isotropic Mises Yield Condition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.7.2 Plane Strain State . . . . . . . . . . . . . . . . . . . . . . . 12.8 Subloading Surface Model with Orthotropic Anisotropy . . 12.8.1 Subloading Surface with Orthotropic Anisotropy 12.8.2 Plastic Strain Rate . . . . . . . . . . . . . . . . . . . . . . 12.8.3 Normal-Yield Ratio . . . . . . . . . . . . . . . . . . . . . 12.8.4 Elastic-Core Yield Ratio . . . . . . . . . . . . . . . . . . 12.8.5 Cyclic Stagnation of Isotropic Hardening . . . . . .
. . . 319 . . . . . . . . .
. . . . . . . . .
. . . . . . . . .
319 321 326 327 329 329 331 342 346
. . . . . . . .
. . . . . . . .
. . . . . . . .
356 359 360 360 361 362 363 363
xvi
13 Constitutive Equations of Soils . . . . . . . . . . . . . . . . . . . . . . . 13.1 Isotropic Consolidation Characteristics . . . . . . . . . . . . . 13.2 Yield Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.2.1 Yield Functions . . . . . . . . . . . . . . . . . . . . . . 13.2.2 Critical State Surface Taken Account of Third Deviatoric Invariant . . . . . . . . . . . . . . . . . . . 13.3 Subloading Surface Model for Soils . . . . . . . . . . . . . . . 13.4 Extension of Material Functions . . . . . . . . . . . . . . . . . . 13.4.1 Yield Surface with Tensile Strength . . . . . . . . 13.4.2 Rotational Hardening . . . . . . . . . . . . . . . . . . 13.5 Extended Subloading Surface Model . . . . . . . . . . . . . . 13.5.1 Superyield, Normal-Yield and Subloading Surfaces . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.5.2 Evolution Rules of Internal Variables . . . . . . . 13.5.3 Plastic Strain Rate . . . . . . . . . . . . . . . . . . . . 13.5.4 Yield Stress Function . . . . . . . . . . . . . . . . . . 13.5.5 Partial Derivatives of Subloading Surface Function . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.5.6 Calculation of Normal-Yield Ratio . . . . . . . . . 13.6 Simulations of Test Results . . . . . . . . . . . . . . . . . . . . . 13.7 Numerical Analysis of Footing Settlement Problem . . . . 13.8 Hyperelastic Equation of Soils . . . . . . . . . . . . . . . . . . .
Contents
. . . .
. . . .
. . . .
. . . .
. . . .
365 365 374 374
. . . . . .
. . . . . .
. . . . . .
. . . . . .
. . . . . .
377 383 393 393 397 401
. . . .
. . . .
. . . .
. . . .
. . . .
401 405 407 409
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
409 413 419 423 430
. . . . .
. . . . .
433 433 435 438 440
. . . . .
. . . . .
442 442 445 447 449
14 Viscoplastic Constitutive Equations with Subloading Surface Concept . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.1 Rate-Dependent Deformation of Solids . . . . . . . . . . . . . . . . 14.2 History of Viscoplastic Constitutive Equations . . . . . . . . . . 14.3 Irrationality of Creep Model . . . . . . . . . . . . . . . . . . . . . . . 14.4 Mechanical Response of Past Overstress Model . . . . . . . . . 14.5 Subloading Overstress Model: Extension to Description of General Rate of Deformation . . . . . . . . . . . . . . . . . . . . . 14.5.1 Static and Limit Subloading Surfaces . . . . . . . . . . 14.5.2 Viscoplastic Strain Rate . . . . . . . . . . . . . . . . . . . 14.5.3 Strain Rate Versus Stress Rate Relation . . . . . . . . 14.6 Comparison with Test Data . . . . . . . . . . . . . . . . . . . . . . . . 14.6.1 Dynamic Loading Process Inducing Elastic-Viscoplastic Deformation . . . . . . . . . . . . . 14.6.2 Quasi-static Loading Process Inducing Elastoplastic Deformation Behaviors . . . . . . . . . . 14.7 Temperature Dependence of Elasto-Viscoplastic Deformation Behavior . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
. . 449 . . 451 . . 454
Contents
xvii
15 Continuum Damage Model with Subloading Surface Concept 15.1 Basic Hypothesis of Strain Equivalence in Constitutive Equation with Brittle Damage . . . . . . . . . . . . . . . . . . . . 15.2 Hyperelastic Equation in Undamaged Variables . . . . . . . 15.3 Hyperelastic Equations with Damage . . . . . . . . . . . . . . . 15.3.1 Bilateral Damage . . . . . . . . . . . . . . . . . . . . . . 15.3.2 Unilateral Damage . . . . . . . . . . . . . . . . . . . . . 15.4 Evolution of Damage Variable . . . . . . . . . . . . . . . . . . . . 15.4.1 Bilateral Damage . . . . . . . . . . . . . . . . . . . . . . 15.4.2 Unilateral Damage . . . . . . . . . . . . . . . . . . . . . 15.5 Elastoplastic-Damage Model with Subloading Surface Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15.5.1 Plastic Strain Rate . . . . . . . . . . . . . . . . . . . . . 15.5.2 Stress(Rate) Versus Strain(Rate) Relation and Stress Integration . . . . . . . . . . . . . . . . . . . 15.6 Anisotropic (Orthotropic) Damage Tensor . . . . . . . . . . . 15.7 Subloading-Overstress Damage Model . . . . . . . . . . . . . . 15.7.1 Bilateral Damage . . . . . . . . . . . . . . . . . . . . . . 15.7.2 Unilateral Damage . . . . . . . . . . . . . . . . . . . . . 15.8 Subloading-Gurson Model for Ductile Damage . . . . . . . . 15.9 High Cycle Fatigue: Redundancy of Two-Scale Damage Model and Necessity of Subloading-Damage Surface Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 Subloading Phase-Transformation Model . . . . . . . . . . . . . . . 16.1 Constitutive Equation . . . . . . . . . . . . . . . . . . . . . . . . . 16.1.1 Elastic Strain Increment . . . . . . . . . . . . . . . . 16.1.2 Plastic Strain Increment Based on Subloading Surface Model . . . . . . . . . . . . . . . . . . . . . . . 16.2 Thermal and Transformation Strain Increments . . . . . . . 16.2.1 Heat-Transformation Strain Increment . . . . . . 16.2.2 Transformation-Plastic Strain Increment . . . . . 16.3 Stress Rate Versus Strain Rate Relation . . . . . . . . . . . . 17 Multiplicative Hyperelastic-Based Plasticity with Subloading Surface Concept . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17.1 Exact Elastic–Plastic Decomposition of Deformation Measure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17.1.1 Necessity of Multiplicative Decomposition of Deformation Gradient Tensor . . . . . . . . . . 17.1.2 Embedded Base Vectors in Intermediate Configuration . . . . . . . . . . . . . . . . . . . . . . . .
. . . . 457 . . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
458 458 460 460 462 472 472 473
. . . . 474 . . . . 474 . . . . . .
. . . . . .
. . . . . .
. . . . . .
477 481 484 485 485 486
. . . . 491
. . . . . 493 . . . . . 493 . . . . . 494 . . . . .
. . . . .
. . . . .
. . . . .
. . . . .
494 496 496 497 498
. . . . . 501 . . . . . 502 . . . . . 502 . . . . . 504
xviii
Contents
17.2
17.3 17.4 17.5 17.6 17.7 17.8 17.9 17.10 17.11
17.12 17.13 17.14
Deformation Tensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17.2.1 Elastic and Plastic Right Cauchy-Green Deformation Tensor . . . . . . . . . . . . . . . . . . . . . . 17.2.2 Strain Rate and Spin Tensors . . . . . . . . . . . . . . . On Limitation of Hypoelastic-Based Plasticity . . . . . . . . . . Further Multiplicative Decomposition of Plastic Deformation Gradient Tensor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Formulation and Calculation in Intermediate Configuration: Isoclinic Concept . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Stress Measures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Internal Variables . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Normal-Yield, Subloading and Elastic-Core Surfaces . . . . . Plastic Flow Rules . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Plastic Strain Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Material Functions for Metals and Soils . . . . . . . . . . . . . . . 17.11.1 Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17.11.2 Soils . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Calculation Procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . Isotropic Hardening Stagnation . . . . . . . . . . . . . . . . . . . . . Subloading-Overstress Model . . . . . . . . . . . . . . . . . . . . . . . 17.14.1 Constitutive Equation . . . . . . . . . . . . . . . . . . . . . 17.14.2 Calculation Procedure . . . . . . . . . . . . . . . . . . . . .
. . 505 . . 505 . . 506 . . 509 . . 510 . . . . . . . . . . . . . .
. . . . . . . . . . . . . .
516 517 519 522 524 527 529 529 531 536 538 542 542 548
18 Viscoelastic-Viscoplastic Model of Polymers . . . . . . . . . . . . . 18.1 Viscoelastic Rheological Model . . . . . . . . . . . . . . . . . . 18.2 Viscoelastic Deformation with Elastic Strain Energy Function . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18.2.1 Elastic Strain Free-Energy Function . . . . . . . . 18.2.2 Second Piola–Kirchhoff Stress Tensor . . . . . . 18.3 Viscoelastic-Damage Model: Subloading-Mullins Effect 18.4 Viscoplastic Constitutive Equation in Glassy State . . . .
. . . . . 549 . . . . . 549 . . . . .
. . . . .
. . . . .
. . . . .
. . . . .
551 551 552 557 561
19 Corotational Rate Tensors . . . . . . . . . . 19.1 Hypoelasticity . . . . . . . . . . . . . . . 19.1.1 Zaremba-Jaumann Rate . 19.1.2 Green-Naghdi Rate . . . . 19.2 Kinematic Hardening Material . . . 19.2.1 Zaremba-Jaumann Rate . 19.2.2 Green-Naghdi Rate . . . . 19.3 Plastic Spin . . . . . . . . . . . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
565 565 566 567 569 571 571 573
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
. . . . . . . .
Contents
20 Localization of Deformation . . . . . . . . . . . . . . . . . . 20.1 Element Test . . . . . . . . . . . . . . . . . . . . . . . . . 20.2 Gradient Theory . . . . . . . . . . . . . . . . . . . . . . 20.3 Shear-Band Embedded Model: Smeared Crack 20.4 Necessary Condition for Shear Band Inception
xix
....... ....... ....... Model . . .......
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21.1 Description of Strain Rate and Spin by Crystal Lattice Vectors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21.2 Resolved Shear Stress (Rate) . . . . . . . . . . . . . . . . . . . . 21.3 Stress Rate Versus Plastic Shear Strain Rate Relation . . 21.4 Conventional Crystal Plasticity Model . . . . . . . . . . . . . 21.4.1 Yield Condition and Flow Rule . . . . . . . . . . . 21.4.2 Evolution of Isotropic Hardening . . . . . . . . . . 21.4.3 Evolution of Kinematic Hardening . . . . . . . . . 21.4.4 Stress Rate Versus Strain Rate Relation . . . . . 21.5 Subloading Crystal Plasticity Model . . . . . . . . . . . . . . . 21.6 Subloading-Overstress Crystal Plasticity Model . . . . . . . 21.7 Extension to Description of Cyclic Loading Behavior . . 21.8 Uniqueness of Slip Rate Mode . . . . . . . . . . . . . . . . . . . 21.9 Various Schemes for Calculation of Shear Strain Rates . 21.9.1 Singular Value Decomposition . . . . . . . . . . . 21.9.2 Regularized Schmid Law . . . . . . . . . . . . . . . 21.9.3 On Creep-Type Crystal Plasticity Model . . . .
. . . . .
. . . . .
. . . . .
. . . . .
. . . . .
583 583 584 588 589
. . . . . 595 . . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . .
596 601 604 605 605 606 607 609 613 619 623 629 632 633 636 637
22 Constitutive Equation for Friction: Subloading-Friction Model . 22.1 History of Constitutive Equation for Friction . . . . . . . . . . . 22.2 Sliding Displacement and Contact Traction . . . . . . . . . . . . . 22.3 Hyperelastic Sliding Behavior . . . . . . . . . . . . . . . . . . . . . . 22.4 Elastoplastic Sliding Velocity . . . . . . . . . . . . . . . . . . . . . . . 22.4.1 Sliding Normal-Yield and Subloading Surfaces . . 22.4.2 Evolution Rule of Sliding Hardening Function . . . 22.4.3 Evolution Rule of Sliding Normal-Yield Ratio . . . 22.4.4 Elastoplastic Sliding Velocity . . . . . . . . . . . . . . . 22.5 Loading Criterion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22.6 Calculation of Normal Sliding-Yield Ratio . . . . . . . . . . . . . 22.7 Fundamental Mechanical Behavior of Subloading-Friction Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22.7.1 Relation of Tangential Contact Stress Rate and Sliding Velocity . . . . . . . . . . . . . . . . . . . . . . 22.7.2 Numerical Experiments and Comparisons with Test Data . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
. . . . . . . . . . .
. . . . . . . . . . .
641 641 643 646 647 647 648 651 654 659 661
. . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . .
. . 661 . . 662 . . 664
xx
Contents
22.8 22.9
Stick–Slip Phenomenon . . . . . . . . . . . . . . . . . . . . . . . . . Friction Condition with Saturation of Tangential Contact Stress . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22.10 Subloading-Overstress Friction Model . . . . . . . . . . . . . . 22.10.1 Subloading-Overstress (Viscoelastic) Friction Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22.10.2 Numerical Experiments . . . . . . . . . . . . . . . . . . 22.10.3 Comparison with Test Data . . . . . . . . . . . . . . . 22.11 Extension to Rotational and Orthotropic Anisotropy . . . .
. . . . 669 . . . . 672 . . . . 683 . . . .
. . . .
. . . .
. . . .
684 689 692 694
Final Remarks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 703 Appendix A: Projection of Area . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 709 Appendix B: Logarithmic Spin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 711 Appendix C: Matrix Representation of Tensor Relations . . . . . . . . . . . . 715 Appendix D: Euler’s Theorem for Homogeneous Function . . . . . . . . . . . 721 Appendix E: Outward-Normal Tensor of Surface . . . . . . . . . . . . . . . . . . 723 Appendix F: Relationships of Material Constants in ln v ln p and e ln p Linear Relations . . . . . . . . . . . . . . . . . . . . . . . . 725 Appendix G: Derivative in Critical State . . . . . . . . . . . . . . . . . . . . . . . . . 727 Appendix H: Convexity of Two-Dimensional Curve . . . . . . . . . . . . . . . . . 729 Appendix I: Normal Tensor to Subloading Surface with Anisotropic Damage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 733 Appendix J: Tensor Exponential Map for Time-Integration of First-Order Linear Differential Equation . . . . . . . . . . . . 737 Appendix K: Eyring Equation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 739 Appendix L: Computer Programs of Subloading Surface Models . . . . . 741 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 817 Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 841
Chapter 1
Mathematical Preliminaries: Vector and Tensor Analysis
Physical quantities appearing in continuum mechanics are mathematically expressed by tensors because they possess not only magnitudes but also directions in multi-dimensional space. Therefore, their relations are described by tensor relations. Before the explanation on the main theme of this book, i.e. the elastoplasticity theory, the mathematical properties of tensors and the mathematical rules on tensor operations are explained on the level necessary to understand the continuum mechanics mainly related to the elastoplasticity theory. The readers will able to understand the formulations almost completely, since the comprehensive explanations and proofs for almost all of the concepts and the equations are given without any logical jumps.
1.1
Conventions and Symbols
Some basic conventions and symbols appearing in the tensor analysis are described in this section.
1.1.1
Summation Convention
We first introduce the Cartesian summation convention. Repeated suffix in a term is summed over numbers that the suffix can take. For instance, ur v r ¼
3 P
ur vr ; Tir vr ¼
r¼1
Trr ¼
3 P
Trr
3 P r¼1
9 > Tir vr > = > > ;
ð1:1Þ
r¼1
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_1
1
2
1
Mathematical Preliminaries: Vector and Tensor Analysis
where the range of suffixes is 1; 2; 3. Because of ur vr ¼ us vs ; Tir vr ¼ Tis vs ; Trr ¼ Tss a letter of the repeated suffix is arbitrary. It is therefore called as the dummy index. The convention described above is also called Einstein’s summation convention. Hereinafter, a repeated index obeys this convention unless otherwise specified by the additional remark “(no sum)”.
1.1.2
Kronecker’s Delta and Permutation Symbol
The symbol dij ði; j ¼ 1; 2; 3Þ defined in the following equation is called the Kronecker’s delta. dij ¼
1: i ¼ j 0: i ¼ 6 j
ð1:2Þ
for which one has dir drj ¼ dij ;
dii ¼ 3
ð1:3Þ
Furthermore, the symbol eijk defined by the following equation is called the alternating (or permutation) symbol or Eddington’s epsilon or Levi-Citiva “e” tensor.
eijk
8
> a a ¼ a ¼ 1 > > > ½a b c > > > > < b c a b ¼ b ¼ 0 ½a b c > > > > > > ½b c c a a b ½a b c 2 1 > > > ða bÞ c ¼ ¼ ¼ : 3 3 ½a b c ½a b c ½a b c
ð1:64Þ
noting Eq. (1.59) with Eq. (1.54). Further, it follows from Eqs. (1.52) and (1.59) that a a þ b b þ c c ¼ 0
ð1:65Þ
noting a a þ b b þ c c ¼
a ðb cÞ þ b ðc aÞ þ c ða bÞ ½abc
(5) Tensor product and component Based on the vectors vð1Þ ; vð2Þ ; ; vðmÞ , one can make the m-th order tensor as follows: ð2Þ ðmÞ vð1Þ vð2Þ vðmÞ ¼ vð1Þ vpm ep1 ep2 epm p 1 vp 2
ð1:66Þ
For two vectors, one has the second-order tensor u v ¼ ui ei vj ej ¼ ui vj ei ej
ð1:67Þ
which is expressed in the matrix form 2
u1 v1 ½u v ¼ 4 u2 v1 u3 v 1
u1 v 2 u2 v 2 u3 v 2
3 u1 v 3 u2 v 3 5 u3 v 3
ð1:68Þ
As described above, one can make a tensor from two vectors. After the scalar product u v and the vector product u v for the two vectors u and v, one calls
14
1
Mathematical Preliminaries: Vector and Tensor Analysis
u v as the tensor (cross) product or dyad which means “one set by two”. Particularly, it holds for three arbitrary vectors u, v and w that ðu vÞw ¼ ðui ei vj ej Þwk ek ¼ ui ei ðvj wk djk Þ ¼ ui ei ðvj wj Þ ¼ uðv wÞ and thus the following expression holds. u v ¼ uðv
ð1:69Þ
It follows from Eq. (1.8) that detðu vÞ ¼ eijk u1 vi u2 vj u3 vk ¼ u1 u2 u3 eijk vi vj vk ¼ u1 u2 u3 ½vvv leading to detðu vÞ ¼ 0
ð1:70Þ
Equation (1.62) is rewritten in terms of the tensor product as follows: v ¼ gv
ð1:71Þ
where g ¼ a a þ b b þ c c ¼ a a þ b b þ c c ð¼gT Þ
ð1:72Þ
g is regarded as the generalized identity tensor and is called the metric tensor transforming the vector to itself in the general coordinate system. In other words, the identity tensor in the general coordinate system is defined by incorporating the reciprocal base vectors ða ; b ; c Þ for the primary base vectors ða; b; cÞ. In particular, g is reduced in the normalized rectangular coordinate system ða ¼ a ¼ e1 ; b ¼ b ¼ e2 ; c ¼ c ¼ e3 Þ as follows: I ¼ dij ei ej ¼ ei ei ð¼IT Þ;
dij ¼ ei ej
ð1:73Þ
which possesses the components of the Kronecker’s delta dij and is called the unit tensor or the identity tensor transforming the vector to itself. Incidentally, the permutation tensor is defined by e ¼ eijk ei ej ek ;
eijk ¼ ½ei ej ek
ð1:74Þ
noting Eq. (1.48). The following relation holds. e :ðu vÞ ¼ u v
ð1:75Þ
1.2 Vector
15
noting ðeijk ei ej ek Þ:ður er vs es Þ ¼ eijk ei ur vs djr dks ¼ eijk ei uj vk ¼ uj ej vk ek
1.2.3
Coordinate Transformation of Vector
Adopt the other normalized orthogonal coordinate system fO xi g with the base fei g in addition to the normalized orthogonal coordinate system fO xi g with the base fei g (Fig. 1.1). Noting v ¼ vj ej ¼ ðv ej Þej in general, the following relations hold for the base vectors. ei ¼ ðei ej Þej ; ei ¼ Qri er ;
ei ¼ ðei ej Þej
ð1:76Þ
ei ¼ Qir er
ð1:77Þ
where the coordinate transformation operator Qij is defined by Qij cosðangle between ei and ej Þ ¼ ei ej
ð1:78Þ
Moreover, because of Qir Qjr ¼ ðei er Þðej er Þ ¼ ei ðej er Þer ¼ ei ej Qri Qrj ¼ ðer ei Þðer ej Þ ¼ ðei er Þer ej ¼ ei ej
It follows that Qir Qjr ¼ Qri Qrj ¼ dij
ð1:79Þ
It is assumed for a while that the relative (parallel and rotational) motion does not exist between the above-described coordinate systems, and that their origins mutually coincide. Then, denoting the component on the base ei by ðÞ , the x2
x*2
v
v2
v*2
e2
v*1
e*2 e*1 0
e1
x*1
θ v1
x1
Fig. 1.1 Coordinate transformation of vector illustrated in the two-dimensional state
16
1
Mathematical Preliminaries: Vector and Tensor Analysis
coordinate transformation rule, i.e. the transformation rule of the components of v in these coordinate systems is given by vi ¼ Qij vj
ð1:80Þ
noting vr er ei ¼ vj ej ei and based on v ¼ vj ej ¼ vr er
ð1:81Þ
Furthermore, noting Qri vr ¼ Qri Qrs vs ¼ dis vs , the inverse relation of Eq. (1.80) is given as vi ¼ Qji vj
ð1:82Þ
Equations (1.80) and (1.82) are expressed in matrix form as 38 9 8 9 2 Q11 Q12 Q13 > < v1 > = < v1 = 6 7 v2 ¼ 4 Q21 Q22 Q23 5 v2 ; > : ; : > ; v3 Q31 Q32 Q33 v3
8 9 2 38 9 Q11 Q21 Q31 < v = > < v1 > = 6 7 1 v2 ¼ 4 Q12 Q22 Q32 5 v2 > : ; : > ; v3 v3 Q13 Q23 Q33
ð1:83Þ
which are often expressed simply as fv g ¼ ½Qfvg;
fvg ¼ ½QT fv g
ð1:84Þ
Needless to say, an equation including ð Þ and ð Þ does not describe the relation between different vectors, but describes the relations between components when a certain vector is described by two different coordinate systems. As known from the following equation, the magnitude of vector is not influenced by the coordinate transformation, whilst it is the basic property of the scalar quantity. kv k ¼
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffi Qir vr Qis vs ¼ Qir Qis vr vs ¼ drs vr vs ¼ vr vr ¼ kvk
The relations described above are shown below.
e e ½Q ¼ 1 1 e2 e1
e1 e2 e2 e2
cos h sin h ½Q½Q ¼ sin h cos h T
e ¼ 1 1 e2 1
cos h sin h
e
1 2 ¼ cos h sin h sin h cos h e2 2
sin h 1 ¼ cos h 0
0 ¼ dij ¼ ½I 1
1.2 Vector
17
ei ¼ ei e1 e1 þ ei e2 e2 ¼ Qi1 e1 þ Qi2 e2 ; e1 ¼ cos he1 þ sin he2 e2 ¼ sin he1 þ cos he2
ei ¼ ei e1 e1 þ ei e2 e2 ¼ Q1i e1 þ Q2i e2 v ¼ v1 e1 þ v2 e2 ¼ v1 e1 þ v2 e2
v1 v2
¼
cos h sin h sin h cos h
v1 ; v2
v1 v2
¼
cos h sin h
sin h cos h
v1 v2
where h designates the angle that the base ei rotates in the anti-clock direction from the base fei g. Choosing the position vector x as the vector v, it follows from Eqs. (1.80) and (1.82) that xi ¼ Qir xr
)
xi ¼ Qri xr
ð1:85Þ
from which one has 9 @xi @Qir xr @xr > ¼ ¼ Qir ¼ Qir djr ¼ Qij > = @xj @xj @xj > @xj @Qrj xr @x ; ¼ ¼ Qrj r ¼ Qrj dir ¼ Qij > @xi @xi @xi
ð1:86Þ
Consequently, Qij can be also described as Qij ¼
1.3
@xi @xj ¼ @xj @xi
ð1:87Þ
Tensor
The vector described in the foregoing possesses the direction in first order but there exit quantities possessing the direction in high order. They are collectively called the tensor. The general definition and mathematical properties of tensor are described in this section.
18
1
1.3.1
Mathematical Preliminaries: Vector and Tensor Analysis
Definition of Tensor
Let the set of nm functions be described as Tðp1 ; p2 ; ; pm Þ in the coordinate system fO xi g with the origin O and the axes xi ði ¼ 1; 2; ;nÞ in the n-dimensional space, where each of the indices p1 ; p2 ; ; pm takes the number 1; 2; ;n. This set of functions is defined as the m th-order tensor in the n-dimension, if the set of functions is observed in the other coordinate system fOxi g with the origin O and the axes xi as follows: T ðp1 ; p2 ; pm Þ ¼ Qp1 q1 Qp2 q2 Qpm qm Tðq1 ; q2 ; qm Þ
ð1:88Þ
or T ðp1 ; p2 ; pm Þ ¼
@xp1 @xp2 @xq1 @xq2
@xpm Tðq1 ; q2 ; qm Þ @xqm
ð1:89Þ
provided that only the directions of axes are different but the origin is common and the relative motion does not exist. Here, Eq. (1.87) is used. Then, designating Tðp1 ; p2 ; pm Þ by the symbol Tp1 p2 pm for the simplicity of notation, Eqs. (1.88) or (1.89) is expressed as
Tp1 p2
¼ Qp1 q1 Qp2 q2 Qpm qm Tq1 q2
pm
ð1:90Þ
qm
or Tp1 p2
pm
¼
@xp1 @xp2 @xpm Tq q @xq1 @xq2 @xqm 1 2
ð1:91Þ
qm
Noting that Qp1 r1 Qp2 r2 Qpm rm Tp1 p2
pm
¼ Qp1 r1 Qp2 r2 Qpm rm Qp1 q1 Qp2 q2 Qpm qm Tq1 q2
qm
¼ ðQp1 r1 Qp1 q1 ÞðQp2 r2 Qp2 q2 Þ ðQpm rm Qpm qm ÞTq1 q2 ¼ dr1 q1 dr2 q2
qm
drm qm Tq1 q2 qm
the inverse relation of Eq. (1.90) is given by Tr1 r2
rm
¼ Qp1 r1 Qp2 r2 Qpm rm Tp1 p2
pm
ð1:92Þ
While the transformation rule of the first-order tensor, i.e. vector is given by Eqs. (1.80) and (1.82), the transformation rule of the second-order tensor is given by Tij ¼ Qir Qjs Trs ;
Tij ¼ Qri Qjs Trs
ð1:93Þ
The transformation between the coordinate systems without relative motion is considered above in the definition of the tensor, whereas the transformation in the
1.3 Tensor
19
form of (1.90) or (1.92) is called as the objective transformation. A tensor that obeys the objective transformation even between the coordinate systems with the relative motion is called an objective tensor.
1.3.2
Quotient Law
One has a convenient law, called the quotient law, which is used to judge whether or not a quantity is a tensor as will be explained below. Quotient law: “If a set of functions Tðp1 ; p2 ; ; pm Þ becomes Bpl þ 1 pl þ 2 pm (m-l-th order tensor lacking the suffices p1 pl ) by multiplying it by Ap1 p2 pl (l-th order tensor ðl mÞ), the set is an m-th order tensor”. (Proof) The proof can be achieved by showing that the quantity Tðp1 ; p2 ; ; pm Þ is the m-th order tensor when it holds that
Tðp1 ; p2 ; ; pm ÞAp1 p2
pl
¼ Bpl þ 1 pl þ 2
ð1:94Þ
pm
which is described in the coordinate system fO xi g as follows: T ðp1 ; p2 ; ; pm ÞAp1 p2
pm
¼ Bpl þ 1 pl þ 2
ð1:95Þ
pm
Here, the following relation holds. Bpl þ 1 pl þ 2
pm
¼ Qpl þ 1 rl þ 1 Qpl þ 2 rl þ 2 Qpm rm Brl þ 1 rl þ 2
rm
¼ Qpl þ 1 rl þ 1 Qpl þ 2 rl þ 2 Qpm rm Tðr1 ; r2 ; ; rm ÞAr1 r2 rl ¼ Qpl þ 1 rl þ 1 Qpl þ 2 rl þ 2 Qpm rm Tðr1 ; r2 ; rm Þ Qp1 r1 Qp2 r2 Qpl rl Ap1 q2 pl |fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} |fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl}
lþ1 m
1 l
ð1:96Þ Substituting Eq. (1.95) into Eq. (1.96) yields
T ðp1 ; p2 ; ; pm Þ Qp1 r1 Qp2 r2 Qpm rm T ðr1 ; r2 ; ; rm Þ Ap1 p2
pl
¼0
ð1:97Þ
from which it holds that T ðp1 ; p2 ; ; pm Þ ¼ Qp1 r1 Qp2 r2 Qpm rm T ðr1 ; r2 ; ; rm Þ
ð1:98Þ
Therefore, taking account of the definition of tensor in Eq. (1.88) into Eq. (1.98), the quantity Tðp1 ; p2 ; ; pm Þ is the m-th order tensor. ðQ.E.DÞ According to the proof presented above, Eq. (1.94) can be written as
20
1
Tp1 p2
Mathematical Preliminaries: Vector and Tensor Analysis pm Ap1 p2 pl
¼ Bpl þ 1 pl þ 2
ð1:99Þ
pm
For instance, if the quantity Tði; jÞ transforms the first-order tensor, i.e. vector vi to the vector ui by the operation Tði; jÞvj ¼ui , one can regard Tði; jÞ as the second-order tensor. Eventually, in order to prove that a certain quantity is a tensor, one needs only to show that it obeys the tensor transformation rule (1.90) or that the multiplication of a tensor to the quantity leads to a tensor by the quotient rule. Tensors fulfill linearity as follows: Tp1 p2 Tp1 p2
pm ðGp1 p2 pl pm ðaAp1 p2
þ Hp1 p2 pl Þ ¼ Tp1 p2 pl Þ ¼ aTp1 p2 pm Ap1 p2
pm Gp1 p2 pl
þ Tp1 p2
pm Hp1 p2 pl
pl
where a is an arbitrary scalar. Therefore, the tensor plays the role to transform linearly a tensor to the other tensor and thus it is called the linear transformation. The operation that lowers the order of tensor by multiplying the other tensor is called the contraction.
1.3.3
Notations of Tensors
When we express the tensor T as T ¼ Tp1 p2
pm ep1 ep2
ep m
ð1:100Þ
in a similar form to the case of vector in Eq. (1.25), Eq. (1.100) is called the component notation with bases, defining ep1 ep2 epm as the base of m-th order tensor. The transformation of T between the bases in Eq. (1.77) leads Eq. (1.100) to T ¼ Tp1 p2
pm Qr1 p1 er1
Qr2 p2 er2 Qrm pm erm
¼ Qr1 p1 Qr2 p2 Qrm pm Tp1 p2 ¼
Tr1 r2 rm er1
er2
p m er 1
er2 erm
erm
The following various notations are used for tensors. Indicial (or component) notation: Tp1 p2 pm . Component notation with base: Tp1 p2 pm ep1 ep2 epm . Symbolic (or direct) notation: T. Matrix notation: Eq. (1.6) for second-order tensor as an example.
ð1:101Þ
1.3 Tensor
21
The matrix notation holds only for a vector or a second-order tensor or for a fourth-order tensor if it is formally expressed by two suffixes. For instance, the stress–strain relation can be expressed in matrix notation by expressing the stress and the strain of second-order tensors as a form of vector and the stiffness coefficient of fourth-order tensor as a form of second-order tensor. Various contractions exist in the operation of higher-order tensors. and thus the symbolic notation is not useful in general. For instance, which of the following does ST mean: Sijk Tjk ; Sijk Tkj ; Sijk Tij ; Sijk Tji ; Sijk Tjl ; Sijk Tkl ; Sijk Til ? In other words, the application of symbolic notation is limited to the multiplication between low order tensors. On the other hand, component notation with bases holds always without defining special rule. Introducing the notation ðQ½½TÞp1 p2 pm Qp1 q1 Qp2 q2 Qpm qm T q1 q2
T Q ½½T p1 p2 pm Qq1 p1 Qq2 p2 Qqm pm Tq1 q2
qm
ð1:102Þ
qm
Equations (1.90) and (1.92) can be expressed by the symbolic notation as follows: T ¼ Q½½T
ð1:103Þ
T ¼ QT ½½T
In particular, transformations of the vector and the second-order tensor are expressed as follows: vi ¼ Qir vr ¼ vr Qir ;
vi ¼ Qri vr ¼ vr Qri
v ¼ Qv ¼ vQT ; v ¼ QT v ¼ v Q Tij ¼ Qir Qjs Trs ;
ð1:104Þ
Tij ¼ Qri Qsj Trs
T ¼ QTQ ; T ¼ Q T Q T
T
ð1:105Þ
noting Tv ¼ vTT ðTij vj ¼ vj Tij Þ in general for Eq. (1.104), where TT ððTT Þij ¼ Tji Þ is the transposed tensor defined exactly in Sect 1.4.3. The component of vector is expressed by the direct notation in Eq. (1.26). Here, consider the component of second-order tensor in the direct notation. The second-order tensor T is expressed from Eq. (1.100) as T ¼ Tij ei ej
ð1:106Þ
22
1
Mathematical Preliminaries: Vector and Tensor Analysis
from which, noting Eq. (1.69), it follows that ei Tej ¼ ei Trs er es ej ¼ Trs dir dsj and thus the component of T in the direct notation is given as Tij ¼ ei Tej
ð1:107Þ
As known from Eq. (1.107), the orthogonal projection of the vector Tej to the base vector ei is the component of the tensor T. Especially, Tii (no sum) and Tij ði 6¼ jÞ are called the normal and the shear components, respectively. Further, noting the relation ðTu TvÞij ¼ Tir ur Tjs us ¼ Tir ur us Tjs ¼ ðTðu vÞTT Þij , one has Tu Tv ¼ Tu vTT
1.3.4
ð1:108Þ
Orthogonal Tensor
The coordinate transformation operator Qij which appeared in Sects. 1.2.3 and 1.3.3 plays an important role in the coordinate transformation. The component notation with bases is obtained from Qij ei ej ¼ ei ðei ej Þej ð ¼ ei ei Þ ¼ ðei er Þer ei ¼ Qri er ei
ð1:109Þ
as follows: Q ¼ Qij ei ej ¼ Qij ei ej
ð1:110Þ
Furthermore, considering Eq. (1.77), the direct notation of Q is given by Q ¼ ei ei
ð1:111Þ
Because of ei ¼ er dir ¼ er er ei ei ¼ er dir ¼ er er ei
)
it follows that ei ¼ Qei ¼ Qri er ;
ei ¼ QT ei ¼ Qir er
ð1:112Þ
1.3 Tensor
23
Furthermore, changing Eq. (1.79) to the direct notation or noting the relation (
QQT ¼ ei ei ej ej ¼ ei dij ej ¼ ei ei QT Q ¼ ei ei ej ej ¼ ei dij ej ¼ ei ei
obtained from Eq. (1.111), it holds that QQT ¼ QT Q ¼ I
ð1:113Þ
where I possesses the components of the Kronecker’s delta, i.e. ðIÞij ¼ dij
ð1:114Þ
which transforms vector and tensor to original vector and tensor as Iv ¼ vI ¼ v and IT ¼ TI ¼ T, and thus it is called the identity tensor. The tensor satisfying Eq. (1.113) is defined as the orthogonal tensor in general, the properties of which are described below. It follows from Eq. (1.113) that QT ¼ Q1
ð1:115Þ
det Q ¼ det QT ¼ 1
ð1:116Þ
Moreover, one has
for the right-handed system, noting detðQQT Þ ¼ detQ detQT ¼ ðdet QÞ2 ¼ det I ¼ 1 which is derived from Eq. (1.113) with Eqs. (1.9), (1.17). Further, from Eq. (1.113) one obtains ðQIÞQT ¼ ðQIÞT Making the determinant of this equation and noting Eqs. (1.12), (1.17) and (1.116), it holds that detðQIÞ ¼ detðQIÞ ! detðQIÞ ¼ 0
ð1:117Þ
Then, one of the principal values of the orthogonal tensor is unity in order to satisfy ðQ1 1ÞðQ2 1ÞðQ3 1Þ ¼ 0 as known from the fact which will be described in Sect. 1.5.
24
1
1.4
Mathematical Preliminaries: Vector and Tensor Analysis
Operations of Tensors
Various basic operations of tensors are shown collectively in this section, which are used for the representations of constitutive relations in the later chapters.
1.4.1
Notations in Tensor Operations
The following notations are used throughout this book for the vectors a; b; c; d and v, the second-order tensors A; B and T, the third-order tensor N and fourth-order tensor T. 8 < a b ¼ trða bÞ for ai bi a b ¼ e:ða bÞ for eijk aj bk ð1:118Þ : a b for ai bj with a b c for ai bj cj
Tv for Tij vj ; vT for vi Tij v T for ðv TÞij ¼ eikl vk Tlj ; T v for ðT vÞij ¼ ejkl Tik vl
ð1:119Þ
v T ¼ vk ek Tlj el ej ¼ vk Tlj ekli ei ej ¼ ekli vk Tlj ei ej ¼ eikl vk Tlj ei ej
!
T v ¼ Tik ei ek vl el ¼ Tik vl ei eklj ej ¼ eklj Tik vl ei ej ¼ ejkl Tik vl ei ej 8 AB for Air Brj > > > A : B ¼ trðABT Þ ¼ trðAT BÞ for A B > > ij ij > > > T : I ¼ I : T ¼ trT for T d ¼ T > ij ij ii > < A B for Aij Bkl T :ða bÞ¼ a Tb for Tij ai bj > > > > ðT aÞb ¼ ða bÞT for Tij ar br ; ða TÞb ¼ a ðTbÞ for ai Tjr br > > > > > ða bÞ:ðc dÞ ¼ ða cÞðb dÞ for ai ci bj dj > : ða b cÞv ¼ ðc vÞa b for ai bj cr vr
N T for Nijk Tkl ; T N for Tij Njkl N : T for Nijk Tjk ; T : N for Tij Nijk
ð1:120Þ
ð1:121Þ ð1:122Þ
ð1:123Þ
1.4 Operations of Tensors
25
pffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffi jjvjj ¼ v v for p viffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi vi ffi pffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffiffi jjTjj ¼ T : T ¼ trðTTT Þ for Tij Tij @f 2 ðTÞ @f 2 ðTÞ for @T @T @Tij @Tkl
9 ðTvÞi ¼ Tir vr ; ðTSÞij ¼ Tir Srj ; T : S ¼ trðTST Þ ¼ Tij Sij = ðRvÞij ¼ Rijr vr ; ðRTÞijk ¼ Rijr Trk ; ðR : TÞi ¼ Rirs Trs ; ðN : TÞij ¼ Nijrs Trs ; ðT : NÞij ¼ Trs Nrsij ; ðN : PÞijkl ¼ Nijrs Prskl
ð1:124Þ
ð1:125Þ
ð1:126Þ
where v; ðT; SÞ, R and ðN; PÞ designate the vector, the second-order, the third-order and the fourth-order tensors, respectively.
1.4.2
Trace
An operation taking the sum of the components having the same suffixes, i.e. the sum of diagonal components in the matrix notation is called the trace and is expressed as tr Tð¼T : IÞ ¼ Trs drs ¼ Trr ¼ T11 þ T22 þ T33
ð1:127Þ
trðABÞð¼A : BT Þ ¼ Air Bri ¼ A11 B11 þ A12 B21 þ A13 B31 þ A21 B12 þ A22 B22 þ A23 B32 þ A31 B13 þ A32 B23 þ A33 B33
ð1:128Þ
The following relations hold for the trace. tr ðA þ SÞ¼ tr A þ tr B; trðaTÞ ¼ atr T; tr ðABÞ ¼ tr ðBAÞ; trðu vÞ ¼ u v ð1:129Þ
1.4.3
Various Tensors
Various basic tensors used widely in tensor operations are defined in this subsection. (1) Transposed tensor The tensor TT satisfying the following equation for any arbitrary vectors a and b is defined as the transposed tensor of a tensor T. a Tb ¼ b TT a
ð1:130Þ
26
1
Mathematical Preliminaries: Vector and Tensor Analysis
Noting a ðu vÞb ¼ b ðv uÞa it follows from Eq. (1.130) that ðu vÞT ¼ v u
ð1:131Þ
Further, comparing the equation ei ðTrs er es Þej ¼ Tij ¼ ej ðTrs es er Þei with Eq. (1.130), one has ðTrs er es ÞT ¼ Trs es er ¼ Tsr er es
ð1:132Þ
ðTT Þij ¼ ðTÞji
ð1:133Þ
ðA þ BÞT ¼ AT þ BT ð½ðA þ BÞT ij ¼ Ajr þ Bri Þ
ð1:134Þ
i.e.
The following relations hold.
It holds from ½ðABÞT ij ¼ Ajr Bri that ðABÞT ¼ BT AT tr TT ¼ tr T;
tr ðABÞ ¼ A : BT ¼ AT : B
ð1:135Þ ð1:136Þ
The magnitude of tensor is defined as the square root of the sum of the squares of each components and thus it is expressed using Eq. (1.133) as kTk ¼
ffi pffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Tij Tij ¼ T : T ¼ tr ðTTT Þð¼ trT2 for T¼TT Þ
ð1:137Þ
The tensor whose magnitude is unity is called the unit tensor. It follows by use of Eq. (1.133) that Tv ¼ vTT ; Tij vj ¼vj Tij Tu v ¼ u TT v ¼ u vT;
ðTij uj Þvi ¼ ui ðvj Tji Þ
Au Bv ¼ u AT Bv (
Tu v ¼ uTT v; ðTir ur Þvi ¼ ður Tir Þvi u Tv ¼ u vTT ; ui ðTjr vr Þ ¼ ui ðvr Tjr Þ
ð1:138Þ ð1:139Þ ð1:140Þ ð1:141Þ
1.4 Operations of Tensors
27
(2) Determinant The determinant is studied already in Sect. 1.1 but it will be studied in more detail in this subsection. The determinant is defined in Eqs. (1.8)–(1.10) already as follows: 9 T11 T12 T13 8 < epqr T1p T2q T3r = or det T ¼ jTij j ¼ T21 T22 T23 ¼ ¼ 1 e e T T T ð1:142Þ : ; 3! abc pqr ap bq cr epqr Tp1 Tq2 Tr3 T31 T32 T33 We obtain the following equations (
detðTÞ ¼ det T;
detðsTÞ ¼s3 det T;
detð ABÞ ¼ det A det B;
det T¼ det TT
detðTn Þ ¼ ð det TÞn ;
detðexp TÞ¼ expðtrTÞ ð1:143Þ
detða bÞ ¼ 0
ð1:144Þ
by virtue of 8 detðTÞ ¼ eijk ðT1i ÞðT2j ÞðT3k Þ ¼ eijk T1i T2j T3k > > > > > < detðsTÞ ¼ eijk sT1i sT2j sT3k ¼ s3 eijk T1i T2j T3k 1 1 1 > det T ¼ eabc epqr Tap Tbq Tcr ¼ epqr eabc Tpa Tqb Trc ¼ eabc epqr Tpa Tqb Trc > > > 3! 3! 3! > : detðABÞ ¼ epqr ðA1a Bap ÞðA2b Bbq ÞðA3c Bcr Þ ¼ epqr A1a A2b A3c Bap Bbq Bcr ¼ A1a A2b A3c eabc det B
ð1:145Þ expðT11 Þ 0 0 ¼ expðT11 þ T22 þ T33 Þ ð1:146Þ detðexp TÞ ¼ 0 0 expðT22 Þ 0 0 expðT33 Þ detða bÞ ¼ eijk ða1 bi Þða2 bj Þða3 bk Þ ¼ a1 a2 a3 eijk bi bj bk ¼ ða1 a2 a3 Þðb bÞ b ð1:147Þ noting Eqs. (1.8), (1.10), (1.17) and (1.45). The following equations hold for the cofactor, noting Eq. (1.13). 8 cofðsTÞ ¼ s2 cofT > > > > < ðcofTÞT ¼ cofðTT Þ; ðcofTÞ1 ¼ cofðT1 Þ; ðcofTÞT ¼ cofðTT Þ cofðABÞ ¼ cofðAÞcofðBÞ > > > trðcofTÞ ¼ 12 ðtr2 T trT2 Þ ¼ II > : detðcofTÞ¼ 1
ð1:148Þ
28
1
Mathematical Preliminaries: Vector and Tensor Analysis
noting 1 1 1 trðcofTÞ ¼ eabc eaqr Tbq Tcr ¼ ðdbq dcr dbr dcq ÞTbq Tcr ¼ ðTqq Trr Trq Tqp Þ 2 2 2 II is the second principal invariant as will be defined in Sect. 1.5. The vector product in Eq. (1.34) and the scalar triple product in Eq. (1.36) are described by the determinant as follows: e 1 e2 e3 u u u u u u 2 3 3 1 1 2 uv ¼ e1 þ e2 þ e3 ¼ u1 u2 u3 v 2 v3 v3 v1 v1 v2 v1 v2 v3 vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u u u2 u3 2 u3 u1 2 u1 u2 2 t jjuvjj ¼ þ þ v2 v3 v3 v1 v1 v2 u 1 u2 u3 u e 1 u e 2 u e 3 ½uvw ¼ v1 v2 v3 ¼ v e1 v e2 v e3 w 1 w 2 w 3 w e1 w e2 w e3 a c b c ða bÞ ðc dÞ¼ a d b d a1 a2 a3 p1 q2 r3 ai pi ½abc½pqr ¼ b1 b2 b3 p2 q2 r3 ¼ bi pi c1 c2 c3 p3 q2 r3 ci pi
ai qi ai ri a p a q a r bi qi bi ri ¼ b p b q b r ci qi ci ri c p c q c r
ð1:149Þ
ð1:150Þ
ð1:151Þ
ð1:152Þ
ð1:153Þ
noting Eq. (1.143)4 . The following equation is derived as the special case of Eq. (1.153) for the three vectors v1 ; v2 ; v3 . v 1 v1 v 1 v2 v1 v3 ð1:154Þ v2 v2 v2 v3 v2 ¼ ½v1 v2 v3 2 ¼ sym: v3 v 3 where v ¼ ½v1 v2 v3 ¼ eijk ðv1 Þi ðv2 Þj ðv2 Þk ð¼eijk ðv1 ei Þðv2 ej Þðv3 ek ÞÞ
ð1:155Þ
1.4 Operations of Tensors
29
is the volume of the parallelepiped formed by the line-elements v1 ; v2 ; v3 . By introducing the symbol vij vi vj
ð1:156Þ
v11 v12 v13 v22 v23 ¼ detðvij Þ v2 ¼ sym: v33
ð1:157Þ
Equation (1.154) can be written as
It follows from Eq. (1.153) that detT ¼ ½Te1 Te2 Te3 ¼ Te1 ðTe2 Te2 Þ
ð1:158Þ
noting T11 T12 T13 ðe1 Te1 Þ ðe1 Te2 Þ detT ¼ T21 T22 T23 ¼ ðe2 Te1 Þ ðe2 Te2 Þ T31 T32 T33 ðe3 Te1 Þ ðe3 Te2 Þ
ðe1 Te3 Þ ðe2 Te3 Þ ðe3 Te3 Þ
¼ ½e1 e2 e3 ½Te1 Te2 Te3 ¼ ½Te1 Te2 Te3 (3) Inverse tensor The tensor T1 fulfilling the following relation is defined as the inverse tensor of the tensor T. TT1 ¼ I;
Tir ðT1 Þrj ¼ dij
ð1:159Þ
Equation (1.14) is expressed in the direct notation as follows: ðdetTÞI ¼ TðcofTÞT ;
ðdetTÞdij ¼ Tip ðcofTÞjp
from which the cofactor is given as follows: cofT ¼ ðdetTÞTT ; (cofTÞij ¼
1 eipq ejrs Tpq Trs ¼ ðdetTÞTji1 2!
ð1:160Þ
and then the inverse tensor is given as follows: T1 ¼
ðcofTÞT ; detT
Tij1 ¼
ðcofTÞji detT
ð1:161Þ
30
1
Mathematical Preliminaries: Vector and Tensor Analysis
Equation (1.161) is represented for the 2 2 and 3 3 matrices as follows:
2 3 T11 T12 T13 " # T22 T33 T23 T32 T32 T13 T33 T12 T12 T23 T13 T22 T T
T T 22 21 11 12 6 7 = T1 ¼ ; T1 ¼4 T23 T31 T21 T33 T33 T11 T31 T13 T13 T21 T11 T23 5= T21 T22 T23 T12 T11 T21 T22 T21 T32 T22 T31 T31 T12 T32 T11 T11 T22 T12 T21 T31 T32 T33
ð1:162Þ Then, det T 6¼ 0 is required in order that T1 exists, so that the tensor fulfilling this condition is called the non-singular (or invertible) tensor. The partial derivative of Eq. (1.18) is rewritten by Eq. (1.160) as @ det T ¼ ðdet TÞTT @T
ð1:163Þ
The derivation of Eq. (1.163) starting from the definition of the total differential equation has been often described in some literatures (cf. Leigh, 1968; Hashiguchi and Yamakawa, 2012) but it needs cumbersome manipulations. Compared with it, the derivation shown above would be concise and straightforward. The following relations hold for the inverse tensor. ðTT Þ1 ¼ ðT1 ÞT ð TT Þ;
ðABÞ1 ¼ B1 A1
ð1:164Þ
ðA1 þ B1 Þ1 ¼ ½B1 ðB þ AÞA1 1 ¼ AðB þ AÞ1 B ¼ BðA þ BÞ1 A ð1:165Þ detTdetðT1 Þ ¼ detð TT1 Þ ¼1
ð1:166Þ
detðT1 Þ ¼ ðdetTÞ1
ð1:167Þ
because of ððTT1 ÞT ¼Þ ðT1 ÞT TT ¼ I ¼ ðTT Þ1 TT ABðABÞ1 ¼ I ! BðABÞ1 ¼ A1 Now, when we regard the transformation of the vector v to the vector u by the tensor T, i.e. Tv ¼ u; Tij vj ¼ ui
ð1:168Þ
as the simultaneous equation in which the components of v are the unknown numbers, solution exists for u 6¼ 0 if detT 6¼ 0 and is given by v ¼ T1 u, noting Eq. (1.161), as
1.4 Operations of Tensors
31
v¼
ðcofTÞT u; detT
vi ¼
ðcofTÞji detT
ð1:169Þ
uj
Here, T must be the non-singular tensor fulfilling det T 6¼ 0 in order that the non-trivial solution v 6¼ 0 exists for u 6¼ 0. On the other hand, T must be the singular tensor fulfilling det T ¼ 0 in order that the solution v 6¼ 0 exists for u ¼ 0. (4) Symmetric and skew-symmetric tensors Tensors TS and TA fulfilling the following relations are defined as the symmetric and the skew-(or anti-)symmetric tensor, respectively. TST ¼ TS ;
TjiS ¼ TijS
ð1:170Þ
TjiA ¼ TijA
ð1:171Þ
and TAT ¼ TA ;
An arbitrary tensor T is uniquely decomposed into the symmetric and the skew (anti)-symmetric tensors. T ¼ TS þ TA 2
3
2
T11 T11 T12 T13 6 T12 þ T21 7 6 6 6 4 T21 T22 T23 5 ¼ 6 2 4 T31 T32 T33 T13 þ T31 2
T12 þ T21 2 T22 T23 þ T32 2
ð1:172Þ
T13 þ T31 3 2 0 7 6 2 6 T23 þ T32 7 7 þ 6 T12 T21 7 6 2 2 5 4 T13 T31 T33 2
1 TS ¼ sym½T ¼ ðT þ TT Þ; 2
T12 T21 2 0 T23 T32 2
1 TA ¼ ant½T ¼ ðT TT Þ 2
sym½A þ B ¼ sym½A þ sym½B; ant½A þ B ¼ ant½A þ ant½B trðA0 BÞ ¼ trðAB0 Þ ¼ trðA0 B0 Þ
T13 T31 3 7 2 T23 T32 7 7 7 2 5 0
ð1:173Þ ð1:174Þ ð1:175Þ
while the components of TS and TA are often denoted by TðijÞ and T½ij , respectively. Equation (1.172) is called the Cartesian decomposition, following the decomposition of a complex number to a real and an imaginary parts. It holds that trTS ¼ trT;
tr TA ¼ 0;
ðTA Þii ¼ 0
tr( AS BA Þ ¼0
ðno sumÞ
ð1:176Þ ð1:177Þ
32
1
Mathematical Preliminaries: Vector and Tensor Analysis
det TA ¼ 0 (
u TS v ¼ v TS u u ðTA vÞ¼ v ðTA uÞ v ðTA vÞ ¼ 0
ð1:178Þ ð1:179Þ ð1:180Þ
noting Eq. (1.130). (5) Spherical and deviatoric parts The tensor T is decomposed as follows: T ¼ Tm þ T0
ð1:181Þ
9 1 1 Tm Tm I; Tm ðtr TÞ¼ Tii = 3 3 ; T0 TTm I ðtrT0 ¼ 0Þ
ð1:182Þ
while Tm and T0 are called the spherical (or mean) part and the deviatoric part of the tensor T. Noting Eq. (1.177), the skew-symmetric tensor T0A of the deviatoric tensor T0 is given by T0 A ¼ TA
ð1:183Þ
Then, the symmetric part of the deviatoric tensor is given by T0 S ¼ T0 TA ¼ TTm ITA
ð1:184Þ
T ¼ Tm I þ T0S þ TA
ð1:185Þ
from which one has
The decomposition of T into the mean component Tm I, the deviatoric symmetric component T0 S and the skew-symmetric component TA is called triple decomposition. The following holds: trðA0 BÞ ¼ trðAB0 Þ ¼ trðA0 B0 Þ
ð1:186Þ
(6) Axial vector The skew-symmetric tensor TA has three independent components in the three-dimensional state. Therefore, the vector tA having the following components is called the axial vector.
1.4 Operations of Tensors
33
A 1 tiA ¼ eirs TrsA t1A t2A t3A ¼ T23 2
A T31
A T12
ð1:187Þ
Inversely from Eq. (1.187) it is obtained that 2 TijA ¼ 4
TijA ¼ eijr trA ;
0
t3A 0 ant:
3 t2A t1A 5 0
ð1:188Þ
Furthermore, noting Eq. (1.36) and the relation TirA vr ¼ eirs tsA vr ¼ eirs trA vs
ð1:189Þ
the following relation holds. TA v ¼ tA v
ð1:190Þ
The relation of TA and tA is shown in Fig. 1.2 in the case that tA is the angular velocity vector and v is the position vector of particle. The quantity in Eq. (1.190) designates the peripheral velocity vector, while TA is called the spin tensor which induces the peripheral velocity by undergoing the multiplication of the position vector.
tA
|| v || sin
n
TAv = t A v = ( || t A |||| v ||sin ) n
v
0 Fig. 1.2 Anti-symmetric tensor and axial vector
34
1
Mathematical Preliminaries: Vector and Tensor Analysis
(7) Fourth-order transformation tensors The fourth-order tracing identity tensor T for a second-order identity tensor I is defined by T ≡ I ⊗ I = δ ijδ kl ei ⊗ e j ⊗ ek ⊗ el = ei ⊗ ei ⊗ e j ⊗ e j ð1:191Þ leading to T : T = T : T = (trT)I = 3Tm I for an arbitrary second-order tensor T. The fourth-order identity tensor I and the fourth-order transposing tensor I for a second-order tensor are defined by I ≡ δ ik δ jl ei ⊗ e j ⊗ ek ⊗ el = ei ⊗ e j ⊗ ei ⊗ e j I ≡ δ il δ jk ei ⊗ e j ⊗ ek ⊗ el = ei ⊗ e j ⊗ e j ⊗ ei
leading
I :T = T: I = T, I :T = T: I = TT ,
to
I :I = I :I = I
ð1:192Þ
and
∂T / ∂T = I , ∂T / ∂T = I .. T
The symmetrizing tensor S and the skew-(or anti-)symmetrizing tensor A are defined by S ≡ 1 (δ ik δ jl + δ ilδ jk )ei ⊗ e j ⊗ ek ⊗ el = 1 (I + I ) 2 2
A ≡ 1 (δ ik δ jl − δ ilδ jk )ei ⊗ e j ⊗ ek ⊗ el = 1 (I − I ) 2 2
ð1:193Þ
leading to S :T = T S , A :T = TA . The deviatoric projection tensor I ' is defined by I ' ≡ (δ ik δ jl − 1 δ ijδ kl )ei ⊗ e j ⊗ ek ⊗ el = I − 1 T 3 3
ð1:194Þ
leading to I ' :T = T : I ' = T' . The following relations hold from Eqs. (1.191)(1.194). ⎫ ⎪ ⎪ ⎪ T :I = I , T : S = T, T :A = O, T : I ' = O, ⎪ , (T : T)ijkl = δ ijTrrkl , I : T = T ⎬ ⎪ (S : T)ijkl = (Tijkl + T jikl ) / 2 , (A : T)ijkl = (Tijkl −T jikl ) / 2 , ⎪ ⎪ ⎪ (I ' : T)ijkl = Tijkl − 1 δ ijTrrkl 3 ⎭
T : T = 3 T, S : S = S , A : A = O, I ' : I ' = I ' ,
ð1:195Þ
where T is an arbitrary fourth-order tensor and O is the fourth-order zero tensor. Further, the following fourth-order tensors are defined by the four types of tensor products of the second-order tensors A; B; C (cf. e.g. del Piero 1979; Steinmann et al. 1997; Kintzel and Bazar 2006; Wang and Dui 2008).
1.4 Operations of Tensors
35
9 ðA BÞijkl ¼ Aij Bkl with A B : C ¼ AðB : CÞ ððA B : CÞij ¼ Aij ðBkl Ckl ÞÞ and > > > > > C : A B ¼ ðC : AÞBððC : A BÞkl ¼ Cij Aij Bkl Þ > > > > > T > ðABÞijkl ¼ Aik Bjl with AB : C ¼ ACB ððAB : CÞij ¼ Aik Bjl Ckl ¼ Aik Ckl Bjl Þ and > > > > > T > C : AB ¼ A CBððC : ABÞkl ¼ Cij Aik Bjl ¼ Aik Cij Bjl Þ > > > > > T T > ðABÞijkl ¼ Ail Bjk with AB : C ¼ AC B ððAB : CÞij ¼ Ail Bjk Ckl ¼ Ail Ckl Bjk Þ and > = C : AB ¼ BT CT AððC : ABÞkl ¼ Cij Ail Bjk ¼ Bjk Cij Ail Þ > > > > > e ijkl ¼ Aik Blj with A B e : C ¼ ACBððA B e : CÞij ¼ Aik Blj Ckl ¼ Aik Ckl Blj Þ and > ðA BÞ > > > > T T > e kl ¼ Cij Aik Blj ¼ Aik Cij Blj Þ > e ¼ A CB ððC : A BÞ C : A B > > > > T > ðA BÞijkl ¼ Ail Bkj with A B : C ¼ AC BððA B : CÞij ¼ Ail Bkj Ckl ¼ Ail Ckl Bkj Þ and > > >
> > > > C : A B ¼ BCT AððC : A BÞ ¼ Cij Ail Bkj ¼ Bkj Cij Ail Þ ;
kl
ð1:196Þ from which it follows that
I
I
ð1:197Þ
1.5
Eigenvalues and Eigenvectors
The application of the tensor T to an arbitrary vector v, i.e. Tv imposes v an extension or a contraction and a rotation in general. However, it imposes only an extension or a contraction without a rotation for the vector possessing the particular direction. Examine this fact in the following. Here, consider the following relation for the unit vector p. Tp ¼ Tp ;
ðTrj er ej Þp ¼ Tp ! ei
Tij pj ¼ Tpi
ðTrj er ej Þp ¼ ei
ð1:198Þ
Tp ! dir Trj pj ¼ Tei ! Tij pj ¼ Tpi
36
1
Mathematical Preliminaries: Vector and Tensor Analysis
i.e. ðTTIÞp ¼ 0; ðTij Tdij Þpj ¼ 0
ð1:199Þ
where the unit vector p is called the eigenvector (or principal or characteristic or proper vector) and the scalar T is called the eigenvalue (or principal or characteristic or proper value) when Eq. (1.198) holds. The eigenvector and the eigenvalue are called the eigenpair. Here, note the fact that the eigenvector is transformed to its scalar multiplication by a linear transformation, i.e. the multiplication of tensor. The necessary and sufficient condition that the simultaneous Eq. (1.199) has a non-zero solution of p is given by detðT TIÞ ¼ 0 ;
Tij Tdij ¼ 0
ð1:200Þ
noting u ¼ 0 in the solution of Eq. (1.169) for the simultaneous Eq. (1.168). Equation (1.200) is called the characteristic equation of the tensor, which is regarded as the cubic equation of T. Unit vectors P1 ; P2 ; P3 are derived for each of solutions T1 ; T2 ; T3 from Eq. (1.199). The mathematical verification of Eq. (1.200) is omitted even in the well-known books (e.g. Fung 1965; Ogden 1984; Bonet and Wood 1997; Holzapfel 2000; de Souza Neto et al., 2008, etc.). Then, it will be explained in detail without fear of redundancy in the following. Besides, the necessary and sufficient condition in Eq. (1.200) for the existence of the non-zero solution of p in Eq. (1.198) is verified by the other way: If the inverse tensor of TTI exists, Eq. (1.199) leads to Ip ¼ p ¼ 0 by multiplying ðTTIÞ1 to Eq. (1.199). In other words, the inverse tensor ðTTIÞ1 must not exist (must be a singular tensor) in order that the non-zero solution of p exists. The condition that the inverse tensor ðTTIÞ1 does not exist is nothing but (1.200), noting Eq. (1.169). Therefore, the simultaneous equation (1.199) for p becomes indeterminate. Then, the vector p cannot be determined only by this equation, while the eigenvalues T are already determined. Thus, an additional property for the magnitude of p must be added in order to determine p. Usually, the incidental condition that p is the unit vector, i.e. ||P|| = 1 is added in order to determined uniquely. The fact that Eq. (1.200) is the necessary and sufficient conditions that non-zero (non-trivial) solution(s) of P exist(s) in Eq. (1.199) is verified below by the elementary method. (Note) Consider the simultaneous equation for the unknown values x1 ; x2 ; x3 : 9 a11 x1 þ a12 x2 þ a13 x3 ¼ c1 > = a21 x1 þ a22 x2 þ a23 x3 ¼ c2 a31 x1 þ a32 x2 þ a33 x3 ¼ c3
> ;
ðaÞ
1.5 Eigenvalues and Eigenvectors
37
where aij ði;j ¼ 1; 2; 3Þ are the known coefficients. Equation (a) is described in the forms: (
Ax ¼ c: direct tensor notation ½Afxg ¼ fcg: matrix form
ðbÞ
setting 2
a11 6 A ¼ 4 a21
a12 a22
3 a13 7 a23 5;
a31
a32
a33
8 9 > = < x1 > x ¼ x2 ; > ; : > x3
8 9 > = < c1 > c ¼ c2 > ; : > c3
Solutions of Eq. (a) are given by c1 a12 a13 c2 a22 a33 c3 a32 a33 ; x1 ¼ a11 a12 a13 a21 a22 a23 a31 a32 a33
a11 c1 a13 a21 c2 a23 a31 c3 a33 ; x2 ¼ a11 a12 a13 a21 a22 a23 a31 a32 a33
a11 a12 c1 a21 a22 c2 a31 a32 c3 x3 ¼ a11 a12 a13 a21 a22 a23 a31 a32 a33
In the case of c1 ¼ c2 ¼ c3 ¼ 0 as seen in Eq. (1.200), only the trivial solutions x1 ¼ x2 ¼ x3 ¼ 0 can exist for det A 6¼ 0. Then, the following equation must hold in order that nontrivial solution exits. a11 a12 a13 T11 T det A ¼ a21 a22 a23 ¼ 0 ! detðTTIÞ ¼ T21 T31 a31 a32 a33
T12 T22 T T32
T13 T23 ¼ 0 T33 T
In what follows, it is proven that the eigenvalues are real and the eigenvectors are mutually orthogonal in the second-order real symmetric tensor. The complex conjugate relation for Eq. (1.198) is generally described as follows: Tp ¼ Tp designating the conjugate quantities by the over bar ðÞ. The multiplication of p and p to Eqs. (1.198) and (1.201) leads to p Tp ¼ Tp p;
p Tp ¼ Tp p
ð1:201Þ
38
1
Mathematical Preliminaries: Vector and Tensor Analysis
from which we have p Tpp Tp ¼ ðTTÞp p The left-hand side of this equation is zero because of p Tpp Tp ¼ p ðTTT Þ p ¼ 0. Then, we have T ¼ T so that T must be a real number. Further, it follows from Eq. (1.198) that Tpa ¼ Ta pa (no sum)
ð1:202Þ
for the eigenvectors pa ða ¼ 1; 2; 3Þ of T. By making the scalar products of Eq. (1.202) and the eigenvectors, we have pb Tpa ¼ Ta pa pb
)
pa Tpb ¼ Tb pb pa
ðno sum)
Subtracting the lower equation from the upper equation, one has pb Tpa pa Tpb ¼ ðTa Tb Þpa pb which reads: ðTa Tb Þpa pb ¼ 0 (no sum)
ð1:203Þ
noting pb Tpa pa Tpb ¼ pb Tpa TT pa pb ¼ pb ðTTT Þpa ¼ 0. The following facts can be concluded from Eq. (1.203). (1) If three principal values are all different to each other, there exist the three principal directions which are perpendicular to each other. (2) If two of three principal values are same, all directions in the plane perpendicular to the principal direction for the other principal value are the principal directions for the same principal value. (3) If all three principal values are same, all directions in the space are the principal directions. Based on the result described above, denoting the eigenvectors by pJ and the corresponding eigenvalues as TJ , one can write TpJ ¼ TJ pJ ðno sum)
ð1:204Þ
In addition, noting that the shear component on the coordinate system with the base vector feP g is zero, i.e. TPQ ¼ 0 ðP 6¼ QÞ
ð1:205Þ
1.5 Eigenvalues and Eigenvectors
39
the symmetric tensor T possessing orthogonal principal directions is expressed by T¼
3 P
TJ pJ pJ
ð1:206Þ
J¼1
which is called the spectral representation. Then, the tensor function of tensor T is described in general as fðTÞ
3 X
f ðTJ ÞpJ pJ
ð1:207Þ
J¼1
and then one has ½fðTÞT ¼ fðTT Þ
ð1:208Þ
The spectral representation of the asymmetric tensor can be referred to Istov (2015). ~ having the eigenvectors p ~J has the same eigenvalues TJ as those If the tensor T of the tensor T possessing the eigenvectors pJ , it holds that ~ pJ ¼ TJ p ~J ðno sumÞ T~
ð1:209Þ
where the orthogonal tensor Q between the eigenvectors of these tensors is given by ~J pK ; Q ¼ pJ p ~J ; pJ ¼ Q~ ~J ¼ QT pJ : QJK ¼ p pJ ; p
ð1:210Þ
Applying QT to Eq. (1.204), one has QT TpJ ¼ QT TJ pJ ¼ TJ QT pJ
ðno sumÞ
from which, considering Eqs. (1.209) and (1.210), it holds that ~J ¼ T ~P ~ J ¼ TJ P ~J QT TQP
ðno sumÞ
ð1:211Þ
Then, one obtains the relation ~ ¼ QT TQ T
ð1:212Þ
As presented above, tensors having identical eigenvalues can be related by the orthogonal tensor; they are called the similar tensor mutually. The coordinate transformation rule (1.105) of a certain tensor and the relation (1.212) of two tensors having identical eigenvalues but different eigenvectors, are of mutually opposite forms.
40
1
Mathematical Preliminaries: Vector and Tensor Analysis
If the function f of tensor A; B; is observed to be identical independent of observers, i.e. if it fulfills the relation f ðA;B; Þ ¼ f ðQ½½A;Q½½B; Þ
ð1:213Þ
using the symbol in Eq. (1.102), f is called the isotropic scalar-valued tensor function, which is none other than the invariant. In particular, if the isotropic scalar-valued tensor function f ðTÞ of single second-order tensor T fulfills f ðTÞ ¼ f ðQTQT Þ
ð1:214Þ
f ðTÞ can be expressed by three principal values in the three-dimensional case, including them in symmetric form so as to be identical even if they are exchanged to each other. Then, there exist three independent invariants for a single tensor. Their explicit forms are presented below. The expansion of the characteristic Eq. (1.200) of T leads to T11 T T21 T31
T12 T22 T T32
T13 T23 T33 T
¼ ðT11 TÞðT22 TÞðT33 TÞ þ T12 T23 T31 þ T21 T32 T13 ðT11 TÞT23 T32 ðT22 TÞT31 T13 ðT33 TÞT12 T21 ¼ T3 þ ðT11 þ T22 þ T33 ÞT2 ðT11 T22 þ T22 T33 þ T33 T11 ÞT þ T11 T22 T33 þ T12 T23 T31 þ T21 T32 T13 þ ðT12 T21 þ T23 T32 þ T31 T13 ÞT T11 T23 T32 T22 T31 T13 T33 T12 T21 ¼ T3 þ ðT11 þ T22 þ T33 ÞT2 ðT11 T22 þ T22 T33 þ T33 T11 T12 T21 T23 T32 T31 T13 ÞT þ T11 T22 T33 þ T12 T23 T31 þ T21 T32 T13 T11 T23 T32 T22 T31 T13 T33 T12 T21 ¼ 0
ð1:215Þ from which the characteristic equation is given as T 3 IT 2 þ IITIII ¼ 0
ð1:216Þ
I T11 þ T22 þ T33 ¼ Tii ¼ trT ¼ T1 þ T2 þ T3
ð1:217Þ
where
T11 T12 T22 T23 T33 T13 1 2 II þ þ ¼ ðT T Trs Tsr Þ¼ 12 ½ðtrTÞ trT2 T21 T22 T32 T33 T31 T11 2 rr ss ¼ T11 T22 þ T22 T33 þ T33 T11 T12 T21 T23 T32 T31 T13 ¼ T1 T2 þ T2 T3 þ T3 T1
ð1:218Þ
1.5 Eigenvalues and Eigenvectors
41
T11 T12 T 13 1 1 1 III T21 T22 T23 ¼ detT ¼ erst Tr1 Ts2 Tt3 ¼ ðtrTÞ3 trTtrT2 þ trT3 6 2 3 T31 T32 T33 ¼ T11 T22 T33 þ T12 T23 T31 þ T21 T32 T13 T11 T23 T32 T22 T31 T13 T33 T12 T21 ¼ T1 T2 T3
ð1:219Þ I, II, III designate the change of the total length, the total surface area and the volume, respectively, of the unit cube in the initial state, if T1, T2, T3 are the principal stretches in Eq. (4.19). The last expressions in III is derived from ðtrT3 I trT2 þ II trT 3III ¼Þ trT3 ðtrTÞtrT2 þ
1 ½ðtrTÞ2 trT2 trT 3III¼ 0 2
which is obtained by taking the trace of Eq. (1.250) in the Cayley–Hamilton theorem which will be described in Sect. 1.8 and by substituting Eq. (1.218) for II. On the other hand, the characteristic Eq. (1.215) is expressed using the three principal values T1 ; T2 ; T3 as follows: T1 T 0 0
0 T2 T 0
0 0 ¼ ðTT1 ÞðTT2 ÞðTT3 Þ ¼ 0 T3 T
ð1:220Þ
i.e. T 3 ðT1 þ T2 þ T3 ÞT 2 þ ðT1 T2 þ T1 T2 þ T3 T1 ÞT T1 T2 T3 ¼ 0 Comparing Eqs. (1.216) and (1.220), coefficients I, II and III are described as I ¼ T1 þ T2 þ T3 II ¼ T1 T2 þ T2 T3 þ T3 T1 III ¼ T1 T2 T3
ð1:221Þ
Equation (1.221) can also be derived by substituting T11 ¼ T1 ; T22 ¼ T2 ; T33 ¼ T3 , T12 ¼ T23 ¼ T31 ¼ 0 in Eqs. (1.217)–(1.219), while hereinafter principal value is denoted by only one suffix. Since I, II and III are the symmetric functions of principal values, they are the invariants and are called the principal invariants. The following invariants are called the moments. I I ¼ trT;
II trT2 ;
III trT3
ð1:222Þ
42
1
Mathematical Preliminaries: Vector and Tensor Analysis
The principal invariants are described in terms of these moments from Eqs. (1.217)–(1.219) as follows: I¼I 2
II ¼ 12 ðI IIÞ
ð1:223Þ
1 3 1 1 III ¼ I I II þ III 6 2 3 Next, consider the deviatoric tensor T0 . The characteristic equation of T0 is given by replacing T to T0 in Eq. (1.216) as follows: T 0 3 þ II0 T0 III0 ¼ 0
ð1:224Þ
0 0 0 I0 tr T0 ¼ T11 þ T22 þ T33 ¼0
ð1:225Þ
where
1 1 1 II0 ¼ Trs0 Tsr0 ¼ tr T02 ¼ jjT0 jj2 2 2 2 1 02 02 02 02 02 02 ¼ ðT11 þ T22 þ T33 Þ þ T12 þ T23 þ T31 2 1 02 02 02 þ T23 þ T31 g ¼ f ½ðT11 T22 Þ2 þ ðT22 T33 Þ2 þ ðT33 T11 Þ2 þ T12 6 1 1 ¼ ðT102 þ T202 þ T302 Þ¼ ½ðT1 T2 Þ2 þ ðT2 T3 Þ2 þ ðT3 T1 Þ2 2 6 0 0 0 0 0 0 0 0 0 0 0 0 ¼ ðT11 T22 þT22 T33 þT33 T11 T12 T21 T23 T32 T31 T13 Þ¼ ðT10 T20 þT20 T30 þT30 T10 Þ ð1:226Þ 0 0 0 T T T 11 12 13 0 0 0 1 1 0 III0 T21 T22 T23 ¼ det T0 ¼ eijk Ti10 Tj20 Tk3 ¼ trT03 ¼ Tij0 Tjk0 Tki0 3 3 0 0 0 T T T 31 32 33 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 ¼ T11 T22 T33 þ T12 T23 T31 þ T21 T32 T13 T11 T23 T32 T22 T31 T13 T33 T12 T21 ¼ T10 T20 T30
ð1:227Þ noting Eqs. (1.218) and (1.219).
1.5 Eigenvalues and Eigenvectors
43
(Note) 1 02 02 ðT þ T 02 þ T33 Þ 2 11( 22 2 2 2 ) 1 1 1 1 T11 ðT11 þ T22 þ T33 Þ þ T22 ðT11 þ T22 þ T33 Þ þ T33 ðT11 þ T22 þ T33 Þ ¼ 2 3 3 3 1 2 1 2 2 2 T11 ¼ T11 ðT11 þ T22 þ T33 Þ þ ðT11 þ T22 þ T33 Þ2 þ T22 T22 ðT11 þ T22 þ T33 Þ 2 3 9 3 1 2 1 2 T33 ðT11 þ T22 þ T33 Þ þ ðT11 þ T22 þ T33 Þ2 þ ðT11 þ T22 þ T33 Þ2 þ T33 9 3 9 1 2 2 3 2 2 2 þ T33 ðT11 þ T22 þ T33 Þ þ ðT11 þ T22 þ T33 Þ2 ¼ T11 þ T22 2 3 9 1 2 1 2 2 2 ¼ T11 þ T22 þ T33 ðT11 þ T22 þ T33 Þ 2 3 1 2 1 2 2 2 2 2 þ T22 þ T33 þ 2T11 T22 þ 2T22 T33 þ 2T33 T11 Þ ¼ T11 þ T22 þ T33 ðT11 2 3 1 2 2 2 2 ðT11 þ T22 þ T33 T11 T22 T22 T33 T33 T11 Þ ¼ 2 3 1 1 ½ðT11 T22 Þ2 þ ðT22 T33 Þ2 þ ðT33 T11 Þ2 ¼ 2 3 1 ¼ ½ðT11 T22 Þ2 þ ðT22 T33 Þ2 þ ðT33 T11 Þ2 6
Then, the following expressions in terms of the principal values hold from Eqs. (1.225)–(1.227) or from Eq. (1.221). I0 ¼ 0 II0 ¼ ðT10 T20 þ T20 T30 þ T03 T10 Þ 03 03 III0 ¼ T01 T02 T03 ¼ 13 ðT03 1 þ T2 þ T3 Þ
1.6
ð1:228Þ
Calculations of Eigenvalues and Eigenvectors
The second-order symmetric tensor can be represented by Eq. (1.206) as the spectral representation in the eigendirections. To express the tensor in the eigendirections, one must calculate the eigenvalues and the eigenvectors of the tensor. The solutions for them (cf. Hoger and Carlson 1984; Carlson and Hoger 1986) are shown in this section.
44
1.6.1
1
Mathematical Preliminaries: Vector and Tensor Analysis
Eigenvalues
In order to obtain eigenvalues, one try to calculate the deviatoric components from the characteristic Eq. (1.224) which is the cubic equation having the coefficients as the functions of invariants. Now, infer the form rffiffiffiffiffiffiffiffi 4 II0 cos w T ¼ 3 0
ð1:229Þ
for the eigenvalues of deviatoric part of tensor T. The substituting Eq. (1.229) into Eq. (1.224), we have 4II0 3=2 3 which is reduced to
cos3 w II0
4II0 1=2 3
cos wIII0 ¼ 0
pffiffiffi 4 pffiffiffi III0 3=2 cos 3wIII0 ¼ 0 3 3
ð1:230Þ
ð1:231Þ
using the trigonometric formula 1 cos3 w ¼ ðcos 3w þ 3 cos wÞ 4
ð1:232Þ
It is obtained from (1.231) that pffiffiffi 3 3III0 cos 3w¼ 2II0 3=2
ð1:233Þ
Noting that the cosine is the periodic function with the period 2p, the angle w is expressed by the following equation with a natural number J in general. w¼
pffiffiffi i 1 h 1 3 3III0 2pJ cos 0 3=2 3 2II
ð1:234Þ
Substituting Eq. (1.234) into Eq. (1.229), one has Tp0
1 ¼ 3
rffiffiffiffiffiffiffi n1h 3pffiffi3ffiIII0 io 4II0 cos 2pJ cos1 3 3 2II0 3=2
ð1:235Þ
and adding the isotropic component I=3, the eigenvalues of T are given as follows:
1.6 Calculations of Eigenvalues and Eigenvectors
1 TP ¼ Iþ 3
1.6.2
45
! rffiffiffiffiffiffiffi n1h pffiffiffi 0 io 4II0 1 3 3III cos 2pJ cos 3 3 2 II0 3=2
ð1:236Þ
Eigenvectors
Equation (1.206) can be expressed as follows: 3 X
T¼
ð1:237Þ
TP EP
P¼1
while the tensor EJ is called the eigenprojection of T, which is defined by EP eP eP (no sum)
ð1:238Þ
fulfilling 3 X
EP ð¼e1 e1 þ e2 e2 þ e3 e3 Þ ¼ I
ð1:239Þ
P¼1
(
EP for P¼Q
EP E Q ¼
O
for P 6¼ Q
EP : EQ ¼ dPQ
9 > > = > > ;
ð1:240Þ
It holds that
9 > > EP ¼ ð TQ EQ ÞeP eP ¼ ð TQ eQ eQ ÞeP eP ¼ TP eP eP > > > = Q¼1 Q¼1 3 X
E P T ¼ e P eP ð
3 X
3 X
TQ EQ Þ ¼ eP eP ð
Q¼1
3 X Q¼1
> > > TQ eQ eQ Þ ¼ eP eP TP > > ;
ðno sum for PÞ
and thus one has TEP ¼ EP T ¼ TP EP (no sum) On the other hand, it holds from Eq. (1.239) that TTQ I ¼
3 X P¼1
TP EP TQ
3 X P¼1
EP
ð1:241Þ
46
1
Mathematical Preliminaries: Vector and Tensor Analysis
and thus it is obtained that T TQ I¼
3 X
ðTP TQ ÞEP
ð1:242Þ
P¼1
from which one has 3 Y Q6¼h Q¼1
ðTTQ IÞ ¼
3 X 3 Y
ðTP TQ ÞEQ ¼
3 hY
Q6¼h P¼1 Q¼1
i ðTh TQ Þ Eh
ð1:243Þ
Q6¼h Q¼1
from which the following Sylvester’s formula is obtained. Eh ¼
3 TT I Q Q T T Q Q 6¼ h h
ð1:244Þ
Q¼1
For instance, E2 ðh¼ 2Þ in the popular case of n ¼ 3 is obtained from Eq. (1.244) as follows: Q ðTTP IÞ ðT T1 IÞðT T3 IÞ P¼1;3 ¼ E2 ¼ Q ðT2 TP Þ ðT2 T 1 ÞðT2 T 3 Þ P¼1;3
1.7
Eigenvalues and Eigenvectors of Skew-Symmetric Tensor
The characteristic equation of skew-symmetric tensor is given by substituting TA into Eq. (1.216) as follows: T 3 tr TA T 2 þ
1 2 A ðtr T tr TA2 ÞTdet TA ¼ 0 2
ð1:245Þ
Noting tr TA ¼ det TA ¼ 0 in Eqs. (1.176) and (1.178), Eq. (1.245) leads to ð2T 2 trTA 2 Þ
T ¼0 2
ð1:246Þ
from which the eigenvalues are given by qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi T ¼ i jtr TA 2 j=2 ¼ ijjtA jj
and 0
ð1:247Þ
1.7 Eigenvalues and Eigenvectors of Skew-Symmetric Tensor
47
noting 02
0 t3A A2 @ 4 trT ¼ tr 0 ant:
32 0 t2A t1A 54 0
t3A 0 ant:
31 t2A t1A 5A¼ 2ðt1A 2 þ t2A 2 þ t3A 2 Þ\0 0
obtained from Eq. (1.188), where tiA is the axial vectors defined in Eq. (1.187). It is known from Eq. (1.247) that one principal value is zero and two principal values are pure-imaginary numbers without a real part. Here, if one of the principal direction e3 is chosen to the direction of the zero A A A principal value of TA leading to T13 ¼ t23 ¼ 0, it follows by denoting T12 ¼ A t12 x that TA ¼ QTA QT 3 2 32 32 sinh 0 cos h sinh 0 0 x 0 cos h 7 6 76 76 ¼4 sinh cos h 0 54 x 0 0 54 sinh cos h 0 5 0 0 1 0 0 0 0 0 0 2 3 2 2 xsinh cos h þ xsinh cos h x sin h þ x cos h 0 6 7 ¼4 xsinh cos h xsinh cos h 0 5 x cos2 h xsin2 h 2
0 6 ¼4 x 0
0 3
x 0
0 7 0 5¼ TA
0
0
0
ð1:248Þ
0
meaning that the components do not change in the coordinate transformation. It is caused from the fact that the independent component of skew-symmetric tensor is only one when the one of bases in the coordinate system is chosen to the principal direction of skew-symmetric tensor.
1.8
Cayley-Hamilton Theorem
Denoting the eigenvector of the tensor T by e, it follows from Eq. (1.198) that Tr e ¼ T r e
ð1:249Þ
by the repeated applications of Eq. (1.198), i.e. T r e ¼ T r1 Te ¼ T r1 Te ¼ T T r2 Te ¼ TT r2 Te ¼ T2 T r2 e ¼ T2 T r3 Te ¼ ¼ Tr e. Equation (1.249) means that the eigenvalues of the tensor Tr are given by TJr where TJ ðJ¼ 1; 2; 3Þ are the
48
1
Mathematical Preliminaries: Vector and Tensor Analysis
eigenvalues of T, and the eigenvectors of the tensor Tr coincide with those of T. Tensors having an identical set of principal directions are called to be coaxial or said to fulfill the coaxiality. Then, the linear associative function fðTÞ of T is coaxial with T and the principal values of fðTÞ are given by f ðTJ Þ. The multiplication of the eigenvector e to the characteristic Eq. (1.216) leads to ðT3 I T2 þ II T III IÞe ¼ 0 noting Eq. (1.249). Because of e 6¼ 0, the following Cayley-Hamilton theorem holds. T3 I T2 þ II T III I ¼ O
ð1:250Þ
It follows from the Cayley-Hamilton theorem that T4 ¼ ðI T2 II T þ IIIIÞT ¼ I T3 II T2 þ III T ¼ IðIT2 II T þ III IÞ II T2 þ III T ¼ (I2 IIÞT2 ðIII IIIÞT þ III I
ð1:251Þ III T1 ¼ T2 I T þ II I
ð1:252Þ
It is concluded that the power of the tensor T is expressed by the linear associative of T2 ; T; I with coefficients consisting of the principal invariants. The Cayley-Hamilton theorem can be also derived from the characteristic equation as follows: Eq. (1.216) holds for the three principal values T1 , T2 and T3 , and thus the following equation must holds. 2
T13 IT12 þ II T1 III 6 0 4
0 0
0
T33 IT32 þ II T3 III
0 2
T13 0 0
3
2
3
0 T23 IT22 þ IIT2 III
T12 0 0
3
2
T1 0 0
3
2
1 0 0
3
7 5
2
0 0 0
3
6 6 6 7 7 7 6 7 6 7 3 2 6 6 7 7 7 6 7 6 7 ¼6 4 0 T2 0 5 I4 0 T2 0 5 þ II4 0 T2 0 5 III4 0 1 0 5 ¼ 4 0 0 0 5 0 0 T3 0 0 1 0 0 0 0 0 T33 0 0 T32 ð1:253Þ the tensor expression of which is nothing but Eq. (1.250).
1.8 Cayley-Hamilton Theorem
1.9
49
Scalar Triple Products with Invariants
The following formulae of the scalar triple products related to the principal invariants hold. ½Tu v w þ ½u Tv w þ ½u vT w ¼ trT½uvw ¼ I½uvw ½u Tv Tw þ ½Tu v Tw þ ½Tu Tv w ¼ II½uvw
ð1:254Þ
½Tu Tv Tw ¼ detT½uvw ¼ III½uvw where I; II; III are the principal invariants of T in Eqs. (1.217)–(1.219), i.e. I trT;
II ðtr2 T trT2 Þ=2; III detT
which will be delineated in the next section. The proof for Eq. (1.254) is given below (cf. Chadwick 1976; Kyoya 2008). Proof of Eq. (1.254)1: We have the relation ½Tu; v; w þ ½u; Tv; w þ ½u; v; Tw ¼ ui vj wk ð½Tei ; ej ; ek þ ½ei ; Tej ; ek þ ½ei ; ej ; Tek Þ ¼ eijk ui vj wk ð½Te1 ; e2 ; e3 þ ½e1 ; Te2 ; e3 þ ½e1 ; e2 ; Te3 Þ
ð1:255Þ
¼ ½uvwð½Te1 ; e2 ; e3 þ ½e1 ; Te2 ; e3 þ ½e1 ; e2 ; Te3 Þ making use of Eq. (1.49) with Eq. (1.47) and noting ½Tei ; ej ; ek þ ½ei ; Tej ; ek þ ½ei ; ej ; Tek ¼ eijk ð½Te1 ; e2 ; e3 þ ½e1 ; Te2 ; e3 þ ½e1 ; e2 ; Te3 Þ
ð1:256Þ Further, the inside of bracket in the last equation in Eq. (1.255) is described as follows: ½Te1 ; e2 ; e3 þ ½e1 ; Te2 ; e3 þ ½e1 ; e2 ; Te3 ¼ ½Tr1 er ; e2 ; e3 þ ½e1 ; Tr2 er ; e3 þ ½e1 ; e2 ; Tr3 er ¼ Tr1 ½er e2 e3 þ Tr2 ½e1 er e3 þ Tr3 ½e1 e2 er ¼ T11 ½e1 e2 e3 þ T22 ½e1 e2 e3 þ T33 ½e1 e2 e3 ¼ T11 þ T22 þ T33 ¼ trT ¼ I
ð1:257Þ Equation (1.254)1 is obtained by substituting Eq. (1.257) into Eq. (1.255). Proof of Eq. (1.254)2: In the similar way as Eq. (1.255), we have first
50
1
Mathematical Preliminaries: Vector and Tensor Analysis
½u; Tv; Tw þ ½Tu; v; Tw þ ½Tu; Tv; w ¼ eijk ui vj wk ð½e1 ; Te2 ; Te3 þ ½Te1 ; e2 ; Te3 þ ½Te1 ; Te2 ; e3 Þ
ð1:258Þ
¼ ½uvwð½e1 ; Te2 ; Te3 þ ½Te1 ; e2 ; Te3 þ ½Te1 ; Te2 ; e3 Þ Here, applying Eqs. (1.47) and (1.256), the inside of bracket in this equation is described as follows: ½e1 ; Te2 ; Te3 þ ½Te1 ; e2 ; Te3 þ ½Te1 ; Te2 ; e3 ¼ Tr2 Ts3 ½e1 er es þ Tr1 Ts3 ½er e2 es þ Tr1 Ts2 ½er es e3 ¼ Tr2 Ts3 e1rs þ Tr1 Ts3 er2s þ Tr1 Ts2 ers3 ¼ ðT22 T33 T23 T32 Þ þ ðT11 T33 T13 T31 Þ þ ðT11 T22 T12 T12 Þ ¼ T11 T22 þ T22 T33 þ T33 T11 ðT12 T12 þ T23 T32 þ T31 T13 Þ
ð1:259Þ
2 2 2 ¼ fT11 þ T22 þ T33 þ 2ðT11 T22 þ T22 T33 þ T33 T11 Þg=2 2 2 2 fT11 þ T22 þ T33 þ 2ðT12 T12 þ T23 T32 þ T31 T13 Þg=2 ¼ ðTrr Tss Trs Tsr Þ=2
leading to ½e1 ;Te2 ;Te3 þ ½Te1 ; e2 ; Te3 þ ½Te1 ; Te2 ; e3 ¼ ½ðtrTÞ2 tr2 T=2
ð1:260Þ
Equation (1.254)2 is obtained by substituting Eq. (1.260) into Eq. (1.258). Proof of Eq. (1.254)3: Changing the representations of the tree vectors to the component-based representations and then applying Eqs. (1.47) and (1.256), we have ½Tu; Tv; Tw ¼ eijk ui vj wk ½Te1 ;Te2 ;Te3 ¼ ½uvw½Te1 ;Te2 ;Te3
ð1:261Þ
Equation (1.254)3 is obtained by substituting Eq. (1.158) into Eq. (1.261). Tep ¼ Tij ei ej ep ¼ Tip ei jTe1 Te2 Te3 j ¼ jTi1 ei Tj2 ej Tk3 ek j ¼ Ti1 Tj2 Tk3 jei ej ek j ¼ Ti1 Tj2 Tk3 eijk je1 e2 e3 j ¼ det T with the aid of Eq. (1.47). It follows from Eq. (1.254)3 for T ¼ Q that Qu ðQv QwÞ ¼ det½Qu ðv wÞ
ð1:262Þ
1.9 Scalar Triple Products with Invariants
51
Then, the scalar product u ðv wÞ is transformed to the identical and the opposite handed system for det½Q ¼ 1 and det½Q ¼ 1, respectively, in the orthogonal transformation. Further, it follows from Eq. (1.254)3 that eijk det T ¼ ½Tei ; Tej ; Tek ; det T ¼ ½Te1 ;Te2 ;Te3
1.10
ð1:263Þ
Positive Definite Tensor
The second-order symmetric tensor P fulfilling Pv v [ 0
ð1:264Þ
for an arbitrary vector vð6¼ 0Þ is called the positive-definite tensor. Here, denoting the principal value and the corresponding eigen values of P as PJ and eJ , respectively, it follows that PeJ eJ ¼ PJ eJ eJ ¼ PJ jjeJ jj2 [ 0 ðno sum)
ð1:265Þ
Therefore, the principal values of the positive-definite symmetric tensor are all real and positive, i.e. PJ [ 0 and further it follows noting Eq. (1.221)3 that det P ¼ III [ 0
1.11
ð1:266Þ
Polar Decomposition
The following positive-definiteness holds for the symmetric tensors TT T and TTT . (
TT Tv v ¼ Tv Tv ¼ jjTvjj2 [ 0 TTT v v ¼ TT v TT v ¼ jjTT vjj2 [ 0
ð1:267Þ
noting Eq. (1.139). Therefore, TT T and TTT are the positive-definite tensors. Then, 2
designating their principal values by k2a and ka and their eigenvectors N a and na ða¼ 1; 2; 3Þ for TT T and TTT , respectively, the following representations in the spectral decomposition hold. 8 3 X > T > > T ¼ k2a N a N a C T > < a¼1
3 > X > 2 > T > ka na na : b TT ¼ a¼1
Further, introduce the following positive-definite tensors.
ð1:268Þ
52
1
Mathematical Preliminaries: Vector and Tensor Analysis
8 3 X > 1=2 > > U ¼ C ¼ ka N a N a > < a¼1
3 > X > > 1=2 > ka na na :V ¼ b ¼
ð1:269Þ
a¼1
which gives UN a ¼ ka N a ; Vna ¼ ka na (no sum)
ð1:270Þ
Further, let the following tensor be introduced. R ¼ TU 1 ;
R ¼ V 1 T
ð1:271Þ
for which one has (
RRT ¼ ðTU 1 ÞðTU1 ÞT ¼ TðU 2 Þ1 TT ¼ TT1 TT TT ¼ I R R ¼ ðV 1 TÞT ðV 1 TÞ ¼ TðV 2 Þ1 TT ¼ TT TT T1 T ¼ I T
ð1:272Þ
Therefore, R and R are the orthogonal tensors. It follows from Eq. (1.271) that T ¼ RU ¼ VR
ð1:273Þ
from which the following expressions are derived. U ¼ RT VR;
V ¼ RUR
T
ð1:274Þ
The substitution of Eq. (1.269)2 into Eq. (1.274)2 gives T
V ¼ RUR ¼ R
3 X
T
ka N a N a R ¼
a¼1
3 X
ka RN a RN a
a¼1
The coincidence of this equation with Eq. (1.269)2 requires the following relations. k a ¼ ka ; na ¼ RN a ;
R¼R N a ¼ RT na
ð1:275Þ ð1:276Þ
Then, it follows from Eqs. (1.269), (1.273), (1.274), (1.275) and (1.276) that
1.11
Polar Decomposition
53
T ¼ RU ¼ VR U ¼ RT VR;
V ¼ RURT
U ¼ C1=2 ¼ ðTT TÞ1=2 ¼ V¼b
T 1=2
¼ ðTT Þ
1=2
¼
3 P a¼1 3 P a¼1
R¼
3 P a¼1
T¼
3 P a¼1
na N a ;
ka na N a ;
ð1:277Þ ð1:278Þ
ka N a N a ð ¼ U T Þ
ð1:279Þ
ka na na ð ¼ V Þ T
det R ¼ 1
T1 ¼
3 1 P N a na a¼1 ka
ð1:280Þ
ð1:281Þ
Equation (1.277) is called the polar (spectral) decomposition, and RU and VR are called the right and left polar decompositions, respectively. U and V possess the same principal values ka but different principal directions N a and na , so that they are called the similar tensors to each other. Now, the following relation holds from Eq. (1.268) that dC¼
3 X
½2ka dka N a N a þ k2a ðdN a N a þ N a dN a Þ
ð1:282Þ
a¼1
from which one has N a dCN a ¼ 2ka dka
ð1:283Þ
noting N a dN a ¼ 0 derived from N a N b ¼ dab . Here, substitutions of N a dCN a ¼ dC:N a N a by virtue of Eq. (1.120)5 and dC ¼ ð@C=@ka Þdka into the left-hand side of Eq. (1.283) leads to @C dka : N a N a ¼ 2ka dka @ka
ð1:284Þ
1 @C dka : N a N a ¼ 1 2ka @ka
ð1:285Þ
leading to
Here, taking account of the contraction @C=@ka :@ka =@C ¼ 1, it follows from Eq. (1.285) that
54
1
Mathematical Preliminaries: Vector and Tensor Analysis
@ka ¼ ð2ka Þ1 N a N a @C
ð1:286Þ
Then, one has @k2a ¼ Na Na @C 9 > @k21 = ¼ IN 3 N 3 > @C > @k23 ; ¼ N3 N3 > @C 3 @k2 X ¼ Na Na ¼ I @C a¼1
for k1 6¼ k2 6¼ k3 6¼ k1 for k1 ¼ k2 6¼ k3
ð1:287Þ
for k1 ¼ k2 ¼ k3 ¼ k
noting @k2a @k2a @ka ¼ ¼ 2ka ð2ka Þ1 N a N a @C @ka @C Equation (1.287) is useful for the derivations of stress in the hyperelastic equation described by the principal stretches as will be shown in Chap. 7.
1.12
Isotropic Tensor-Valued Tensor Function
If the tensor-valued function f of tensors S; T; fulfills the following equation, it is called the isotropic function. Q½½fðS; T; Þ ¼ fðQ½½S; Q½½T; Þ
ð1:288Þ
where use is made of the symbol in Eq. (1.103). If f is a scalar, it is to be the invariant defined in Eq. (1.213) and if it is a tensor, it is called the isotropic tensor-valued tensor function. Now, consider the isotropic second-order tensor function B of a single second-order tensor A as the simplest case of Eq. (1.288), i.e. B ¼ fðAÞ
ð1:289Þ
fðQAQT Þ ¼ QfðAÞQT
ð1:290Þ
where f fulfills
First introducing the coordinate system with the bases e1 ; e2 ; e3 , which are the unit eigenvector of the tensor A and further adopting the another coordinate system
1.12
Isotropic Tensor-Valued Tensor Function
55
rotated 180° around the base e3 , the orthogonal tensor between the bases of these coordinate systems is given by 2
1 Q0 ¼ 4 0 0
3 0 0 1 0 5 0 1
ð1:291Þ
where Q0 fulfills Q0 ¼ QT0 resulting in the symmetric tensor and it holds that 2
1 4 0 0
0 1 0
38 0 9 8 0 9 0 > < > = > < > = 5 0 0 ¼ 0 ; > > > > 1 :1; :1;
i:e: Q0 e3 ¼ e3
ð1:292Þ
Then, it is known that e3 is one eigenvector not only of A but also of Q0 . Furthermore, denoting the principal values of A by A1 ; A2 ; A3 , it holds that 2
1 4 0 0
0 1 0
32 0 A1 0 54 0 0 1
0 A2 0
32 0 1 0 54 0 A3 0
0 1 0
3 2 0 A1 0 5¼4 0 0 1
0 A2 0
3 0 0 5; A3
i.e: Q0 AQT0 ¼ A
ð1:293Þ and thus it holds that fðQ0 AQT0 Þ ¼ fðAÞ ¼ B
ð1:294Þ
On the other hand, from Eqs. (1.289), (1.290) and (1.293) one has B ¼ fðAÞ ¼ fðQ0 AQT0 Þ ¼Q0 fðAÞQT0 ¼ Q0 BQT0
ð1:295Þ
Then, one obtains B ¼ Q0 BQT0 leading to the commutative law Q0 B ¼ BQ0
ð1:296Þ
and further, noting Eq. (1.292), the following relation is obtained. Q0 Be3 ¼ BQ0 e3 ¼ Be3
ð1:297Þ
which means that Be3 is the eigenvector of Q0 . Then, reminding that the eigenvector of Q0 is e3 , it follows that Be3 has the same direction as e3 . Then, it follows that Be3 ¼ B3 e3
ð1:298Þ
where B3 is the eigenvalue of B for the eigenvector e3 . Performing the similar manipulations also for e1 and e2 , it can be concluded that the tensor B has the same eigenvectors as the tensor A, leading to the coaxiality. Therefore, the principal
56
1
Mathematical Preliminaries: Vector and Tensor Analysis
values B1 ; B2 ; B3 of the tensor B can be represented in unique relation to the principal values A1 ; A2 ; A3 of the tensor A, i.e. ^ i ðA1 ; A2 ; A3 Þ Bi ¼ B
ði ¼ 1; 2; 3Þ
ð1:299Þ
Now, if the principal values Ai is different from each other, the most general expression of Bi is given by ^ i ðAj Þ ¼ /0 þ /1 Ai þ /2 A2i Bi ¼ B ði ¼ 1; 2; 3Þ Bi ei ei ¼ ð/0 þ /1 Ai þ /2 A2i Þei ei ðno sumÞ
ð1:300Þ
i.e. 9 B1 ðA1 ; A2 ; A3 Þ ¼ /0 þ /1 A1 þ /2 A21 > = B2 ðA1 ; A2 ; A3 Þ ¼ /0 þ /1 A2 þ /2 A2 ; 2 > B3 ðA1 ; A2 ; A3 Þ ¼ /0 þ /1 A3 þ /2 A23 ;
8 9 2 > 1 = < B1 > i:e: B2 ¼ 4 1 > ; : > 1 B3
A1 A2 A3
38 / 9 A21 > = < 0> 25 /1 A2 > > A23 : / ; 2
ð1:301Þ where /0 ; /1 ; /2 are scalar functions of A1 ; A2 ; A3 . The rightness of Eq. (1.300) or (1.301) can be verified by the reason that /0 ; /1 ; /2 can be determined if A1 ; A2 ; A3 and B1 ; B2 ; B3 are given, because the following Vandermonde’s determinant for the above simultaneous equation of the unknown variables /0 ; /1 ; /2 is not zero for mutually different values of A1 ; A2 ; A3 , i.e. 1 A1 1 A2 1 A3 leading to the B0 B1 B2 /0 ¼ 1 1 1
A21 A22 ¼ ðA1 A2 ÞðA2 A3 ÞðA3 A1 Þ 6¼ 0 A23
ð1:302Þ
solutions A1 A21 A2 A22 A3 A23 ; A1 A21 A2 A22 A3 A23
1 1 1 /1 ¼ 1 1 1
B1 B2 B3 A1 A2 A3
A21 A22 A23 ; A21 A22 A2 3
1 1 1 /2 ¼ 1 1 1
A1 A2 A3 A1 A2 A3
B1 B2 B3 A21 A22 A2
ð1:303Þ
3
as known from the description in Sect. 1.5. While Eq. (1.301) is regarded as a representation of the relation of the tensors A and B in their common principal coordinate system, it is expressed in the direct notation of tensor as
1.12
Isotropic Tensor-Valued Tensor Function
57
B ¼ /0 I þ /1 A þ /2 A2
ð1:304Þ
This fact can also be verified simply using Cayley-Hamilton theorem (Sect. 1.8) for the special case that f is a polynomial equation of the power of A in Eq. (1.289). However, for the case in which f is the general function of A, one must depend on the above-mentioned proof. It follows analogously to the above-mentioned method that B ¼ /0 I þ / 1 A
ð1:305Þ
when the two of principal values are same, and ð1:306Þ
B ¼ /0 I
when all principal values are same. In the particular case in which f is the linear function of the tensor A, Eq. (1.304) is reduced to B ¼ aðtr AÞI þ bA
ð1:307Þ
where a and b are the constants. Equation (1.307) is rewritten as B ¼ N: A
ð1:308Þ
1 N aT þ bS; Nijkl adij dkl þ bðdik djl þ dil djk Þ 2
ð1:309Þ
where
While the second-order isotropic tensor-valued tensor function of single tensor is considered above, the representation theorem of the second-order isotropic tensorvalued symmetric tensor function f S and anti-symmetric tensor function f A of the two tensors A and B are represented as follows (Rivlin 1955; Truesdell and Noll 1965). 9 f S ðA; BÞ ¼ u0 I þ u1 A þ u2 B þ u3 A2 þ u4 B2 þ u5 ðAB þ BAÞ > > > > 2 2 2 2 2 2 2 2 þ u ðA B þ BA Þ þ u ðAB þ B AÞ þ u ðA B þ B A Þ = 6
7
8
f ðA; BÞ ¼ g1 ðABBAÞ þ g2 ðA BBA Þ þ g3 ðAB2 B2 AÞ > > > > ; 2 2 2 2 þ g4 ðABA A BAÞ þ g5 ðBAB B ABÞ A
2
2
ð1:310Þ
where u0 ; u1 ; ; u8 and g1 ; g2 ; ; g5 are the scalar-valued isotropic functions of invariants of A and B, i.e. tr A; tr A2 ; tr A3 ; tr B; tr B2 ; tr B3 ð1:311Þ tr ðABÞ; tr ðAB2 Þ; tr ðA2 BÞ; tr ðA2 B2 Þ
58
1
Mathematical Preliminaries: Vector and Tensor Analysis
noting that all tensors can be represented by powers of tensor lower than the second power by the Cayley-Hamilton theorem in Sect. 1.8.
1.13
Representation of Tensor in Principal Space
Second-order tensor T is described by only three independent components, i.e. the principal values, in the directions of the eigenvectors e1 ; e2 ; e3 . Designating the principal values by T1 ; T2 ; T3 , it can be represented as follows: T ¼ T1 e1 þ T2 e2 þ T3 e3
ð1:312Þ
Equation (1.312) may be called the representation of tensor in principal space (Fig. 1.3) by which the second-order tensor can be visualized in the three dimensional space. Equation (1.312) is rewritten by decomposing T into the mean and the deviatoric components as follows: 0 ÞeIII ¼ Tm þ T0 T ¼ ðTm þ TI0 Þe1 þ ðTm þ TII0 Þe2 þ ðTm þ TIII
ð1:313Þ
where pffiffiffi Tm Tm Im ¼ 3Tm em ðTm ðTI þ TII þ TIII Þ=3Þ
ð1:314Þ
1 Im eI þ eII þ eIII ; em pffiffiffi Im ðjjem jj¼1Þ 3
ð1:315Þ
0 eIII ¼ jjT0 jjt0 T0 TI0 eI þ TII0 eII þ TIII
ð1:316Þ
TIII
' TIII TI' e III
0
Space diagonal
T
em TII e II TI e I
T' t'
Deviatoric plane
TII'
Tm
Fig. 1.3 Representation of second-order tensor in principal space
1.13
Representation of Tensor in Principal Space
59
TIII
n'III θ + 2π 3
_
0
n'I
T 'I
θ T'
TI
_
T 'II
n'II
2 π −θ 3
TII
_
( −) TIII '
Fig. 1.4 Deviatoric part of tensor in deviatoric plane (p-plane)
t0 T0 =jjT0 jjðjjt0 jj¼ 1Þ
ð1:317Þ
9 0 TI0 TI Tm ; TII0 TII Tm ; TIII TIII Tm = qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 02 ¼ pffiffiffi jjT0 jj ¼ TI0 2 þ TII0 2 þ TIII ðTI TII Þ2 þ ðTII TIII Þ2 þ ðTIII TI Þ2 ; 3
ð1:318Þ
Here, introduce the following unit vectors n0I , n0II and n0III which are the projections of the unit base vectors eI , eII and eIII onto the deviatoric plane as shown in Fig. 1.4. 9 1 > > n0I pffiffiffi ð2eI eII eIII Þðjjn0I jj ¼ 1Þ > > > 6 > > = 1 0 0 nII pffiffiffi ðeI þ 2eII eIII ÞðjjnII jj ¼ 1Þ > 6 > > > > 1 > 0 0 nIII pffiffiffi ðeI eII þ 2eIII ÞðjjnIII jj ¼ 1Þ > ; 6
ð1:319Þ
fulfilling n0I em ¼ n0II em ¼ n0III em ¼ cos ½ð1=2Þp ¼0; pffiffiffiffiffiffiffiffi n0I eI ¼ n0II eII ¼ n0III eIII ¼ 2=3; n0I n0II ¼ n0II n0III ¼ n0III
9 > > > =
> > > ; n0I ¼ cos ½ð2=3Þp ¼ 1=2
ð1:320Þ
60
1
Mathematical Preliminaries: Vector and Tensor Analysis
It follows from Eqs. (1.316) and (1.319) for the projections of T0 onto the directions n0I , n0II and n0III that 9 0 T I T0 n01 ¼ jjT0 jjt0 n01 ¼ jjT0 jj cos h > > > 2 > = 0 T II T0 n0II ¼ jjT0 jjt0 n0II ¼ jjT0 jj cos h p 3 > > 2 > > 0 0 0 0 0 0 0 T III T nIII ¼ jjT jjt nIII ¼ jjT jj cos h þ p ; 3
ð1:321Þ
where h is the angle measured in the anti-clock wise direction from n0I to T0 in the deviatoric plane as shown in Fig. 1.4 and it is called the Lode angle. Further, one 0 has the following relation for T I for example. 0
ðT I n0I Þ eI ¼ ½ðT0 n0I Þn0I eI 1 1 0 ¼ ½ðTI0 eI þ TII0 eII þ TIII eIII Þ pffiffiffi ð2eI eII eIII Þ pffiffiffi ð2eI eII eIII Þ eI 6 6 1 0 0 0 0 0 0 0 ¼ ð2TI TII TIII Þ 2 ¼ T1 ð*TI þ TII þ TIII ¼ 0Þ 6
Then, we have rffiffiffi 9 2 0 > > eI ¼ jjT eI ¼ jjT jj cos h > > > 3 > > rffiffiffi = 2 2 0 2 0 0 0 0 0 jjT jj cosðh pÞ > ð1:322Þ TII ¼ ðT II nII Þ eII ¼ jjT jj cosðh pÞnII eII ¼ 3 3 3 > > rffiffiffi > > 2 2 0 2 > > 0 0 0 0 0 jjT jj cosðh þ pÞ ; TIII ¼ ðT III nIII Þ eIII ¼ jjT jjðh þ pÞnIII eIII ¼ 3 3 3 T10 ¼
0 ðT I n0I Þ
0
jj cos h n0I
The product of the three components is given by III0
0 T 0 TII0 TIII 2 ¼ 10 0 0 ¼ 0 3=2 jjT jj jjT jj jjT jj 3 ð2II Þ
rffiffiffi 2 2 2 1 cos h cosðh pÞ cosðh þ pÞ ¼ pffiffiffi cos 3h 3 3 3 3 6 ð1:323Þ
noting Eqs. (1.226) and (1.228)3 and cos h cosðh þ
h1 2 2 4 i pÞ cosðh pÞ ¼ cos h ðcos 2h þ cos pÞ 3 3 2 3 1h 4 4 i ¼ cos 3h þ cos h þ cosðh þ pÞ þ cosðh pÞ 4 3 3 1h 4 i 1 ¼ cos 3h þ cos h þ 2 cos h cos p ¼ cos 3h 4 3 4
1.13
Representation of Tensor in Principal Space
61
exploiting the formula cos A cos B ¼ ½cosðA þ BÞ þ cosðABÞ=2, cos A þ cos B ¼ 2cos½ðA þ BÞ=2cos½ðABÞ=2. It follows from Eq. (1.323) that pffiffiffi pffiffiffi 3 3 III0 cos 3h ¼ ¼ 6tr t0 3 0 3=2 2 II
ð1:324Þ
noting Eqs. (1.226) and (1.227) for the third term. The different proof of this equation is shown in Ottosen and Ristinmaa (2005). Incidentally, the variable lð1 l 1Þ defined by l
0 pffiffiffi 2TII TI TIII 2TII0 TI0 TIII ¼ ¼ 3 tanðh p=6Þ 0 0 TI TIII TI TIII
ð1:325Þ
is called the Lode variable which is reduced to ( l¼
1 for TII ¼ TIII ðh ¼ 0 : triaxial extensionÞ þ 1 for TII ¼ TI ðh ¼ p=3 : triaxial compressionÞ
ð1:326Þ
describing the state of the intermediate principal stress. The following relation is exploited in Eq. (1.325)4. 2T2 T1 T3 T1 T3 2 cosðh 23 pÞ cos h cosðh þ 23 pÞ cosðh 23 pÞ cos h þ cosðh 23 pÞ cosðh þ 23 pÞ ¼ ¼ cos h cosðh þ 23 pÞ 2 sinðh þ 13 pÞ sinð 13 pÞ pffiffi pffiffi 1 1 4 2 sinðh 3 pÞ sinð 3 pÞ 2 sin h sinð 3 pÞ sinðh 13 pÞð 23Þ sin hð 23Þ pffiffi pffiffi ¼ ¼ 2 cosðh 16 pÞð 23Þ cosðh 16 pÞð 23Þ pffiffi sinðh 13 pÞ þ sin h 2 sinðh 16 pÞ cosð 16 pÞ 2 sinðh 16 pÞð 23Þ pffiffiffi 1 ¼ ¼ ¼ 3 tanðh pÞ ¼ 6 cosðh 16 pÞ cosðh 16 pÞ cosðh 16 pÞ
ð1:327Þ
1.14
Two-Dimensional State
Consider the two-dimensional state in which the components related to the e3 direction in the coordinate system ðx1 ; x2 ; x3 Þ with the fixed base ðe1 ; e2 ; e3 Þ are zero, i.e. T33 ¼ T 31 ¼ T23 ¼ 0. Furthermore, introduce the coordinate system ðx1 ; x2 ; x3 Þ with the bases ðe1 ; e2 ; e3 ð¼ e3 ÞÞ which is rotated by the angle a in the counterclockwise direction around the axis x3 as shown in Fig. 1.5, where the normal components Tij ði; j ¼ 1; 2; i ¼ jÞ and the shear components Tij ði 6¼ jÞ in
62
1
x2
Mathematical Preliminaries: Vector and Tensor Analysis
x2
Tt
Material
x1* * * (T11 , T12 )
x*2 x1*
RT
(T22 , T12 ) x*2
α 0
x1
2
T2 P
2 p T1
Tm
Tn
(T11 , T12 ) x1
(T22* , T12* ) (a) Physical plane
(b) (Tn , Tt ) plane
Fig. 1.5 Mohr’s circle
the base ðe1 ; e2 Þ are designated by Tn and by Tt , respectively. The orthogonal tensor between these bases is given from Eq. (1.78) as follows: 2
3 0 05 1
cos a sina ½Q ¼4 sina cosa 0 0
ð1:328Þ
Substituting Eq. (1.328) into Eq. (1.93), one has 9 T11 ¼ T11 cos2 a þ T22 sin2 a þ 2T12 cos a sin a > =
T22 ¼ T11 sin2 a þ T22 cos2 a 2T12 sin a cos a T12
¼ ðT22 T11 Þ cos a sin a þ T12 ðcos a sin aÞ 2
2
> ;
ð1:329Þ
which is rewritten as 9 T11 ¼ Tm þ T cos 2a þ T12 sin 2a > =
T22 ¼ Tm T cos 2a T12 sin 2a > ; ¼ T sin 2a þ T12 cos 2a T12
ð1:330Þ
where Tm
T11 þ T22 ; 2
T
Furthermore, it follows from Eq. (1.330) that
T11 T22 2
ð1:331Þ
1.14
Two-Dimensional State
63 T11 þ T22 ¼ T11 þ T22
@T11 ¼ 2T12 ; @a
ð1:332Þ
@T22 ¼ 2T12 @a
ð1:333Þ
While the axis x3 ð¼ x3 Þ is one of the principal directions, the other principal directions exist on the plane ðx1 ; x2 Þ. Denoting the principal direction from the x1 =@a ¼ T12 ¼ 0 or @T22 =@a ¼ 0 in axis by a, it is obtained by substituting @T11 Eq. (1.333) into Eq. (1.330) that tan 2ap ¼
T12 T
ð1:334Þ
from which one obtains T ¼ RT cos 2ap ; qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 RT ¼ T þ T212
9 T12 ¼ RT sin 2ap = ;
ð1:335Þ
Substituting Eq. (1.335) into the upper two of Eq. (1.330) with specifying a as ap , the maximum and the minimum principal values T1 and T2 are described by T1 T2
) ¼ Tm R T
ð1:336Þ
Equation (1.336) can also be derived directly from the quadratic equation 2 ¼0 T 2 ðT11 þ T22 ÞT þ T11 T22 T12 2 which is obtained by inserting I ¼ T11 þ T22 ; II ¼ T11 T22 T12 ; III ¼ 0 ðT33 ¼ T31 ¼ T23 ¼ 0Þ in Eq. (1.216). Furthermore, denoting a for the extremal value of T12 as as , it follows by taking @T12 =@a ¼ 0 in Eq. (1.330) that
tan 2as ¼
T 2T12
ð1:337Þ
Equations (1.334) and (1.337) yield the relation as ¼ ap p=4 (tan 2ap tan 2as ¼ 1) and thus there exist the two directions for the extremal values of T12 and they divide the two principal directions into two equal angles, i.e. p=4. The denoted by TM is given from Eq. (1.335)2 as follows: extremal values of T12
64
1
Mathematical Preliminaries: Vector and Tensor Analysis
TM ¼ RT
ð1:338Þ
which is also expressed by Eq. (1.336) as TM ¼
T1 T2 2
ð1:339Þ
The following equation is derived from Eqs. (1.330) and (1.335)3. 2
ðTn Tm Þ2 þ Tt2 ¼ RT
ð1:340Þ
noting the following equations obtained from Eq. (1.335)1;3 or (1.335)2;3 with the replacement of T11 ðT22 Þ ! Tn ; T12 ! Tt . 2
2 ðTn Tm Þ2 ¼ T cos2 2a þ T12 sin2 2a þ 2TT12 sin 2a cos 2a
)
2
2 Tt2 ¼ T sin2 2a þ T12 cos2 2a 2TT12 sin 2a cos 2a
Consequently, the components on an arbitrary plane is expressed by the point on the circle with the radius RT centering at ðTm ; 0Þ in the two-dimensional plane ðTn ; Tt Þ as shown in Fig. 1.5. This circle is called the Mohr’s circle. Substituting Eq. (1.335) into Eq. (1.330), we have the expressions 9 ¼ Tm þ RT cosð2a 2ap Þ; > T11 =
T22 ¼ Tm RT cosð2a 2ap Þ; > ; T12 ¼ RT cosð2a 2ap Þ
ð1:341Þ
Therefore, T11 and T12 are shown by the values in the ordinate and abscissa axes, respectively, of the point rotated 2a counterclockwise from point T11 ; T12 on the Mohr’s circle as shown in Fig. 1.5, provided that the definition for the sign of shear component is altered to be positive when it applies to the body surface in the clockwise direction, in the Mohr’s circle. As shown in Fig. 1.5, the intersecting angle of the two straight lines drawn in parallel to the physical plane x1 and x1 stemming from the points ðT11 ; T12 Þ and ðT11 ; T12 Þ, respectively, on the Mohr’s circle is a which is the angle of circumference of Mohr’s circle and thus the intersecting point lies on the circle. This point is called the pole. Generally speaking, the normal component Tn and the shear component Tt applying to a certain physical plane are given by the intersecting point of Mohr’s circle and the straight line drawn parallel to that physical plane from the pole P.
1.14
Two-Dimensional State
1.15
65
Tensor Functions
The following various tensor functions are defined. sin T ¼
1 X ð1Þn n 1 1 TP eP eP ¼ ðTP TP3 þ TP5 ÞeP eP ð2n þ 1Þ! 3! 5! n¼0
1 X ð1Þn 2n þ 1 1 1 T ¼ T T3 þ T5 ¼ ð2n þ 1Þ! 3! 5! n¼0
cos T ¼
1 1 TP2n eP eP ¼ 1 TP2 þ TP4 eP eP ð2nÞ! 2! 4!
1 X ð1Þn n¼0
1 X ð1Þn 2n 1 1 T ¼ I T2 þ T4 ¼ ð2nÞ! 2! 4! n¼0
exp T ¼
1 X 1 n 1 1 TP eP eP ¼ ð1 þ TP þ TP2 þ TP3 þ n! 2! 3! n¼0
1 1 ¼ I þ T þ T2 þ T 3 þ 2! 3! log T ¼
1 X ð1Þn1 n¼1
ð1:342Þ
n
ð1:343Þ
ÞeP eP
1 X 1 n T ¼ n! n¼0
ð1:344Þ
1 1 ðTp 1Þn eP eP ¼ ðTp 1Þ ðTp 1Þ2 þ ðTp 1Þ3 eP eP 2 3
1 X 1 1 ð1Þn1 ¼ ðT IÞ ðT IÞ2 þ ðT IÞ3 ¼ ðT IÞn 2 3 n n¼1
ð1:345Þ These functions satisfy the following basic properties noting Eqs. (1.207) and (1.344). ðexp TA ÞT ¼ ðexp TAT Þ ¼ expðTA Þ
ð1:346Þ
ðexp TÞT ¼ expðTT Þ
ð1:347Þ
expðnTÞ ¼ ðexp TÞn
ð1:348Þ
expðA þ BÞ ¼ exp A exp B
ð1:349Þ
det½expðA þ BÞ ¼ expðtrAÞ expðtrBÞ
ð1:350Þ
ðexp TA Þðexp TA ÞT ¼ ðexp TA Þðexp TAT Þ ¼ expðTA TA Þ ¼ I
ð1:351Þ
66
1
Mathematical Preliminaries: Vector and Tensor Analysis
noting 1 X 1 ðA þ BÞn n! n¼0
¼
1 1 1 X X 1 n X 1 n 1 n n1 n A Þð AÞ ðA þ nAn1 B þ þ nAB þ B Þ ¼ ð n! n! n! n¼0 n¼0 n¼0
det½exp T ¼ ð exp TÞ1 ðexp TÞ2 ðexp TÞ3 ¼ exp T1 exp T2 exp T3 ¼ expðT1 þ T2 þ T3 Þ for Eqs. (1.349) and (1.350). For tensor functions of diagonalizable tensors, one can obtain alternative closed-form expressions for the tensor functions and their derivatives in the spectral representations. Numerical computation of the exponential and logarithmic functions of a tensor, as well as their derivatives, can be referred to de Souza Neto (2001), Ortiz et al. (2001), Itskov and Aksel (2002), Itskov (2003, 2019), etc.
1.16
Partial Differential Calculi
Partial derivatives of symmetric tensors appearing often in elastoplasticity are shown this section. (1) Power functions n @Tn X ~ þ Tn1 I ~ ~ n1 þ TT ~ n2 þ þ Tn2 T ~ nk ¼ IT Tk1 T ¼ @T k¼1
@Tn @T
ijkl
¼
n X
ðTk1 Þik ðTnk Þlj
ð1:352Þ
k¼1
Examples for low-order tensors are shown below. @Tij ¼ dik djl @Tkl @ðTir Trj Þ ¼ dik drl Trj þ Tir drk djl ¼dik Tlj þ Tik djl @Tkl @ðTir Trs Tsj Þ ¼ dik drl Trs Tsj þ Tir drk dsl Tsj þ Tir Trs dsk djl ¼dik Tls Tsj þ Tik Tlj þ Tir Trk djl @Tkl
ð1:353Þ
1.16
Partial Differential Calculi
67
using the symbol in Eq. (1.196)3 . (2) Invariants @ðTrs drs Þ ¼ dir djs drs ¼ dij @Tij @ðTrs Tsr Þ ¼ dri dsj Tsr þ Trs dsi drj ¼ 2Tji @Tij @ðTrs Tst Ttr Þ ¼ dri dsj Tst Ttr þ Trs dsi dtj Ttr þ Trs Tst dti drj ¼Tjt Tti þ Tri Tjr þ Tjs Tsi ¼ 3Tjt Tti @Tij
@Tij0 ¼ I 0ijkl @Tkl
@Tij0 @ðTij Tm dij Þ 1 ¼ ¼ dik djl dij dkl @Tkl @Tkl 3
@trT @I @I ¼ ¼ ¼ I; @T @T @T
@trT2 @ II ¼ ¼ 2TT ; @T @T
@trT3 @ III ¼ ¼ 3TT 2 ð1:354Þ @T @T
2 @II @ 12 ððtrTÞ trT2 Þ ¼ I I TT ¼ @T @T @III @ detT ¼ ¼ ðdetTÞTT ¼ II I I TT þ T2T ¼ ðII I IT þ T2 ÞT @T @T ð1:355Þ
noting Eq. (1.163) and h1 1 1 3i 3 2 @III @ 6 ðtrTÞ 2 trTtrT þ 3 trT 3 1 1 1 ¼ ðtrTÞ2 I ðtrT2 ÞI ðtrTÞ2TT þ 3TT2 ¼ @T @T 6 2 2 3 1 2 2 T T2 ¼ ½ðtrTÞ trT I ðtrTÞT þ T 2
Then, the following equations are derived.
and
pffiffiffiffiffiffiffiffiffiffi @ detT 1 pffiffiffiffiffiffiffiffiffiffi T ¼ detTT @T 2
ð1:356Þ
pffiffiffiffiffiffiffiffiffiffi @ ln detT 1 1 ¼ T 2 @T
ð1:357Þ
! pffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffi @ ln det T 1 @ det T 1 1 ¼ pffiffiffiffiffiffiffiffiffiffi ¼ pffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffi ðdet TÞT1 @T det T @T det T 2 det T
68
1
Mathematical Preliminaries: Vector and Tensor Analysis
Further one has pffiffiffiffiffiffiffiffiffiffiffi @ Trs Trs 1 Tij @jjTjj T ¼ ðTrs Trs Þ1=2 ðdri dsj Tsr þ Trs dri dsj Þ ¼ pffiffiffiffiffiffiffiffiffiffiffi ; ¼ @Tij 2 @T jjTjj Trs Trs
@
ð1:358Þ
ð1:359Þ
pffiffiffiffiffiffiffiffiffiffiffi Trs0 Trs0 1 0 0 1=2 @ðTrs0 Trs0 Þ 1 0 0 1=2 ¼ ðTrs Trs Þ ¼ ðTrs Trs Þ 2dir djs Trs0 ¼ tij0 ; 0 @Tij0 2 2 @Tij ð1:360Þ
(3) Deviatoric tensors
@tij0 @Tkl0
¼
Tij0 @ pffiffiffiffiffiffiffiffiffiffiffiffiffi 0 T0 Tpq pq @Tkl0
dik djl ¼
pffiffiffiffiffiffiffiffiffiffiffiffiffi Tkl0 0 T 0 T 0 pffiffiffiffiffiffiffiffiffiffiffiffiffi Tpq pq ij 0 T0 Tpq pq 0 T0 Tpq pq
1 ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffi ðdik djl tij0 tkl0 Þ; 0 T0 Tpq pq
ð1:361Þ
@tij0 @tij0 @Trs0 1 1 0 ¼ ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffi Þðdrk dsl drs dkl Þ ðdir djs tij0 trs 0 T0 @Tkl @Trs0 @Tkl 3 Tpq pq 1 1 ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffi ðdik djl dij dkl tij0 tkl0 Þ; 0 T0 3 Tpq pq
cos 3h
pffiffiffi 0 3 6trt
ð1:362Þ ð1:363Þ
pffiffiffi 0 0 0 pffiffiffi pffiffiffi tqr trp @ cos 3h @ 6tpq 0 0 ¼ ¼ 3 6dip djq tqr trp ¼ 3 6tir0 trj0 ; 0 0 @tij @tij pffiffiffi @ cos 3h ¼ 3 6t 0 2 0 @t
ð1:364Þ
1.16
Partial Differential Calculi
69
0 pffiffiffi 0 0 1 @ cos 3h @ cos 3h @trs 0 0 ¼ ¼ 3 6trp tps 0 ðdri dsj trs tij Þ 0 @Tij0 @Tij0 @trs jjT jj 3 pffiffiffi ¼ 0 ð 6tit0 ttj0 cos 3htij0 Þ jjT jj
@ cos 3h 3 pffiffiffi ¼ 0 ð 6t0 cos 3ht0 Þ 0 @T jjT jj
ð1:365Þ
0 @ cos 3h @ cos 3h @trs ¼ 0 @Tij @Tij @trs pffiffiffi 1 1 0 0 d ¼ 3 6trt0 tts0 pffiffiffiffiffiffiffiffiffiffiffiffiffi d d t t d ir js ij rs ij rs 0 T0 3 Tpq pq pffiffiffi 1 1 0 0 0 0 0 0 0 0 t t t d t t t t t ¼ 3 6 pffiffiffiffiffiffiffiffiffiffiffiffiffi ij ir rj rs st tr ij ; 0 T0 3 rt tr Tpq pq
pffiffiffi @ cos 3h 3 6 0 2 1 1 0 2 0 ¼ 0 t pffiffiffi cos 3ht jjt jj I @T jjT jj 3 6
ð1:366Þ
(4) Various tensor functions Differentiating Tir1 Trs ¼ dis , one has @ðTir1 Trs Þ @Tir1 @T 1 ¼ Trs þ Tir1 drk dsl ¼ 0 ! ir Trs Tsj1 þ Tir1 drk dsl Tsj1 ¼ 0 @Tkl @Tkl @Tkl
d ij ¼ ðTir1 Trs Þ ¼ T
1 ir
Trs þ Tir1 T
rs
¼0
which leads to
ð1:367Þ
Further, one has 1 n 1 n @ exp T X @ exp T X 1X 1X ~ nk ; ¼ Tk1 T ¼ ðTk1 Þik ðTnk Þlj ijkl @T n! k¼1 @T n! k¼1 n¼1 n¼1
ð1:368Þ
70
1
Mathematical Preliminaries: Vector and Tensor Analysis
noting 1 2 1 3 @ exp T @ðI þ T þ 2! T þ 3! T þ Þ ¼ @T @T due to Eq. (1.344) with Eq. (1.352). One should notice @½f ðAÞBij @f ðAÞ @Bij ¼ Bij þ f ðAÞ @Tkl @Tkl @Tkl @½f ðAÞB @f ðAÞ @B @f ðAÞ @B ¼B þ f ðAÞ B þ f ðAÞ 6¼ @T @T @T @T @T For symmetric tensors, the fourth-order identity tensor replaced to the fourth-order symmetrizing tensor equations.
1.17
ð1:369Þ can be in the above
Differentiation and Integration in Tensor Field
Scalar s, vector v, and tensor T are called the scalar field, the vector field, and the tensor field, respectively when they are functions of the position vector x. Their differentiation and integration in fields are shown below, in which the following operator, called the nabra or Hamilton operator, is often used. $
@ @ er ¼ @xr @x
ð1:370Þ
(1) Gradient Scalar field : grads ¼ $s ¼
@s er @xr
8 @ @vi > > ej ¼ ei ej : rear (right) form < v $ ¼ vi ei @xj @xj Vector field : gradv ¼ @ @vj > > :$ v ¼ ei vj ej ¼ ei ej : front (left) form @xi @xi
ð1:371Þ
ð1:372Þ
1.17
Differentiation and Integration in Tensor Field
71
Second-order tensor field: 8 @ @Tij > > ek ¼ ei ej ek : rear (right) form < T $ ¼ Tij ei ej @xk @xk grad T ¼ > @ @Tjk > :$ T ¼ ei Tjk ej ek ¼ ei ej ek : front (left) form @xi @xi ð1:373Þ (2) Divergence Vector field: divv ¼ $ vð ¼ v $Þ ¼ vi ei
@ @vi ej ¼ @xi @xj
ð1:374Þ
Second-order tensor field: 8 @ @Tir > > ek ¼ ei : rear (right) form < T$ ¼ Tij ei ej @xr @xk divT ¼ > @ @Tri > : $T ¼ ei Tjk ej ek ¼ ei : front (left) form @xr @xi
ð1:375Þ
(3) Rotation (or curl) Vector field:
rotv ¼
8 @ @vi > > < v $ ¼ vi ei @x ej ¼ eijk @x ek : rear (right) form j
j
> > : $ v ¼ @ ei vj ej ¼ eijk @vj ek : front (left) form @xi @xi
ð1:376Þ
noting Eq. (1.35). Second-order tensor field: 8 @ > > T $ ¼ Tij ei ej ek > > > @xk > > > > > @Tij @Tij > > ¼ ei ðej ek Þ ¼ ejkr ei er : rear (right) form < @xk @xk rotT ¼ > @ > > ei Tjk ej ek $T¼ > > @x > i > > > > @Tjk @Tjk > > ¼ ðei ej Þ ek ¼ eijr er ek : front (left) form : @xi @xi ð1:377Þ
72
1
Mathematical Preliminaries: Vector and Tensor Analysis
The symbol $ is regarded as a vector, and the scalar product of itself, i.e. D r2 $ $ ¼
@ @ @2 er es ¼ @xr @xr @xr @xs
ð1:378Þ
has the meaning of r2 ð Þ divðgradð ÞÞ. The symbol D is called the Laplacian or Laplace operator, which is often used for scalar or vector fields as @2s @xr @xr
ð1:379Þ
@ 2 vs es @xr @xr
ð1:380Þ
Ds ¼ Dv ¼
The following relations hold between the above-mentioned operators. 9 gradðsvÞ ¼ v grads þ sgradv; > > > > divðsvÞ ¼ sdivv þ v grads; > > > > divðu vÞ ¼ v rotu u rotv; > > = rotðu vÞ ¼ ðgraduÞvðgradvÞu þ ðdivvÞuðdivuÞv; ð1:381Þ gradðu vÞ ¼ ðgradvÞu þ ðgraduÞv þ u rotv þ v rotu; > > > T > divðsTÞ ¼ T grads þ sdivT; > > > > > divðTvÞ ¼ v divT þ trðTgradvÞ > ; T divðABÞ ¼ B divA þ divA B (4) Gauss’ divergence theorem Consider the physical quantity TðxÞ in the zone surrounded by a smooth surface inside a material. Then, suppose the slender prism cut by the four planes perpendicular to the x2 -axis and x3 -axis in infinitesimal intervals from a zone inside the material. The following equation holds for the prism possessing the infinitesimal volume @v ¼ dx1 dx2 dx3 . Z
@T dv ¼ @x 1 @v
Z
@T xþ dx1 dx2 dx3 ¼ ½Tx11 dx2 dx3 ¼ ðT þ T Þdx2 dx3 @v @x1
ð1:382Þ
where ðÞ þ and ðÞ designate the values of physical quantity at the maximum and the minimum x1 -coordinates, respectively. The neighborhood of the surface cut by the prism is magnified in Fig. 1.6. Consider the infinitesimal rectangular surface PQRS of the prism exposed at the surface in the maximum x1 -coordinate and the infinitesimal rectangular section PQ R S cut by the plane passing through the point P and perpendicular to the x1
!
!
axis by the prism. Then, denoting QQ ¼ dxQ ; SS ¼ dxS , the vectors PQ; PS are given by
1.17
Differentiation and Integration in Tensor Field
73
S*
R* dx
dx2
1
R
dx s S
e3
n
dx3
e2
Q*
0
dx Q
Q
P
e1
Fig. 1.6 Infinitesimal square pillar cut from a zone in material
! PQ ¼ dx2 e2 þ dxQ e1 ;
!
PS ¼ dx3 e3 þ dxS e1
ð1:383Þ
and thus it holds that !
!
n þ da þ ¼ PQ PS ¼ ðdx2 e2 þ dxQ e1 Þ ðdx3 e3 þ dxS e1 Þ ¼ dx2 dx3 e2 e3 þ dxQ dx3 e1 e3 þ dxS dx2 e2 e1
ð1:384Þ
¼ dx2 dx3 e1 dxQ dx3 e2 dxS dx2 e3 Comparing the components in the base e1 on the both sides in Eq. (1.384), one has n1þ da þ ¼ dx2 dx3
ð1:385Þ
In a similar manner for the surface of the prism exposed on surface in the minimum x1 -coordinate, one has n 1 da ¼ dx2 dx3
ð1:386Þ
The general expression of projected area is given in Appendix A. Adopting Eqs. (1.385) and (1.386) in Eq. (1.382), it holds for the prism that Z @T þ þ þ dv ¼ T þ n1þ da þ T ðn ð1:387Þ 1 da Þ ¼ T n1 da þ T n1 da @v @x1 Then, the following equation holds for the whole zone. Z v
@T dv ¼ @x1
Z a
Tn1 da
ð1:388Þ
In a similar manner also for the x2 - and x3 -directions, the following Gauss’ divergence theorem holds.
74
1
Mathematical Preliminaries: Vector and Tensor Analysis
Z v
@T dv ¼ @xi
Z Tni da
ð1:389Þ
a
The following equations for the scalar s, the vector v and the tensor T hold from Eq. (1.389).
1.18
Ð Ð Ð Ð @s v @x dv ¼ a sni da; v $sdv ¼ a snda i
ð1:390Þ
Ð @vi Ð Ð Ð v @x dv ¼ a vi ni da; v $ vdv ¼ a v nda i
ð1:391Þ
Ð @Tij Ð Ð Ð v @x dv ¼ a Tij ni da; v $Tdv ¼ a Tnda i
ð1:392Þ
Representation in General Coordinate System
The physical meanings of finite strain (rate) tensors are captured clearly by their representations in the embedded (convected) coordinate system, since the components in the primary bases are kept constant. The embedded coordinate system changes to a curvilinear coordinate system even if it is the rectangular coordinate system at the initial state of deformation process. Then, mathematics in the general curvilinear coordinate system with the primary and the reciprocal base vectors, the representations and the coordinate transformations of vector and tensor, the metric tensor, the scalar, the vector and the tensor products, etc. are addressed in this section.
1.18.1
Primary and Reciprocal Base Vectors
The following reciprocal base vectors g1 ; g2 ; g3 are defined by the primary base vectors g1 ; g2 ; g3 as follows: gi ¼ 12 eijk
gj gk ; vg
i:e:
g1 ¼
g2 g3 2 g3 g1 3 g1 g2 ;g ¼ ;g ¼ ð1:393Þ ½g1 g2 g3 ½g1 g2 g3 ½g1 g2 g3
noting Eq. (1.59) and e1jk gj gk ¼ 2g2 g3 for example, where vg is the volume of parallelepiped made up of the primary vectors g1 ; g2 ; g3 in the right hand coordinate system, i.e.
1.18
Representation in General Coordinate System
vg ½g1 g2 g3
75
ð1:394Þ
noting Eq. (1.47). The inverse expression of Eq. (1.393) is given by 1 gi ¼ eijk vg g j gk ; 2
i:e:
g1 ¼
g2 g3 g3 g1 g1 g2 ; g ; g ¼ ¼ 2 3 ½g1 g2 g3 ½g1 g2 g3 ½g1 g2 g3 ð1:395Þ
noting Eq. (1.59), where it is satisfied noting Eq. (1.63)3 that ½g1 g2 g3 ½g1 g2 g3 ¼ 1
ð1:396Þ
1 ¼ ½g1 g2 g3 vg
ð1:397Þ
leading to
Here, note that the primary and the reciprocal basis in Eqs. (1.393) and (1.395) satisfy the following relations. gi g j ¼ dij ; gi gi ¼ 3
ð1:398Þ
gi gi ¼0
ð1:399Þ
which are equivalent to Eqs. (1.63)1; 2 and (1.65). Now, we consider the general curvilinear coordinate systems fhi g with the primary base fgi g and its reciprocal base fgi g. The infinitesimal line-element dx at the material point vector x is described as follows: ð1:400Þ dx ¼ dxi ei ¼ dhi gi where dxi ¼ dx ei ; dhi ¼dx gi
ð1:401Þ
by virtue of Eq. (1.62). It follows from Eq. (1.400) with Eq. (1.398) that dx dx ¼ dxi dxi ¼ dhi dh j gi gj
ð1:402Þ
Further, one has the expressions dx ¼
@x @x dxi ¼ i dhi @xi @h
ð1:403Þ
76
1
Mathematical Preliminaries: Vector and Tensor Analysis
Then, the following expressions are derived from Eqs. (1.400) and (1.403). ei ¼
@x ; @xi
gi ¼
@x @hi
ð1:404Þ
The reciprocal base vector gk is always perpendicular to the coordinate plane formed by the primary base vectors ðgi ; gj Þðk 6¼ i; k 6¼ jÞ. Note, however, that fgi g and fhi g are merely defined momentarily from the primary base vector fgig. Therefore, a continuous coordinate system fhi g with the reciprocal vector base fgi g does not exist in actuality when the coordinate system fhi g with the primary base fgig is a curvilinear coordinate system.
1.18.2
Metric Tensor and Base Vector Algebra
Now, we introduce the scalar quantities gij gi gj ; gij gi g j ¼ dij ; gij gi gj ¼ dij ; gij gi g j
ð1:405Þ
noting Eqs. (1.398), where the symmetries gij ¼ gji ; gij ¼ gji ; gij ¼ gij hold. Here, note that gij and gij are necessary to express the length of the line-element and the angle between line-elements which are most basic geometrical ingredients. Then, they are called the metric tensors. It follows from Eq. (1.405) that gir grj ¼ dij ;
½gir ½grj ¼ ½gir grj ¼ ½dij
ð1:406Þ
noting gir grj ¼ ðgi gr Þðgr g j Þ ¼ gi g j ¼ dij or gir grj ¼
@hr @h j @h j ¼ i ¼ dij @hi @hr @h
Hence, we have gij ¼ g1 ij ;
½gij ¼ ½gij 1
ð1:407Þ
1.18
Representation in General Coordinate System
77
from which it follows that detðgij Þ detðgij Þ ¼1 (no sum); detðgij Þ ¼ 1= detðgij Þ
ð1:408Þ
and g detðgij Þ ¼ ½g1 g2 g3 2 ¼ v2g ;
detðgij Þ ¼ 1=g ¼ 1=v2g
ð1:409Þ
noting Eqs. (1.17) or (1.143) 4 , (1.154) and (1.157). Equation (1.107) means that the matrices with the components gij and gij are inverse tensors to each other, while this property is based on their non-singularities, i.e., ½gij 6¼ 0 and ½gij 6¼ 0. By use of the notation in Eq. (1.405), the following transformation rules between the primary and the reciprocal base vectors hold. gi ¼ gir gr ;
gi ¼ gir gr
ð1:410Þ
It is known from Eq. (1.410) that gij and gij play the role of the shifter which makes the index go up and down, respectively. The generalized identity tensor in the general coordinate system is given from Eq. (1.72) with Eq. (1.405) as follows: gð¼ ei ei Þ ¼ gi gi ¼ gi gi ¼ gij gi gj ¼ gij gi g j ð¼ gT Þ
ð1:411Þ
fulfilling pffiffiffi jjgjj ¼ 3; detg ¼ 1; g ¼ gT ¼ g1
ð1:412Þ
The fact that the tensor g is the generalized identity tensor is confirmed noting Eq. (1.72) as follows: ( gv ¼ ( gT ¼
gi gi v ¼ ðv gi Þgi ¼ vi gi gi gi v ¼ ðv gi Þgi ¼ vi gi
) ¼v
gi gi Trs gr gs ¼ dir Trs gi gs ¼ Tis gi gs
gi gi Tr s gr gs ¼ dri Tr s gi gs ¼ Ti s gi gs
ð1:413Þ
) ¼T
ð1:414Þ
The components of g shown in Eq. (1.411) are necessary for the description of the geometrical elements, e.g., the length of line-element and the angle between line elements in the general space as will be shown in Subsection 1.18.3. Then, it is called the metric tensor. Further, Eqs. (1.393) and (1.395) are written as
78
1
Mathematical Preliminaries: Vector and Tensor Analysis
1 gi ¼ eijk gj gk ; 2
1 gi ¼ eijk g j gk 2
ð1:415Þ
and one has the following equations from Eqs. (1.394) and (1.397) ½gi gj gk ¼ eijk ;
½gi g j gk ¼ eijk
ð1:416Þ
eijk vg
ð1:417Þ
eijk ¼ gir gjs gkt erst
ð1:418Þ
introducing the symbols eijk eijk vg ; eijk which satisfy the following transformation rules. eijk ¼ gir gjs gkt erst ; noting eijk ¼ vg detðgpq Þeijk ¼ vg gir gjs gkt erst ¼ gir gjs gkt vg erst vg vg eijk ¼
1 1 1 detðgpq Þeijk ¼ gir gjs gkt erst ¼ gir gjs gkt erst vg vg vg
with the aid of Eqs. (1.15) and (1.409). The following relations hold by virtue of Eq. (1.21). 8 ijk > < e eirs ¼ ejki ersi ¼ djr dks djs dkr eijk eijs ¼ 2dks > : ijk e eijk ¼ 3! ¼ 6
ð1:419Þ
Further, the vector products of the both primary bases or reciprocal bases are given by gi gj ¼ eijk gk ; gi g j ¼ eijk gk
ð1:420Þ
noting 8 1 1 1 > < gi gj ¼ ðgi gj gj gi Þ ¼ ðdir djs dis djr Þgr gs ¼ ekij ekrs gr gs ¼ ekij gk 2 2 2 > : gi g j ¼ 1 ðgi g j g j gi Þ ¼ 1 ðd d d d Þgr gs ¼ 1 ekij e gr gs ¼ ekij g ir js is jr krs k 2 2 2
by virtue of Eqs. (1.39), (1.415) and (1.419).
1.18
Representation in General Coordinate System
1.18.3
79
Tensor Representations
The vector is described by Eq. (1.62) as v ¼ ðv gi Þgi ¼ ðv gi Þgi ð¼ gi gi v ¼ gi gi vÞ
ð1:421Þ
v ¼ vi g i ¼ v i g i
ð1:422Þ
leading to
with v i ¼ v gi ;
vi ¼v gi
ð1:423Þ
where vi and vi are called the contravriant component and the covariant component, respectively. The relation of vi and vi are given as vi ¼gir vr ; vi ¼gir vr
ð1:424Þ
noting v i ¼ v r gr gi ;
vi ¼ v r gr gi
The metric tensor component plays the role of shifter as known from Eq. (1.424). The scalar, the vector and the tensor products are given by Eqs. (1.420) and (1.422) as u v ¼ ui vi ¼ ui vi ¼ gij ui v j ¼ gij ui vj jjvjj ¼
qffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffi v v ¼ vi vi ¼ gij vi v j ¼ gij vi vj
ð1:425Þ ð1:426Þ
u v ¼ eijk ui v j gk ¼ eijk ui vj gk
ð1:427Þ
ðu vÞ w ¼ eijk ui v j wk ¼ eijk ui vj wk
ð1:428Þ
u v ¼ ui v j gi gj ¼ ui vj gi g j ¼ ui v j gi gj ¼ ui vj gi g j
ð1:429Þ
The square of the length of an infinitesimal line-element vector dx is expressed using the metric tensor by Eq. (1.426) as follows: jjdxjj2 ¼ dx dx ¼ gij dhi dh j ¼ gij dhi dhj ¼ dhi dhi
ð1:430Þ
80
1
Mathematical Preliminaries: Vector and Tensor Analysis
noting dx dx ¼ dhi gi dh j gj ¼ dhi gi dhj g j ¼ dhi gi dhj g j The infinitesimal surface vector dak of the surface formed by the line-elements dhi gi (no sum) and dh j gj (no sum) along the primary base fgi g is given from Eq. (1.420) by dak ¼ dhi gi dh j gj ¼ eijk dhi dh j gk
ð1:431Þ
(no sum)
where i; j; k are different from each other and are ordered in an even permutation, which is called the surface element vector. The volume dV of the infinitesimal parallelepiped formed by the sides along the primary base fgi g is given by dv ¼ ½dh1 g1 dh2 g2 dh3 g3 leading to dv ¼ ½g1 g2 g3 dh1 dh2 dh3 ¼ vg dh1 dh2 dh3
ð1:432Þ
Following the tensor product of two vectors in Eq. (1.429), the second-order tensor T is defined as follows:
T ¼ T ij gi gj ¼ Tij gi g j ¼ Ti j gi gj ¼ Tij gi g j
ð1:433Þ
where the components of T are derived from as follows: T ij ¼ gi Tg j ; Tji ¼ gi Tgj ; Tij ¼ gi Tg j ; Tij ¼ gi Tgj
ð1:434Þ
exploiting Eq. (1.398)1 . Here, we cannot write Tji because of Tij 6¼ Ti j . Then, the relations between the components in the different variance are given by substituting Eq. (1.433) into Eq. (1.434) as follows: (
i rj g ¼ gir Tr j ¼ gir Trs gsj ; T ij ¼ Tr
Ti j ¼ gir T rj ¼ gir Trs gsj ¼ Tir grj ;
Tij ¼ T ir grj ¼ gir Tr s gsj ¼ gir Trj
Tij ¼ gir T rs gsj ¼ gir Trj ¼ Ti r grj
ð1:435Þ
Analogously to the vector described in the foregoing, T ij , Tij , Ti j and Tij are called the contravariant, the contravariant-covariant, the covariant-contravariant and the covariant components, respectively. The scalar product of second-order tensors A and B is expressed in the components as follows:
1.18
Representation in General Coordinate System
81
j i ij ij rs js A : B ¼ Aij Bij ¼ Aij Bj i ¼ Ai Bj ¼ Aij B ¼ gir gjs A B ¼ gir g Aij Brs
ð1:436Þ
noting 8 ij ðA gi gj Þ:ðBrs gr gs Þ ¼ ðgi gr Þðgj gs ÞAij Brs > > > > i j s r j i s > > r > > ðAj gi g Þ:ðBr g gs Þ ¼ ðgi g Þðg gs ÞAj Br > > < ðAj gi g Þ:ðBr g gs Þ ¼ ðgi g Þðg gs ÞAj Br s r j r j i i s i j rs i s > ðAij g g Þ:ðB gr gs Þ ¼ ðg gr Þðgj g ÞAij Brs > > > > > > ðAij gi gj Þ:ðBrs gr gs Þ ¼ gir gjs Aij Brs > > > : ðAij gi g j Þ:ðBrs gr gs Þ ¼ gir gjs Aij Brs : The magnitude of tensor T is given by setting A ¼ B ¼ T in Eq. (1.436) as qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffi jjTjj ¼ T : T ¼ T ij Tij ¼ T ij Ti j ¼ gir gjs T ij T rs
ð1:437Þ
The definition of the trace trT T : I in the Cartesian space is extended as the scalar product with the metric tensor in the Riemann space, leading to the sum of the diagonal components of the mixed variant form, i.e.
trT T : g ¼ g : T ¼ T ii ¼ Tii ¼ Ti i ¼ Tii
ð1:438Þ
noting 8 i ðg gi Þ:ðT rs gr gs Þ ¼ T ii > > > > < ðgi g Þ:ðT r g gs Þ ¼ T i s r i i s r i > ðgi g Þ:ðTr g gs Þ ¼ Tii > > > : ðgi gi Þ:ðTrs gr gs Þ ¼ Tii where use is made of Eq. (1.120)7 . The following relations hold. 8 i g ðT rs gr gs Þg j ¼ T ij ¼ g j ðT rs gs gr Þgi > > > > < gi ðT r g gs Þg ¼ T i ¼ g ðT r g gr Þgi j j s r j s s > gi ðTrs gr gs Þg j ¼ Tij ¼ g j ðTrs gs gr Þgi > > > : i g ðTrs gr gs Þg j ¼ Trs ¼ g j ðTrs gs gr Þgi Comparing these relations with Eq. (1.130), the transposed tensor TT is given by
82
1
Mathematical Preliminaries: Vector and Tensor Analysis
TT ¼T ji gi gj ¼ Tj i gi g j ¼ Tji gi gj ¼ Tji gi g j
ð1:439Þ
Therefore, the transposed tensor is given by exchange of the front and the rear base vectors. However, it is not obtained by the exchange of front and rear suffices in components, differing from the tensor expressions in the Cartesian coordinate system, as known from ( T ¼ T
Tj i g i g j
)
( 6¼
Tji gi gj
Tji gi g j
Tj i g i g j
ð1:440Þ
If T is the symmetric tensor, it follows from Eqs. (1.433) and (1.439) that
T ij ¼ T ji ; Tij ¼ Tj i ðTi j ¼ Tji Þ;
Tij ¼ Tji
ð1:441Þ
However, the indices cannot be exchanged between the contravariant and the covariant ones, i.e.
Tij 6¼ Tji ; Ti j 6¼ Tj i
ð1:442Þ
Therefore, the mixed components Tij and Tji possess the nine independent values even in the symmetric tensor, leading to the inconvenience of analysis.
Chapter 2
Description of Motion
The tensor analysis providing the mathematical foundation for the continuum mechanics was described in Chap. 1. Basic concepts and quantities for continuum mechanics will be studied in the four chapters up to Chap. 6. The appropriate descriptions of motion of basic material elements are of importance in the formulation of constitutive relations. The variations of some basic material elements during the deformation will be studied in this chapter prior to the explanation of the deformation/rotation measures in the subsequent chapter.
2.1
Motion of Material Point
A material body is assembly of material particles (or material elements). The map of positions of material particles in a space is referred to as the configuration. Here, the configurations in the initial time t ¼ t0 and the current time t are called the initial (or Lagrangian) configuration and the current (or Eulerian) configuration, respectively. Deformation is described by the change of configuration from a particular configuration which is called the reference configuration. Here, the reference configuration can be chosen at arbitrary intermediate time sðt0 s tÞ which is called the reference time. The position vectors of material particle in the initial and the current configurations are designated by X and xðtÞ, respectively. Here, X is fixed and thus it can be regarded as a label of each material particle. The motion of material point during the time t0 ! t is described as x ¼ vðX; tÞ;
X ¼ v1 ðx; tÞ
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_2
ð2:1Þ
83
84
2
Description of Motion
Incidentally, the motion of material point during the time t0 ! s is described as xðsÞ ¼ vðX; sÞ;
X ¼ v1 ðxðsÞ; sÞ
ð2:2Þ
The fact that a material point does not overlap or separate by the motion of material requires the existence of the one-to-one correspondence between X and x (x is uniquely determined by X and vice versa) so that x1 ðX1 ; X2 ; X3 Þ; x2 ðX1 ; X2 ; X3 Þ and x3 ðX1 ; X2 ; X3 Þ must be mutually independent. Now, let the mathematical requirement for this fact be derived by the reductive absurdity. x1 ; x2 ; x3 are not mutually independent if they satisfy the constraint f ðx1 ðX1 ; X2 ; X3 Þ; x2 ðX1 ; X2 ; X3 Þ; x3 ðX1 ; X2 ; X3 ÞÞ ¼ 0
ð2:3Þ
from which it follows that 9 2 @x @f @x1 @f @x2 @f @x3 1 > þ þ ¼ 0> > > 6 @X1 @x1 @X1 @x2 @X1 @x3 @X1 > = 6 @x @f @x1 @f @x2 @f @x3 6 1 þ þ ¼ 0 ; i.e:6 > 6 @X2 @x1 @X2 @x2 @X2 @x3 @X2 > > 4 @x > @f @x1 @f @x2 @f @x3 > 1 ; þ þ ¼0 @X3 @x1 @X3 @x2 @X3 @x3 @X3
@x2 @X1 @x2 @X2 @x2 @X3
9 8 9 8 @x3 3> @f > > 0 > > > > > > > > > > > > @X1 7 @x1 > > > > > = < = < 7 @x3 7 @f ¼ 0 7 > > @X2 7> > > > > @x2 > > > > > > > > @x3 5> @f > > > ; ; : > : 0 @X3 @x3 ð2:4Þ
The equation J¼0
ð2:5Þ
must hold in order that @f =@x1 ; @f =@x2 ; @f =@x3 are determined uniquely on account of Eq. (1.169) regarding TiJ ¼ @xi =@XJ and vi ¼ @f =@xi , where J is defined by @xi @x1 @x2 @x3 ¼ eIJK J det @XJ @XI @XJ @XK
ð2:6Þ
and is called the functional determinant or Jacobian. In contrast, in order that they are mutually independent, it must hold that J 6¼ 0
ð2:7Þ
being free from the constraint in Eq. (2.3). The transformation between x and X is called the admissible transformation, if f1 ; f2 ; f3 in x1 ¼ f1 ðX1 ; X2 ; X3 Þ; x2 ¼ f2 ðX1 ; X2 ; X3 Þ, x3 ¼ f3 ðX1 X2 X3 Þ are single-valued and continuous functions, so that the Jacobian is not zero as shown in Eq. (2.7). Further, if the Jacobian is positive, a right-hand coordinate system is transformed to other right-hand one, and it is called the positive transformation. Inversely, if the Jacobian is negative, a
2.1 Motion of Material Point
85
right-hand coordinate system is transformed to a left-hand one, and it is called the negative transformation. Admissible and positive transformation with J [ 0 is assumed throughout this book. Physical quantity, say T, in the body changes generally with the position and the time. Physical quantity at current time is described TðX; tÞ in terms of the reference configuration X and the current time t. This type of description of mechanical state is called the Lagrangian (or material) description. On the other hand, the physical quantity at current time is described Tðx; tÞ in terms of the current configuration x and the current time t. This type of description of physical quantity is called the Eulerian description or spatial description. Further, the physical quantity at current time can be described in terms of the configuration xðsÞ at arbitrary reference time s and the current time t as T v1 ðxðsÞ; sÞ; t ¼ s TðxðsÞ; tÞ
ð2:8Þ
() designating to choose the reference time s as s [ t0 . Equation (2.8) is called the relative description. Specifically, t TðxðtÞ; tÞ is called the updated Lagrangian description, where the reference configuration is taken as the current configuration, choosing the reference time as the current time, i.e. s ¼ t. In contrast, the description TðX; tÞ will be called the total Lagrangian description. s
2.2
Time-Derivatives
The time-derivative of the tensor in the spatial description @Tðx; tÞ @t
ð2:9Þ
describes the rate of the physical quantity at a certain spatial point and thus it is called the spatial-time (or local) derivative. In many cases of fluid mechanics, a motion and its history of individual particle from the initial state is immaterial and thus the spatial-time derivative is often adopted. In contrast, the time-derivative of the tensor in the material description @TðX; tÞ @t
ð2:10Þ
describes the rate of the physical quantity in a certain material particle and thus is called the material-time derivative. It is denoted by the symbol
T
@TðX; tÞ DT @TðX; tÞ or @t Dt @t
ð2:11Þ
86
2
Description of Motion
In solid mechanics, the rate of deformation and its history of individual material particle is required and thus the material-time derivative is used usually. The material-time derivative in Eq. (2.11) and the spatial-time derivative in Eq. (2.9) are related by
T
@Tðx; tÞ @Tðx; tÞ @Tij ðx; tÞ @Tij ðx; tÞ þ v ; Tij þ vk @t @x @t @xk
ð2:12Þ
where v @x=@t is the velocity vector of material particle. The first term in the right-hand side signifies the non-steady (or local time derivative) term describing the change with time at fixed spatial point and the second term signifies the steady (or convective) term describing the change due to the movement of material, which results from the existence of a spatial gradient of the physical quantity T. Rate-type constitutive equations for the irreversible deformation of solids, e.g. the viscoelastic, the elastoplastic and the viscoplastic deformation, must be described by the material-time derivative pursuing a material particle because they must describe the relation of physical quantities in each material particle. Here, it should be noticed that the material-time derivative of physical quantity describes the rate observed by moving in parallel with material particle as known from Eq. (2.11) which concerns only with the position vector of material particle and the time. Then, the objective time-derivative based on the rate of physical quantity observed by the coordinate system deforming/rotating with a material must be used for constitutive equations of solids as will be described in Chap. 6.
2.3
Variations and Rates of Geometrical Elements
Variations of line, surface and volume elements and their rates during the deformation are described in this section.
2.3.1
Deformation Gradient and Variations of Line, Surface and Volume Elements
Variations of line, surface and volume elements are described below. Relation of the current infinitesimal line element dx to the initial infinitesimal line element dX is represented as
2.3 Variations and Rates of Geometrical Elements
87
dx ¼ FdX; dX ¼ F1 dx
ð2:13Þ
where F
@x @X
ð2:14Þ
which is referred to as the deformation gradient tensor, for which it follows that 8 @x @xi > > ei ej ¼ F¼ > > @X @Xj > > T > > > T @x T @xi @xi @xj > > ¼ ei ej ¼ ej ei ¼ ei ej
> F1 ¼ ei ej ¼ > > > @xj @x > > T > > @X T @Xi @Xj > > : FT ¼ ¼ ei ej ¼ ei ej @xj @xi @x
ð2:15Þ
noting FF1 ¼
@xi @Xr @xi @Xj ei ej er es ¼ ei es ¼ dis ei es ¼ I @Xj @xs @Xj @xs
Equation (2.15) is represented in the fixed coordinate system, while it is represented concisely by the embedded base vectors in the initial and the current configurations in the embedded coordinate system as will be described in the next chapter. The following expression holds by introducing the polar decomposition for the deformation gradient tensor in Eq. (1.273) into Eq. (2.13). dx ¼ RUdX ¼ VRdX;
dX ¼ U1 RT dx ¼ RT V1 dx
ð2:16Þ
The current position vector is related to the reference position vector by use of the displacement vector u as x ¼ X þ u ¼ ðXi þ ui Þei
ð2:17Þ
Then, the deformation gradient is represented as follows: @ðX þ uÞ @u F¼ ¼ Iþ ¼ @X @X
@ui ei ej dij þ @Xj
ð2:18Þ
88
2
Description of Motion
from which it follows that dF ¼
@du @dui ¼ ei ej @Xj @X
ð2:19Þ
The current and the reference infinitesimal volume elements dv and dV formed by the infinitesimal line-elements dxa; dxb; dxc and dXa; dXb; dXc are related by dv ¼ dxa dxb dxc ¼ FdXa FdXb FdXc ¼ det F dXa dXb dXc ¼ ðdet FÞdV ð2:20Þ by virtue of Eq. (1.254)3 . Then, the ratio of the current to the reference infinitesimal volume elements, i.e., Jacobian is given by J det F ¼
m q0 ¼ q V
ð2:21Þ
where q0 and q are the mass densities in the reference and the current configurations, respectively. The following equalities hold for the infinitesimal volume elements, designating the infinitesimal reference and current surface element vectors as dA ¼ dXa dXb and da ¼ dxa dxb , respectively, noting Eq. (1.139). dv ¼
da dxc JdV ¼ JdA dXc ¼ JdA F1 dxc ¼ JFT dA dxc
ð2:22Þ
where the area vectors da and dA in the current and the reference configurations are given by da ¼ dan;
dA ¼ dAN
ð2:23Þ
designating the infinitesimal area and the unit outward-normal as ðda; nÞ and ðdA; NÞ in the current and the reference configurations, respectively. The following relations are obtained from Eq. (2.22). da ¼ JFT dA ¼ ðcofFÞdA;
dA ¼ FT da=J ¼ ðcofFÞ1 da
ð2:24Þ
noting cofF ¼ JFT
ð2:25Þ
2.3 Variations and Rates of Geometrical Elements
89
by virtue of Eq. (1.160), leading to da ¼ FT N nJdA;
dA ¼ FT n Nda=J
ð2:26Þ
n ¼ FT NJdA=da;
N ¼ FT nda=ðJdAÞ
ð2:27Þ
Equation (2.26) is referred to as the Nanson’s formula. Here, the following Euler’s formula holds for the cofactor. rX ðcofFÞ ¼
@cofF ¼ 0; @X
@ðcofFÞap @Xp
¼
@ 12 eabc epqr Fbq Fcr ¼0 @Xp
ð2:28Þ
noting Eq. (1.13).
2.3.2
Velocity Gradient and Rates of Line, Surface and Volume Elements
Rates of line, surface and volume elements are described below. Differentiating Eq. (2.13), we have
ðdxÞ ¼ F dx;
ðdxÞ ¼ dv ¼
@v dx ¼ ldx @x
ð2:29Þ
where l gradv ¼ v rx ¼
@v 1 ¼ FF @x
ð2:30Þ
noting
@v @ x @ x @X ¼ ¼ @x @x @X @x
ð2:31Þ
The tensor l is called the velocity gradient tensor. The following expression holds noting FF1 ¼ F F1 þ F F1 ¼ l þ F F1 ¼ O with the caution ðF1 Þ ¼
F1 F F1 6¼ ðFÞ1 ¼ F1 ½ðF1 Þ 1 F1 that F F1 ¼ l
ð2:32Þ
90
2
Description of Motion
It follows from Eqs. (2.21) and (2.30) that
trl ¼ divv ¼ J =J
ð2:33Þ
J ¼ Jtrl
ð2:34Þ
ðdvÞ ¼ ð J =JÞdv ¼ J dV ¼ ðtrlÞdv
ð2:35Þ
noting
J ¼ ðdet FÞ ¼
@ det F : F ¼ ðdet FÞFT : F ¼ ðdet FÞtr F1 F ¼ ðdet FÞtr F F1 ¼ Jtrl @F
or
trl ¼
@vr @v1 @v2 @v3 @v1 @x2 @x3 þ @x1 @v2 @x3 þ @x1 @x2 @v3 ð@x1 @x2 @x3 Þ v ¼ þ þ ¼ ¼ ¼ @xr @x1 @x2 @x3 @x1 @x2 @x3 @x1 @x2 @x3 v
with Eqs. (1.163). Further, it follows from Eq. (2.25) that ðcofFÞ ¼ ~lcofF
ð2:36Þ
~l ðtrlÞI lT
ð2:37Þ
where
noting
ðcofFÞ ¼ J F
T
þF
T
J J J ¼ cofF þ FT FT cofF ¼ cofF FT FT cofF J J
with the use of Eq. (1.160). ~l is referred to as the surface strain rate tensor. Further, differentiating Eq. (2.24) and noting Eq. (2.36), we have ðdaÞ ¼ ðcofFÞ dA ¼ ~lðcofFÞdA which is rewritten as ðdaÞ ¼ elda in the spatial description.
ð2:38Þ
2.3 Variations and Rates of Geometrical Elements
91
Now, we express the surface vectors as follows: da nda;
dA NdA
ð2:39Þ
where da and dA are the infinitesimal areas, and n and N are the unit normal vectors of the current and the reference infinitesimal surface vectors, respectively. Here, noting n n ¼ 0 from n n ¼ 1 for the unit vector n, it follows that h i ðdaÞ ¼ n nðdaÞ ¼ n ðndaÞ n da ¼ n ðdaÞ
ð2:40Þ
Substituting Eq. (2.38) into Eq. (2.40), one obtains the rate of the current infinitesimal area as follows: ðdaÞ ¼ n elda
ð2:41Þ
ðdaÞ ¼ ðtrl n dnÞda
ð2:42Þ
d sym½l
ð2:43Þ
or
where
noting n wn ¼ 0 because of Eq. (1.180), where w ant½l
ð2:44Þ
Here, d and w are called the strain rate tensor and the continuum spin tensor, respectively. Further, the rate of the unit normal of the current surface element is given from Eqs. (2.37), (2.39), (2.41) and (2.42) as follows: n ¼ ðn dnÞI lT n ¼ ðn lnÞI lT n
ð2:45Þ
noting
n da ¼ ðndaÞ nðdaÞ ¼ ðtrlÞI lT nda nðtrl n dnÞda ¼ lT nda þ n dnda The formula for the physical quantities on the geometrical variation of material formulated in this section will be often used in the subsequent chapters.
92
2
2.4
Description of Motion
Material-Time Derivative of Volume Integration
Supposing that the zone of material occupying the volume v at the current moment ðt ¼ tÞ changes to occupy the volume v þ dv after the infinitesimal time R ðt ¼ t þ dtÞ, the material-time-derivative of the volume integration v Tðx; tÞdv of the physical quantity Tðx; tÞ involved in the volume is given by the following equation. Z
Tðx; tÞdv
1 dt!0 dt
¼ lim
v
1 dt!0 dt
Z
Z v þ dv
Tðx; t þ dtÞdv
Z
Tðx; tÞdv v
Z fTðx; t þ dtÞ Tðx; tÞgdv þ
¼ lim
Tðx; t þ dtÞdv
ð2:46Þ
dv
v
The integration of the first term in the right-hand side in Eq. (2.46) is transformed as 1 lim dt!0 dt
Z
Z ½Tðx; t þ dtÞ Tðx; tÞdv ¼
v
v
@Tðx; tÞ dv @t
ð2:47Þ
On the other hand, the second term in Eq. (2.46) describes the influence caused by the change of volume during the infinitesimal time increment. Here, the volume increment dv is given by subtracting the volume flowing out from the boundary of the zone from the volume flowing into the boundary, which is the sum of dvð¼ v ndadtÞ over the whole boundary surface (Fig. 2.1). Therefore, substituting the Gauss’ divergence theorem in Eq. (1.389) and ignoring the second-order infinitesimal quantity, the integration of the second term in the right-hand side of Eq. (2.46) is given by
Fig. 2.1 Translation of a zone in material
v t n
a da
n
v
v v
2.4 Material-Time Derivative of Volume Integration
93
Z 1 Tðx; t þ dtÞdv ffi lim Tðx; tÞdv dt!0 dt dv dv Z Z 1 Tðx; tÞvr nr dadt ¼ Tðx; tÞvr nr da ¼ lim dt!0 dt a a Z Z Z @Tðx; tÞvr @Tðx; tÞ @vr dv ¼ vr dv þ Tðx; tÞ dv ¼ @xr @xr @xr v v v
1 lim dt!0 dt
Z
The sum of the first term in the right-hand side in this equation and Eq. (2.47) is equal to the material-time derivative of Tðx; tÞ by virtue of Eq. (2.12). Then, Eq. (2.46) is given by
R
v
Tðx; tÞdv
¼
R
@vr dv ¼ T ðx; tÞ þ Tðx; tÞ v @xr
R
v
½T ðx; tÞ þ Tðx; tÞdiv v dv ð2:48Þ
which is called the Reynolds’ transportation theorem. Equation (2.48) can be obtained also by the following simple manner. Z
Tðx; tÞdv
Z
¼
TðX; tÞJdV
v
Z ¼
V
_ ðTðX; tÞJ þ TðX; tÞ JÞdV
v
Z @vr dv ¼ Tðx; tÞ þ Tðx; tÞ @xr v where V is the initial volume, while Eq. (2.34) is used. For the physical quantity T kept constant in a volume element, say a mass, Eq. (2.48) leads to Z ðT ðx; tÞ þ Tðx; tÞdivvÞdv ¼ 0 ð2:49Þ v
The local (weak) form of Eq. (2.49) is given as
T ðx; tÞ þ Tðx; tÞdivv ¼ 0
ð2:50Þ
Chapter 3
Description of Tensor (Rate) in Convected Coordinate System
Material involved in a unit mass in the current state (configuration) changes by the deformation. Therefore, constitutive equation describing the inherent property of material cannot be formulated by tensor variables in the current configuration, i.e. the spatial, i.e. Eulerian tensors but must be formulated by tensor variables in the reference (initial) configuration, i.e. the material, i.e. Lagragian tensors. Needless to say, however, variables which one can actually observe are the one in the current configuration. Therefore, the transformation of tensor variables from the current to the reference configuration, called the pull-back operation, and it inverse transformation, called the push-forward operation, are required in the formulation of constitutive equation and the deformation analysis. These operations can be clearly interpreted by the description of tensors in the convected (embedded) coordinate system in which the coordinate axes are caved inside the material itself. The description of tensors in the convected coordinate system and the pull-back, push-forward operations are described comprehensively in this chapter. The convected derivatives, the corotational rates and the objectivity of rate tensors are also described in detail.
3.1
Reference and Current Primary and Reciprocal Base Vectors
Suppose the coordinate system Hi with the primary base fGi g ði ¼ 1; 2; 3Þ, possessing the origin P0 at the fixed material point, which are embedded inside the material itself at the reference time t0 , changes to the coordinate system hi with the primary base fgi ðtÞg at the current time t under the deformation. Here, it should be noted that hi ¼ Hi holds even during the deformation since they are the components in the embedded coordinate system, although the primary base fgi ðtÞg © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_3
95
96
3
Description of Tensor (Rate) in Convected Coordinate System
changes. The coordinate system Hi (and hi ðtÞ ) is called the convective (convected or embedded) coordinate system and the primary base fGi g (and fgi ðtÞg) is called the convective (convected or embedded) base. On the other hand, the reciprocal bases in the reference configuration and in the current configuration are denoted by Gi and fgi ðtÞg, respectively, and their components are designated as fHi g and fhi ðtÞg, respectively. Here, it should be noted that the reciprocal bases and coordinates must be reformed at every moment with a deformation in order to keep the mathematical relation in Eq. (1.393) to the primary base so that they cannot be embedded. Here, the covariant components change in general as known from dhi ðtÞ ¼ gij ðtÞdh j ¼ gij ðtÞdH j 6¼ Gij dH j ¼ dHi noting Eq. (1.424) because of gij ðtÞ 6¼ Gij . Consequently, it follows that hi ¼ H i ;
dhi ¼ dHi ðhi ðtÞ 6¼ Hi ;
dhi ðtÞ 6¼ dHi Þ
ð3:1Þ
The following relations hold for the primary and the reciprocal base vectors by virtue of Eq. (1.405)2,3. Gi hG j ¼ dij ; Gi ¼ Gir Gr ;
gi ðtÞhg j ðtÞ ¼ dij
Gi ¼ Gir Gr
gi ðtÞ ¼ gir ðtÞgr ðtÞ;
gi ðtÞ ¼ gir ðtÞgr ðtÞ
ð3:2Þ ð3:3Þ
where Gij ¼ Gi hGj ;
Gij ¼ Gi hG j
gij ðtÞ ¼ gi ðtÞhgj ðtÞ;
gij ðtÞ ¼ gi ðtÞhgJ ðtÞ
ð3:4Þ
The following tensors are nothing else the metric tensors defined by Eq. (1.411). (
G Gi Gi ¼ Gi Gj ¼ Gij Gi Gj ¼ Gij Gi G j ¼ GT gðtÞ gi ðtÞ gi ðtÞ ¼ gi ðtÞ gi ðtÞ ¼ gij ðtÞgi ðtÞ gj ðtÞ ¼ gij ðtÞgi ðtÞ g j ðtÞð¼ gT ðtÞÞ
ð3:5Þ The vector and the tensor based in the reference and the current configuration are called the Lagrangian and the Eulerian vector and tensor, respectively. In principle, the Lagrangian and the Eulerian vector and tensor are denoted by the uppercase and the lowercase letters, respectively, throughout this book. Further, the tensor based in both of the reference and the current configurations is called the Lagrangian– Eulerian or Eulerian–Lagrangian two-point tensor and denoted by the uppercase letter throughout this book.
3.2 Description of Deformation Gradient Tensor by Embedded Base Vectors
3.2
97
Description of Deformation Gradient Tensor by Embedded Base Vectors
The infinitesimal line-element dX in the reference configuration and dxðtÞ in the current configuration are described as follows:
dX ¼ dHi Gi ¼ dhi Gi ¼ dHi Gi 6¼ dhi Gi dxðtÞ ¼ dhi gi ðtÞ ¼ dHi gi ðtÞð¼ dhi gi ðtÞ 6¼ dHi gi ðtÞÞ
ð3:6Þ
where
dHi ¼ dXhGi dhi ¼ dxðtÞhgi ðtÞ
ð3:7Þ
noting Eq. (1.423). The relation of the primary base vectors gi ðtÞ and Gi is given by replacing dxðtÞ and dX, respectively, in Eq. (2.13) as follows: gi ðtÞ ¼ FðtÞGi Gi ¼ F1 ðtÞgi ðtÞ
ð3:8Þ
or noting 8 @xðtÞ @xðtÞ @xðtÞ @X > > ¼ ¼ < gi ðtÞ ¼ @X @Hi @Hi @hi @X @X @X @xðtÞ > > ¼ ¼ : Gi ¼ @Hi @hi @xðtÞ @hi by of virtue of Eq. (1.404). Then, it follows from Eq. (3.8) that @hi @hi FðtÞ ¼ gi ðtÞ Gi ¼ Fi j gi ðtÞ G j ; Fi j ¼ dij ¼ ¼ @H j @h j i i @Hi @hi F1 ðtÞ ¼ Gi gi ðtÞ ¼ F1 j Gi g j ðtÞ; F1 j ¼ dij ¼ ¼ @h j @h j noting (
FðtÞ ¼ FðtÞG ¼ FðtÞGi Gi F1 ðtÞ ¼ F1 ðtÞg ¼ F1 ðtÞgi ðtÞ gi ðtÞ
F i ¼ gi ðtÞhFðtÞGj ¼ gi ðtÞhðgr ðtÞ Gr ÞGj ¼ dir drj j1 i F j ðtÞ ¼ Gi hF1 ðtÞgj ðtÞ ¼ Gi hðGr gr ðtÞÞgj ðtÞ ¼ dir dtj
ð3:9Þ
98
3
Description of Tensor (Rate) in Convected Coordinate System
with Eqs. (3.1) and (3.8). The fact that the deformation gradient is the Eulerian– Lagrangian two point tensor based in the initial and the current base vectors can be understood by the representation in the embedded coordinate system, i.e. Equation (3.9), although this fact is hidden in the rectangular coordinate system. Further, it is noteworthy that the deformation gradient tensor is regarded as the identity tensor possessing the Kronecker’s delta dij in the convected coordinate system with both the covariant and the contravariant bases as known from Eq. (3.9), although it is represented with the components @xi =@Xj in the fixed coordinate system as seen in Eq. (2.15). Here, it follows from Eq. (3.9) that i i FT ðtÞ ¼ Gi gi ðtÞ ¼ FT j G j gi ðtÞ; FT j ¼ dij j j FT ðtÞ ¼ gi ðtÞ Gi ¼ FT i gi ðtÞ Gj ; FT i ¼ dij
ð3:10Þ
from which one has gi ðtÞ ¼ FT ðtÞGi
ð3:11Þ
Gi ¼ FT ðtÞgi ðtÞ for the reciprocal base vectors, noting (
gi ðtÞ ¼ g j ðtÞdij ¼ g j ðtÞ Gj Gi ¼ FT ðtÞGi Gi ¼ G j dij ¼ G j gj ðtÞ gi ðtÞ ¼ FT ðtÞgi ðtÞ
The symbol ðtÞ designating the time-dependence will be eliminated for the sake of simplicity in the following. The rates of the current base vectors are given in terms of the velocity gradient tensor l in Eq. (2.30) from Eqs. (3.8) and (3.11) as follows: h
gi ¼ F Gi ¼ lgi gi ¼ FT Gi ¼ lT gi
F ¼ g Gi ¼ lF;
ð3:12Þ
FT ¼ Gi gi ¼ FT lT
ð3:13Þ
i 1
F ¼ Gi
gi
1
¼ F l;
F
T
¼
gi
T
Gi ¼ l F T
3.2 Description of Deformation Gradient Tensor by Embedded Base Vectors
99
noting 8 h > 1 > < F F FGi ¼ lgi T
T T i T T T i T T i > > F G ¼ F F F G ¼ F F g ¼ F F1 gi ¼ F F1 gi ¼ lT gi :
Further, it follows from Eqs. (3.9) and (3.12) that l ¼ gi gi
ð3:14Þ
lT ¼ gi gi noting
lgi gi ¼ gi gi
3.3
Pull-Back and Push-Forward Operations
The Eulerian tensor defined by the current base fgi g or fgi g is transformed to the Lagrangian tensor by replacing these bases to the reference base fGi g or Gi , keeping the conponents. This operation is called the pull-back and its inverse operation is called the push-forward. These operations can be performed also by the multiplication of the deformation gradient tensor because the current and the reference base vectors are related by the deformation gradient tensor as shown in Eqs. (3.8) and (3.11). The Eulerian vector v is expressed by the following contravariant and the covariant forms in the current configuration by Eq. (1.422), i.e. v ¼ v i gi ¼ v i gi
ð3:15Þ
The pull-back transformations from the Eulerian vector to the Lagrangian vectors are performed by replacing the current base fgi g or fgi g to the reference base fGi g or Gi and designated as follows: v G vi Gi ;
v
G
vi G i
ð3:16Þ
where the over arrow directed to the left ð Þ is added (the right arrow ð!Þ for the push-forward operation described later). Further, the uppercase index G is added in order to stipulate the replacement of the current base to the reference base (the lowercase index g for the push-forward operation described later) and it is put in a
100
3
Description of Tensor (Rate) in Convected Coordinate System
lower or upper position of suffix of the transposed base vector. These symbols for the pull-back and the push-forward operations have been introduced by Hashiguchi (2011). Equation (3.16) is rewritten as v G ¼ vi F1 gi ¼ F1 vi gi ;
v
G
¼ v i F T gi ¼ F T v i gi
using Eqs. (3.8) and (3.11), and thus the following direct notations are obtained. v G ¼ vi Gi ¼ F1 v;
vG ¼ v i G i ¼ F T v
ð3:17Þ
which are called the contravariant and the covariant pull-back operation, respectively. In particular, for the pure rotation with F ¼ R leading to Gi ¼ RT gi ; Gi ¼ RT gi , the two relations in Eq. (3.17) unified to the single relation. R v R ¼ vi RT gi ¼ V ¼ vi RT gi ¼ RT v
ð3:18Þ
in which the distinction between the covariant and the contravariant transformation diminishes. In the inverse to the above, the Lagrangian vector defined on the reference base fGi g or Gi , i.e. V ¼ V i Gi ;
V ¼ Vi Gi
ð3:19Þ
!g V ¼ Vi gi
ð3:20Þ
is transformed to the Eulerian vectors ! V g ¼ V i gi ;
by the push-forward operations replacing the reference base fGi g or Gi to the ! !g current base fgi g or fgi g, where V g and V are called the contravariant and the covariant push-forward operation, respectively. Here, because of ! V g ¼ V i FGi ;
!g V ¼ Vi FT Gi
noting Eq. (3.11), the following direct notations hold. ! V g ¼ V i gi ¼ FV;
!g V ¼ Vi gi ¼ FT V
ð3:21Þ
Further, for the pure rotation with F ¼ R leading to gi ¼ RGi ; gi ¼ RGi , it follows from Eq. (3.21) that
3.3 Pull-Back and Push-Forward Operations
101
!R ! V R ¼ V i RGi ¼ V ¼ Vi RGi ¼ RV
ð3:22Þ
Next, the second-order tensor t in the current configuration is represented by the following four types from Eq. (1.433). i t ¼ tij gi gj ¼ tj gi g j ¼ tij gi gj ¼ tij gi g j
ð3:23Þ
Here, when tij ¼ tji ; ti j ¼ tji ; tij ¼ tji ; tij ¼ tji , t is the symmetric tensor and thus t ¼ tT holds by virtue of Eq. (1.441). Here, note that the primary base vector gi and that of the the component vi obey the ordinary coordinate-transformation
rule as
derivative @f =@hi of scalar function f , i.e. @f =@hi ¼ @f =@h
j
j
@h =@hi , while gi
and vi obey the opposite transformation rule as explained in detail in Hashiguchi and Yamakawa (2012) and Hashiguchi (2020). Then, gi and vi are called the covariant base vector and the covariant component, respectively, and gi and vi are called the contravariant base vector and the contravariant component, respectively. The Eulerian tensor in Eq. (3.23) is replaced to the Lagrangian tensors by the pull-back operations replacing both of the current base vectors to the reference base vectors as follows: 8 > t > > > < t > t > > > : t
1 1 ij T ij ij 1 GG t Gi Gj ¼ t F gi F gj ¼ F t gi gj F G j T j 1 i i i 1 j G tj Gi G ¼ tj F gi F g ¼ F tj gi g F j i j T i 1 G gj ¼ FT tij gi gj FT G ti G Gj ¼ ti F g F i j T i T j T i j GG tij G G ¼ tij F g F g ¼ F tij g g F
leading to t
T
¼ tij Gi Gj ¼ F1 tFT ¼ t GG ;
i T T GT tG ¼ t G G ¼ F tF ¼ 6 t j G i G ;
GG
T i ¼ tj Gi G j ¼ F1 tF 6¼ t G G
GG t ¼ tij Gi G j ¼ FT tF ¼ t GG T
t
G G
ð3:24Þ where t gg , t gg t gg and t gg are called the contravariant, the contravariantcovariant, the covariant-contravariant and the covariant push-forward operation, respectively. For the pull-back operation by which only one of the two current base vectors is exchanged to the reference base vector, the hat symbol (^) is added to the unchanged one as follows:
102
3
Description of Tensor (Rate) in Convected Coordinate System
8 9 9 ij ii 1 ij ij 1 > = = t ¼ t G g ¼ t F g g t ¼ t g G ¼ t g F g G^ g i g ^ j > j r J i i j G > 1 > > ¼ F ¼ tFT t; > ^ g ^ g > ; j i j i i j i 1 1 ; > < t G ¼ tj Gi g ¼ tj F gi gj t G ¼ ti g Gj ¼ ti g F gj 9 9 G^ g > > i j T i j > t ^gG ¼ tij gi G j ¼ tij gi FT g j = ¼ tij G g ¼ tij F g g = > t > > ¼ FT t; ¼ tF > > > j i i T j; : t G ¼ tj Gi g ¼ tj FT gi g > ; t G ^ g ¼ tj gi G ¼ tj gi F g j j i i ^ g
ð3:25Þ Incidentally, when only the rotation is given to both base vectors, the pull-backed tensor are described by the following unique equation in which the distinction of covariant, contravariant and mixed-variant operations vanishes: ^
T t
R
^
¼ RT tR ¼ TT
ð3:26Þ
^
which is derived by setting F ¼ R in Eq. (3.24). T is called the corotational Lagrangian tensor. On the other hand, the Lagrangian tensor defined by the reference bases fGi g; Gi is represented in the four types T ¼ T ij Gi Gj ¼ Tji Gi G j ¼ Tij Gi Gj ¼ Tij Gi G j
ð3:27Þ
The following push-forwarded Eulerian tensors are obtained by replacing the reference base vectors to the current base vectors. 8! > T gg T ij gi gj ¼ T ij FGi FGj ¼ FT ij Gi Gj FT > > > !g < T g ¼ T ji gi g j ¼ T ji FGi FT G j ¼ FTij Gi G j F1 g >! T g ¼ Tij gi gj ¼ Tij FT Gi FGj ¼ FT Tij Gi Gj FT > > > : !gg T ¼ Tij gi g j ¼ Tij FT Gi FT Gj ¼ FT Tij Gi G j F1 leading to !
! T T gg ¼ T ij gi gj ¼ FTFT ¼ T gg ;
!
! gT T g ¼ Tji gi g j ¼ FTF1 6¼ T g !ggT
!g !gT !gg T g ¼ Tij gi gj ¼ FT TFT ð6¼ T g Þ; T ¼ Tij gi g j ¼ FT TF1 ¼ T
ð3:28Þ !gg ! !g !g where T gg , T g , T g and T are called the contravariant, the contravariantcovariant, the covariant-contravariant and the covariant push-forward operation, respectively.
3.3 Pull-Back and Push-Forward Operations
103
For the push-forward operation in which only one of the two reference base vectors is changed to the current base vector, it is denoted by adding the hat symbol (^) to the unchanged one as follows:
ð3:29Þ Incidentally, when only the rotation is given to both base vectors, the push-forwarded tensor is represented by the following unified equation in which the distinction of covariant, contravariant and mixed-variant operations vanishes: _
!R t T ¼ RTRT ¼ t T
_
ð3:30Þ
which is called the corotational Eulerian tensor. The metric tensors in the reference and the current configurations are related in the pull-back and the push-forward operations as follows: (
G Gi Gi ¼ F1 gi FT gi ¼ F1 gF ¼ g G G¼ G Gi Gi ¼ FT gi F1 gi ¼ FT gFT ¼ g G 8 g < g gi ¼ FG FT Gi ¼ FGF1 ¼ ! Gg i i g¼ g : gi g ¼ FT Gi FG ¼ FT GFT ¼ ! G i
i
g
9 > > > > > > = > > > > > > ;
ð3:31Þ
The tensors in the current configuration and in the referent configuration are referred to as the Eulerian tensor and the Lagrangian tensor, respectively. It is shown that we have the two and the four types of transformations from the Eulerian to the Lagrangian vector and second-order tensor or its inverse transformations, respectively, since the vector and the second-order tensor possesses one and two base vectors, respectively. In general, there exist n 2 types of transformations for the n-th order tensor. It should be noticed from the physical point of view that practically the Eulerian tensor is defined first and thereafter the Lagrangian tensors are produced from the Eulerian tensor by the pull-back operation. It will be shown for stress tensors in Sect. 5.1.
104
3
Description of Tensor (Rate) in Convected Coordinate System
A material involved in the unit volume in the current configuration changes but that in the reference configuration does not change during a deformation. Therefore, a constitutive equation describing the inherent property of material must be formulated by the tensors based in the reference configuration, which are given by the pulled-back of the tensors in the current configuration and the deformation analysis must be performed by the constitutive equation in the reference configuration. Note, however, that the boundary condition is given in the current configuration and thus first the variables must be pulled-back to the reference configuration. Therefore, the deformation analysis must be performed by the constitutive equation in the reference configuration and after that the calculated result is push-forwarded to the current configuration, by which the actual states of stress and deformation are captured. Gurtin et al. (2010) proposed the symbols specifying the pull-back operations P½t FT tF; P½t F1 tFT and the push-forward operations P1 ½t FT tF1 ; P1 ½t FtFT However, the covariant-contravariant mixed operations cannot be specified by these symbols. On the other hand, the symbols with the arrow and superscript or subscript shown in this section are capable of specifying the four types of the pull-back operations and the four thypes of the push-forward operations, providing the whole informations for these operations in the concise forms.
3.4
Convected Time-Derivative
As will be described in detail in Chap. 6, the constitutive relation of material does not depend on the rigid body rotation of material. Here, the material-time derivatives of the stress and the tensor-valued internal variables observed by a fixed coordinate system are influenced by the rigid-body rotation of material. Therefore, the material-time derivatives in Eq. (2.12) described in a fixed coordinate system cannot be used in the description of constitutive equation. On the other hand, the rates of stress and internal variables observed by the embedded coordinate system, which deforms and rotates with the material itself, is irrelevant to the rigid body rotation. Therefore, the constitutive equation must be represented in terms of the embedded derivatives of variables. The physical meaning and the mathematical expression of the time-derivatives of vector- and tensor-valued variables observed by the embedded coordinate system will be explained in this section, which is called the embedded time-derivative or the convective time-derivative or the convected time-derivative.
3.4 Convected Time-Derivative
105
3.4.1 General Convected Derivative The material-time derivative of the second-order tensor t based in the current configuration in Eq. (1.433) is described by 8 ij t gi gj ¼ tij gi gj þ t ij gi gj þ tij gi g > > > j > > >
> j < i i j j i j i tj gi g ¼ t j gi g þ tj gi g þ tj gi g t¼ > j i > > ti g gj ¼ tij gi gj þ tij gi gj þ ti gi gj > > > > : tij gi g j ¼ tij gi g j þ tij gi g j þ tij gi g j
ð3:32Þ
from which the convective derivatives, i.e. the rate of the tensor t observed by the observer moving with the embedded coordinate system are given by 8 ij > t gi gj ¼ t tij gi gj tij gi gj ¼ t tij lgi gj tij gi gj lT ¼ t lt tlT > > > > > > i < ti j gi g j ¼ t tj gi gj ti j gi gj ¼ t ti j lgi gj þ ti j gi g j l ¼ t lt tl > > > ti jgi gj ¼ t tij gi gj tij gi gj ¼ t þ ti lT gi gj ti j gi gj lT ¼ t lT t tlT > > > > : t_ij gi g j ¼ t tij gi g j tij gi g j ¼ t þ tij lT gi g j þ tij gi g j l ¼ t lT t tl
ð3:33Þ noting Eq. (3.12). The following four types of convected derivatives can be defined from Eq. (3.33).
o
o
o
o
o
o
o
o
ð3:34Þ
106
3
Description of Tensor (Rate) in Convected Coordinate System
noting F F1 ¼ l; FT FT ¼ FT FT ¼ lT because of FF1 ¼ F F1 þ gg g g F F1 ¼ O. The four types of the convected time-derivatives tgg , tg , tg and t are the push-forward of the material-time derivative of the pull-backed tensor by the contravariant, the contravariant-covariant, the covariant-contravariant and the covariant type operation, respectively. Namely, the convective time-derivative are attained by. 1. first pulling-back the tensor in the current configuration to the reference configuration, 2. then taking its time-differentiation, 3. finally returning (pushing-forward) it to the current configuration, There are the four types of the pull-back operations in the second-order tensor which possesses the two base vectors. Here, the variant type of the pull-back and the push-forward operation is same to each other. Incidentally, the particular operators, e.g. the identity tensor I, R are also used in addition to the original operator F. The convective time-derivative has been called often the Lie derivative and denoted by the symbol Lv ðf Þ defined by Lv ðf Þ v
D 1 v f Dt
where f is scalar, vector or tensor variable, and v and v1 designate the push-forward and the pull-back operation, respectively (cf. e.g., Truesdell and Noll 1965; Marsden and Hughes 1983; Simo and Hughes 1998; Holzpfel 2000; Bonet and Wood 2008; de Souza Neto et al. 2008; Belytschko et al. 2014). However, the symbol with the double arrows and the two indices shown in this section would express the operations and modes (covariant or contravariant) of the convected time-derivatives clearly and intuitively. It was introduced by Hashiguchi (2011). However, the hypoelastic-based constitutive equation represented by the above-mentioned convective time-derivative will be disused sooner or later by the adoption of the multiplicative-based constitutive equation as will be explained in detail in Chap. 17.
3.4.2 Corotational Rate An irreversible deformation exhibits the loading-path dependency. Therefore, a constitutive equation for an irreversible deformation must be described by the rate-form. Therein, pertinent rates of stress and internal variables must be used in the constitutive equation, which satisfy the objectivity described the next
3.4 Convected Time-Derivative
107
subsection. Here, consider the particular cases in which the deformation and its rate are negligible compared with the rotation, i.e. U ffi I or d ffi O: For the case that the deformation is small compared with the rotation, i.e. F ffi R, all the four kinds of derivatives in Eq. (3.34) reduce to the single equation: oR
o
o• g
og
o gg
t ≡ t gg = t g = t • g = t
→
)• )• R • • • = R (R tR) R = R T RT ≡ T = t − R RT t + t R RT •
T
T
tR
ð3:35Þ
is zero which is referred to the Green-Naghdi rate (Green and Naghdi 1965). when the quantity RT tR, which is observed by the coordinate system moving and rotating with a material line-element, is kept constant. The relation of the base vectors are given from Eq. (3.8) under F ¼ R as follows: 8 < gi ¼ RGi ; gi ¼ RGi G ¼ RT gi ; Gi ¼ RT gi : i R ¼ gi Gi ¼ gi Gi
ð3:36Þ
Then, the rate of the current base vector is given as follows:
gi ¼ XR gi
ð3:37Þ
noting gi ¼ R Gi ¼ R RT gi , where
XR R RT
ð3:38Þ
is referred to the relative spin. It designates the spin of the normalized convected covariant base vector gi =kgi k under the rigid-body rotation in which ðgi =kgi kÞ ¼ XR ðgi =kgi kÞ holds because of kgi k ¼ const. and kðgi =kgi kÞk ¼ 1. The physical meaning of the relative spin based on the name capped with the adjective “relative” will be given in Sect. 4.4. The Green-Naghdi rate in Eq. (3.35) is represented in terms of the relative spin as follows:
t R ¼ t XR t þ tXR ¼ t RT
ð3:39Þ
Further, in the case that the strain rate d in Eq. (2.43) is infinitesimal compared with the continuum spin w in Eq. (2.44), leading to l ¼ w, Eq. (3.34) is reduced to
t w t wt þ tw ¼ t wT
ð3:40Þ
108
3
Description of Tensor (Rate) in Convected Coordinate System
which is called the Zaremba-Jaumann rate (Zaremba 1903; Jaumann 1911). Here,
tw ignoring the strain rate d is related to the convected derivative in Eq. (3.34) as follows: o
o
o
o
g
o
gg • g t w = t gg + t d +dt = t g − td +dt = t • g + t d − dt = t −td −dt
ð3:41Þ
resulting in o o o o g og o g o g og og t w = 1 (t gg + t gg ) = 1 (t •g + t • g ) = 1 [t •g + (t •g )T ] = 1 [t • g + (t • g )T ] 2 2 2 2
ð3:42Þ
The corotational rates in Eqs. (3.39) and (3.40) depend only on the geometrical change of material which can be observed from the outside appearance of material. However, the physically meaningful rotation is the rotation of the substructure of material which is induced by the elastic distortion and the rigid-body rotation but is not induced by the plastic deformation related merely to the mutual slips between material particles, and thus the rotation of the substructure of material cannot be known only from the outside appearance of material. Based on this physical interpretation, it has been pointed out by Mandel (1971, 1973), Kratochvil (1971), etc. that the rotation of substructure is suppressed by the plastic deformation such that the substructure spin wS is given by the subtraction of the plastic spin wp (Dafalias 1984, 1985a, 1998) from the continuum spin w, i.e.
t t ws t þ tws
ð3:43Þ
wS ¼ w wp ¼ we
ð3:44Þ
with
The explicit equation of the plastic spin wp will be shown in the subsequent chapters. The oscillations of the stress and the kinematic hardening variable are suppressed by the corotational rate with the modified spin tensor in Eq. (3.44).
3.4.3 On Adoption of Convected Rate Tensor in Hypoelastic Constitutive Equation The influence of the rigid-body rotation occurs in the deformation analysis by the constitutive equation described in terms of the material time-derivative denoted by the symbol ð Þ. The influence can be reduced greatly by exploiting the convected (embedded) time-derivative which is the variation of the tensor observed from the coordinate system deforming/rotating with the material itself. However, there exist
3.4 Convected Time-Derivative
109
the four types of the convected time-derivative. Unfortunately, which one is rational in them cannot be concluded. In addition, the time-integration of them by the forward Euler method leads to an erroneous result as far as the calculation by infinitesimal increments is not performed. Instead, the time-integration must be executed by the pull-back, the time-integration and the push-forward operation in the three steps as will be explained in the next subsection.
3.4.4 Time-Integration of Convected Rate Tensor The exact time-integration of the convective rate variable, e. g. stress rate and rate of internal variables is of crucial importance in the deformation analysis by the hypoelastic-based constitutive equation. The forward-Euler calculation method of the stress and internal variables in the hypoelastic constitutive relation has to be done in infinitesimally small incremental steps, otherwise large error will be induced. Then, an efficient calculation method has been developed and installed as the standard-function in the commercial FEM software, e.g. Mark and Abaqus, etc., while a detailed theoretical background is not written in their user-manuals. In what follows, it will be explained systematically from the general aspect of the convected rate tensors and the corotational tensors, although it has been delineated in fragments taking account of only the material rotation by some literatures (e.g. Hughes and Winget 1980; Pinsky et al. 1983; Rubinstein and Atluri 1983; Flagnan and Taylor 1987; Simo and Hughes 1998; de Souza Neto et al. 2008). (a) Pull-back and Push-forward time-integration method The rigid-body rotation of material occurs finitely, which is larger compared with the deformation in many cases as known from the screw, the turbine, the whleel, etc., while needless to say the rigid-body rotation is irrelevant to the material property, i.e. the constitutive relation. Therefore, the variation of tensor variables must be executed by an infinitesimal increment in the current configuration in general. However, the time-integration of tensor variables can be executed in finite increments by the calculation with the three steps, i.e. the pull-back operation to the reference configuration, the incremental calculation by the constitutive equation in the reference configuration and the push-forward operation to the current configuration as described in Sect. 3.4.1.
The convected rate t in Eq. (3.34) can be represented collectively in terms of the second-order tensors F l and F r which play the roles of the pull-back and push-forward operations, noting that the operators for the pull-back and the push-forward are the inverse relation to each other, as follows: 1
t ¼ F l F l tF r F r1
ð3:45Þ
110
3
Description of Tensor (Rate) in Convected Coordinate System
leading to ð3:46Þ where the second-order tensors F l and F r are given for the four types of the
convective time-derivative t in Eq. (3.34) as follows: ⎧ → • ⎪ l o −T r • for t gg ≡ ← t gg = F(F −1tF −T )• FT = t − lt − tl T ⎪ = F, = F ⎪ → ⎪ • •g g • −1 −1 ⎪ l = F, r = F for to g• ≡ ← t g = F (F tF)• F = t − lt + tl ⎪⎪ ⎨ → ⎪ l •g og ← r − ⎪ = F −T , = F T for t • g ≡ t • g = F −T (FT tF −T )• FT = t• + l T t −t l T ⎪ ⎪ → gg ⎪ • gg ← ⎪ l = F −T , r = F for to ≡ t = F −T (FT tF)• F −1 = t• + l T t + t l ⎪⎩
ð3:47Þ
We consider the four examples of the typical convective rates.
(1) The Oldroyd rate is based on the contravariant rate tgg in Eq. (3.34)1 , i.e. (3.47)1 : F l ¼ F; F r ¼ FT
(2) The Cotter-Rivlin rate is based on the covariant rate tgg in Eq. (3.34)4 , i.e. (3.47)4 : F l ¼ FT ; F r ¼ F (3) The Green-Naghdi rate in Eq. (3.39) is derived by setting Fl ¼ Fr ¼ R (4) The Jaumann rate in Eq. (3.40) is given by 1 ¼ F r F r1 ¼ w F l F l i.e.
F l ¼ wF i ;
F r ¼ wF r
ð3:48Þ
ð3:49Þ ð3:50Þ . Then, denoting F l
noting
and F r collectively by F , i.e. F l ¼ F r ¼ F , Eq. (3.50) is represented in the unified form as follows:
F ¼ wF
ð3:51Þ
3.4 Convected Time-Derivative
111
The incremental relation of Eq. (3.51) is given by F n þ 1 ¼ exp wn þ 1=2 Dt F n
ð3:52Þ
noting Eq. (A.57) with h ¼ 1=2.
In what follows, we consider the time-integration of the convected rate tensor t for the all cases in Eq. (3.47). Firstly, the pull-back of t in the step n þ 1 to the reference configuration is given by t
nþ1
r ¼ t n þ tn þ 1=2 Dt ¼ F nl1 tn F rn þ F nl1 þ 1=2 tn þ 1=2 F n þ 1=2 Dt
ð3:53Þ
noting Eq. (3.45), where t is given by the constitutive relation in the current configuration. Then, the updated tensor tn þ 1 in the current configuration is calculated by the push-forward of Eq. (3.53) as follows: tn þ 1 ¼ F ln þ 1 t
r1 n þ 1F n þ 1
r l1 r ¼ F ln þ 1 F l1 t F þ F t F Dt F r1 n n þ 1=2 n n n þ 1=2 n þ 1=2 nþ1
ð3:54Þ leading to
tn þ 1 ¼ F lD tn F rD þ F ld tn þ 1=2 F rd Dt
ð3:55Þ
where (
F lD F ln þ 1 F nl1 ; F rD F rn F r1 nþ1 r r r1 F ld F ln þ 1 F nl1 ; F F þ 1=2 d n þ 1=2 F n þ 1
ð3:56Þ
The variables in Eq. (3.56) are represented by F noting Eq. (3.47) as follows:
112
3
8 > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > < > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > > :
Description of Tensor (Rate) in Convected Coordinate System
r T T F lD ¼ Fn þ 1 F1 n ; F D ¼ Fn Fn þ 1
9 =
T ; F ld ¼ Fn þ 1 Fn þ 1=2 ; F rd ¼ FT n þ 1=2 Fn þ 1
F lD ¼ Fn þ 1 F1 n ;
for contravariant: Oldroyd rate
9 = F rD ¼ Fn F1 nþ1
; F rd ¼ Fn þ 1=2 F1 nþ1 9 T = F lD ¼ F1 F rD ¼ FT n þ 1 Fn ; n Fn þ 1
F ld ¼ Fn þ 1 F1 n þ 1=2 ;
for contravariant-covariant
for covariant-contravariant T ; F rd ¼ FT n þ 1=2 Fn þ 1 9 T = F lD ¼ FT F rD ¼ Fn F1 n þ 1 Fn ; nþ1 for covariant: Cotter-Rivlin rate T ; F ld ¼ FT F rd ¼ Fn þ 1=2 F1 n þ 1 Fn þ 1=2 ; nþ1 9 F lD ¼ Rn þ 1 RTn ; F rD ¼ Rn RTn þ 1 = for Green-Naghdi rate F ld ¼ Rn þ 1 RTn þ 1=2 ; F rd ¼ Rn þ 1=2 RTn þ 1 ; ) F lD ¼ exp wn þ 1=2 Dt ; F rD ¼ F lT D ¼ exp wn þ 1=2 Dt for Jaumann rate F ld ¼ exp wn þ 1=2 Dt=2 ; F rd ¼ F lT d ¼ exp wn þ 1=2 Dt=2 F ld ¼ F1 n þ 1 Fn þ 1=2 ;
ð3:57Þ noting ) F lD ¼ exp wn þ 1=2 Dt F n F 1 ¼ exp w Dt ; n þ 1=2 n1 F n ¼ exp wn þ 1=2 Dt F rD ¼ F n exp wn þ 1=2 Dt
ð3:58Þ
1 ) F ld ¼ exp wn þ 1=2 Dt F n exp wn þ 1=2 Dt=2 F n ¼ exp wn þ 1=2 Dt=2 1 F rd ¼ exp wn þ 1=2 Dt=2 F n exp wn þ 1=2 Dt F n ¼ exp wn þ 1=2 Dt=2 ð3:59Þ with Eq. (3.52). The following approximate equations for the Jaumann rate in Eq. (3.57) were proposed by Hughes and Winget (1980) taking only the first and second-order terms in the Taylor expansion in Eq. (1.344).
F TD þ tn þ 1=2 Dt tn þ 1 ¼ f F D tn F n f
ð3:60Þ
f F D I þ wn þ 1=2 Dt
ð3:61Þ
where
3.4 Convected Time-Derivative
113
noting (
exp wn þ 1=2 Dt ffi I þ wn þ 1=2 Dt; exp wn þ 1=2 Dt ffi I wn þ 1=2 Dt exp wn þ 1=2 Dt=2 ffi I þ wn þ 1=2 Dt=2 ffi I; exp wn þ 1=2 Dt=2 ffi I wn þ 1=2 Dt=2 ffi I
ð3:62Þ Only the influence caused by the rigid-body rotation is excluded as seen in the first term in the right-hand side in Eq. (3.60). Equation (3.60) is described for the Cauchy stress by
rn þ 1 ¼ f F D rn F n f F TD þ E : dn þ 1=2 dpn þ 1=2 Dt
ð3:63Þ
and for the kinematic hardening variable a described in Chap. 8 by F D an F n f F TD þ ck dpn þ 1=2 an þ 1 ¼ f
1 bk Fn þ 1=2
p dn þ 1=2 an þ 1=2 Dt
ð3:64Þ
(b) Numerical examples The numerically calculated results by the above-mentioned Jaumann rate and the Green-Naghdi rate in the simple shear deformation of the hypoelastic material are shown in Figs. 3.1 and 3.2. These calculations have been executed by Prof. Yuki Yamakawa (Tohoku university) and they are published here by his courtesy. The calculation by the forward-Euler method deviates from the exact calculation curve with the decrease of the number of calculation, i.e. the increase of strain increment as shown in Figs. 3.1a and 3.2a for the Jaumann rate and the
(a) Forward-Euler method
Fig. 3.1 Numerical time-integrations of Jaumann rate
(b) Pull-back and push-forward method
114
3
Description of Tensor (Rate) in Convected Coordinate System
(a) Forward-Euler method
(b) Pull-back and push-forward method
Fig. 3.2 Numerical time-integrations of Green-Naghdi rate
Green-Naghdi rate, respectively. Here, the exact calculation for the Jaumann rate is shown by the sinusoidal curve, because the material rotates constantly by the constant spin w if the plastic spin is not incorporated as described in Sect. 3.4.2. On the other hand, the exact calculations can be executed by the pullback-push forward calculation method in Eq. (3.55) with Eq. (3.57) independent of the shear strain increment as shown in Figs. 3.1b and 3.2b. There does not exist the definite answer to the question which convected rate tensor is best. The Jaumann rate is practically accepted widely because it based on the continuum spin as the anti-symmetric part of the velocity gradient tensor which is the most basic quantity describing the rate of deformation and the incorporation of the plastic spin tensor is physically acceptable. However, the hypoelastic equation holds only for the infinitesimal elastic deformation as will be described in Sect. 17.3. Such defect in the hypoelasticty can be excluded in the multiplicative hyperelastic-based plastic constitutive equation described in Chap. 17.
Chapter 4
Deformation/Rotation Tensors
There are various deformation/rotation (rate) measures, from which the most suitable measure must be adopted in the description of constitutive equation, depending on the purpose, e.g. the description of infinitesimal deformation, finite deformation, hyper- or hypo-elastic deformation, (visco)plastic deformation, etc. In this chapter, definitions and physical meanings of main deformation/rotation (rate) measures used in continuum mechanics will be explained based on the deformation gradient tensor from which all the deformation/rotation (rate) measures are deduced. In particular, various strains, strain rates and spins will be introduced as the deformation/rotation (rate).
4.1
Deformation Tensors
Applying the polar decomposition in Sect. 1.11 to the deformation gradient F, we have F ¼ RU ¼ VR ;
FiA ¼ RiR U RA ¼ Vir RrA
ð4:1Þ
U ¼ RT F ¼ ðFT FÞ1=2 ðU2 ¼ FT FÞ ðUT ¼ UÞ
ð4:2Þ
V ¼ FRT ¼ ðFFT Þ1=2 ðV2 ¼ FFT Þ ðVT ¼ VÞ
ð4:3Þ
where
R ¼ FU1 ¼ FðFT FÞ1=2 ; R ¼ V1 F ¼ ðFFT Þ1=2 F ðdetR ¼ 1Þ V ¼ RURT ; U ¼ RT VR
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_4
ð4:4Þ ð4:5Þ
115
116
4
Deformation/Rotation Tensors
U and V are the symmetric tensors describing the pure deformation which possess the different principal directions depending on the orthogonal tensor R describing the rotation. They are the similar tensors to each other, since they are related by Eq. (4.5) through the orthogonal tensor R as was described in Sect. 1.11. Therefore, they possess the same principal stretches, say ka ð[ 0Þ ða ¼ 1; 2; 3Þ. Denoting the bases for the principal directions of U and V by fNa ðtÞg and fnðaÞ ðtÞg, respectively, they can be written as U¼
3 P a¼1
ka NðaÞ ðtÞ NðaÞ ðtÞ; V ¼
3 P a¼1
ka nðaÞ ðtÞ nðaÞ ðtÞ
ð4:6Þ
where the relation of NðaÞ ðtÞ and nðaÞ ðtÞ is given from Eq. (1.276) as follows: nðaÞ ðtÞ ¼ RðtÞNðaÞ ðtÞ; NðaÞ ðtÞ ¼ RT ðtÞnðaÞ ðtÞ
ð4:7Þ
with RðtÞ ¼
3 P
nðaÞ ðtÞ NðaÞ ðtÞ
ð4:8Þ
a¼1
NðaÞ ðtÞ and nðaÞ ðtÞ are called the Lagrangian triad and the Eulerian triad, respectively. Substituting Eqs. (4.6) and (4.8) into Eq. (4.1), F and its inverse tensor are described by FðtÞ ¼
3 P a¼1
ka ðtÞnðaÞ ðtÞ NðaÞ ðtÞ; F1 ðtÞ ¼
3 P 1 NðaÞ ðtÞ nðaÞ ðtÞ a¼1 ka ðtÞ
ð4:9Þ
Let the mechanical meanings of U; V and R be examined below. The variation of infinitesimal line-element is given by the polar decomposition F ¼ RU as follows: dx ¼ FdX ¼ RUd X ¼ R
3 X b¼1
kb NðbÞ NðbÞ
3 X a¼1
dXa NðaÞ ¼ R
3 X
ka dXa NðaÞ
a¼1
ð4:10Þ Equation (4.10) means that the infinitesimal line-elements dXa NðaÞ (no sum) in the principal directions NðaÞ are first stretched by ka times to ka dXa NðaÞ (no sum) and then undergoes the rotation R as shown in Fig. 4.1. On the other hand, the change of the infinitesimal line-element by the polar decomposition VR is described as
4.1 Deformation Tensors
117
RL
RE
RL
R
RE
R
L
Fig. 4.1 Polar decomposition of deformation gradient
dx ¼ FdX ¼ VRdX ¼
3 X
kb nðbÞ nðbÞ R
b¼1
¼
3 X a¼1
ka dXa nðaÞ ¼
3 X
dXa NðaÞ ¼
a¼1 3 X
3 X b¼1
kb nðbÞ nðbÞ
3 X
dXa nðaÞ
a¼1
ka RdXa NðaÞ
a¼1
ð4:11Þ
118
4
Deformation/Rotation Tensors
Equation (4.11) means that the infinitesimal line-elements dXa NðaÞ (no sum) in the principal directions NðaÞ first becomes dXa nðaÞ (no sum) by rotation R and then are stretched by ka times to ka dXa nðaÞ (no sum) (see Fig. 4.1). As described above, U and V designates the deformation and R the rotation. ka is called the principal stretch, and U and V are called the right and left stretch tensor, respectively. Letting RL and RE designate the rotations of the Lagrangian triad fNðaÞ g and the Eulerian triad fnðaÞ g, respectively, from the fixed base fea gða ¼ 1; 2; 3Þ, they are given by RL
3 P a¼1
NðaÞ ea ; RE
3 P a¼1
nðaÞ ea
ð4:12Þ
where the following relations hold. NðaÞ ¼ RL ea ; nðaÞ ¼ RE ea
ð4:13Þ
RE ¼ RRL
ð4:14Þ
Considering the particle P and the adjacent particles P0 and P00 , we designate their position vectors before and after the deformation by X; X þ dX; X þ dX and x; x þ dx; x þ dx, respectively. Then, noting (1.139), one has
dx dx ¼ FdX FdX ¼ FT FdX dX ¼ CdX dX 1 dX dX ¼ F1 dx F1 dx ¼ FT F1 dx dx ¼ FFT dx dx ¼ b1 dx dx ð4:15Þ
where C U2 ¼ FT F ¼ FT g F ¼ g GG ¼ gij Gi G j ¼ CT C1 ¼ F1 FT ¼ F1 g FT ¼ g ¼ gij Gi Gj ¼ CT
ð4:16Þ
! b V2 ¼ FFT ¼ FGFT ¼ G gg ¼ Gij gi gj ¼ bT !gg b1 ¼ FT F1 ¼ FT GF1 ¼ G ¼ Gij gi g j ¼ bT
ð4:17Þ
GG
as ascertained by
FT F ¼ Gi gi gj Gj ¼ gij Gi Gj FFT ¼ gi Gi Gj gj ¼ Gij gi gj
4.1 Deformation Tensors
119
so that C and b are related to each other as follows: (
C ¼ FT bFT ¼ b G
G
ð4:18Þ
G
b ¼ FT CT ¼ C G
C and b are referred to as the right and left Cauchy-Green deformation tensor, respectively. In accordance with Eq. (4.6) they are described by C¼
3 P a¼1
k2a NðaÞ NðaÞ ; b ¼
3 P a¼1
k2a nðaÞ nðaÞ
ð4:19Þ
where ka is called the principal stretch. The principal values k2a are obtained by the solutions T of the characteristic equation T 3 Ic T 2 IIc T þ IIIc ¼ 0
ð4:20Þ
based on Eq. (1.216), where Ic trC; IIc
i 1h 1 1 1 ðtrCÞ2 trC2 ; IIIc ðtrCÞ3 trCtrC2 þ trC3 2 6 2 3
ð4:21Þ
The principal values and directions are calculated by the method described in Sect. 1.6. The similar equations hold for b instead of C. Using the relative description (2.8), the relative deformation gradient tensor in the reference configuration xðsÞ is defined as s FðtÞ
¼
@xðtÞ @xðsÞ
ð4:22Þ
which is related to the deformation gradient FðtÞ ð 0 FðtÞÞ as FðtÞ ¼
@xðtÞ @xðtÞ @xðsÞ ¼ ¼ s FðtÞFðsÞ @X @xðsÞ @X
ð4:23Þ
and is further expressed in the polar decomposition as s FðtÞ
¼ s RðtÞs UðtÞ ¼ s VðtÞs RðtÞ
ð4:24Þ
while s CðtÞ and s bðtÞ are defined by s CðtÞ
¼ ðs FðtÞÞ T s FðtÞ ¼ s U2 ðtÞ
s bðtÞ
¼ s FðtÞ ð s FðtÞÞT ¼ s V2 ðtÞ
) ð4:25Þ
which are called the relative right and the left Cauchy-Green deformation tensors.
120
4.2
4
Deformation/Rotation Tensors
Strain Tensors
Consider the scalar quantity which changes only by the pure deformation but is independent of the rotation. Subtracting the upper equation from the lower equation in Eq. (4.15), one has ( 2EdX dX ð¼ 2EAB dXA dXB Þ dx dx dX dX ¼ ð4:26Þ 2edx dx ð¼ 2eij dxi dxj Þ where 1 1 1 1 @x @x E ðC IÞ ¼ ðU2 IÞ ¼ ðFT F IÞ ¼ ð ÞT ð Þ I 2 2 2 2 @X @X 1 1 @xk @xk EAB ðFkA FkB dAB Þ ¼ dAB 2 2 @XA @XB T 1 1 1 1 @X @X e ðIb1 Þ ¼ ðIV2 Þ ¼ ðI FT F1 Þ ¼ I 2 2 2 2 @x @x 1 1 @XK @XK eij ½dij ðF1 ÞKi ðF1 ÞKj ¼ dij @xi @xj 2 2
9 > > > > > > > > > > > > > = > > > > > > > > > > > > > ;
ð4:27Þ
which are defined by C and b describing the pure deformations, noting dx dx ¼ FdX FdX ¼ FT FdX dX and dX dX ¼ F1 dx F1 dx ¼ FT F1 dx dx¼ ðFFT Þ1 dx dx. Applying the quotient law described in Sect. 1.3.2 to Eq. (4.26), it is confirmed that E and e are the second-order tensors. If a deformation is not induced, the triangle PP0 P00 keeps the same shape and thus the left-hand side in Eq. (4.26) is zero so that E and e are zero independent of rotation. Conversely, if E 6¼ O; e 6¼ O, the left-hand side in Eq. (4.26) is not zero so that the shape of the triangle is not same as that in the initial, resulting in a deformation. Therefore, E and e are the quantities describing the deformation independent of rigid-body rotation and called the Green strain tensor and the Almansi strain tensor, respectively. Using the displacement vector u ¼ x X ¼ ui ei ;
@u ¼FI @X
ð4:28Þ
they are expressed by " 9 T # 1 @u @u @u @u 1 @uA @uB @uK @uK > > > þ Þð ; EAB ¼ E¼ þ þ þ > = 2 @X @X @X @X 2 @XB @XA @XA @XB > " # T T > > 1 @u @u @u @u 1 @ui @uj @uk @uk > > > e¼ þ þ ; eij ¼ ; @xi @xi @xj 2 @x @x @x @x 2 @xj ð4:29Þ
4.2 Strain Tensors
121
The following relation exists between them, noting Eq. (3.24)4 for the pull-back operation and Eq. (3.28)4 for the push-forward operation. 9 GG > T = E ¼ F eF ¼ e ; EAB ¼ FiA FjB eeij ð4:30Þ !gg > T 1 e ¼ FT EF1 ¼ E ; eij ¼ FiA FjB EAB ; The Green strain tensor E and the Almansi strain tensor e are called also the Lagrangian strain tensor and the Eulerian strain tensor, respectively, since the former and the latter are represented by the reference and the current variables. Now, consider the symmetric part of the displacement gradient tensor u which is the eliminations of the third terms, i.e. the second-order infinitesimal terms in the brackets E and e in Eq. (4.29), i.e. e
@u @X
s
" T # 1 @u @u 1 1 @uA @uB T þ ¼ þ ¼ F þ F I ; eAB 2 @X @X 2 2 @XB @XA ð4:31Þ
or " s T # @u 1 @u @u 1 1 @ui @uj þ e ¼ þ ¼ I F1 þ FT ; eij @xi @x 2 @x @x 2 2 @xj ð4:32Þ which describe roughly deformation, depending not only on U or V but also on the rotation tensor R, noting F ¼ RU ¼ VR. Incidentally, the distinction between the infinitesimal line elements dX in the initial state and dx in the current state becomes meaningless. Then, e is called the infinitesimal (nominal) strain tensor, while the Green and the Almansi strain tensors are called the finite (true) strain tensor. Consequently, the distinction of the Lagrange and the Eulerian description, i.e. the difference of Eqs. (4.31) and (4.32) vanishes because of dx ffi dX in an infinitesimal deformation. Hereinafter, unless otherwise noted, the infinitesimal strain tensor refer to Eq. (4.31) in the Langrange description. The defects of the infinitesimal (nominal) strain tensor will be explained in Sect. 4.5. In what follows, the geometrical interpretation of E and e will be given. Considering the case that the two infinitesimal line-elements PP0 and PP00 coincide to each other, i.e. dX ¼ dX; dx ¼ dx and denoting their direction vectors in the initial and current configurations by NðjjNjj ¼ 1Þ and nðjjnjj ¼ 1Þ leading to dX ¼ jjdXjjN and dx ¼ jjdxjjn, it follows from Eq. (4.26) that
122
4
( 2
2
jjdxjj jjdXjj ¼
Deformation/Rotation Tensors
2EN NjjdXjj2 2en njjdxjj2
ð4:33Þ
By selecting the X1 -axis and the x1 -axis in the reference and the initial stage, respectively, of this line-element, ðN1 ; N2 ; N3 Þ ¼ ð1; 0; 0Þ and ðn1 ; n2 ; n3 Þ ¼ ð1; 0; 0Þ hold leading to EN N ¼ E11 and En n ¼ e11 and thus we have " #! 9 > 1 jjdxjj2 jjdXjj2 1 jjdxjj 2 > > > ¼ 1 E11 ¼ > 2 = 2 2 jjdXjj jjdXjj " # ! ð4:34Þ > > 1 jjdxjj2 jjdXjj2 1 jjdXjj 2 > > e11 ¼ ¼ 1 > ; 2 2 jjdxjj jjdxjj2 Therefore, E11 and e11 describes the half of the change in the square of the line-element vector. Equation (4.34) becomes E11 e11
9 1 ðjjdxjj jjdXjjÞðjjdxjj þ jjdXjjÞ > > ¼ > = 2 jjdXjj2 1 ðjjdxjj jjdXjjÞðjjdxjj þ jjdXjjÞ > > > ¼ ; 2 jjdXjj2
ð4:35Þ
E11 and e11 are simplified to the following equation in the case of the infinitesimal deformation fulfilling jjdxjj ffi jjdXjj. 9 jjdxjj jjdXjj > ¼ e11 > = jjdXjj > jjdxjj jjdXjj ; ¼ e11 > ffi jjdxjj
E11 ffi e11
ð4:36Þ
which coincides with the infinitesimal strain e in Eq. (4.31) or (4.32). On the other hand, denoting the direction vectors of the two distinct infinitesimal line-element PP0 and PP00 as N0 and N00 , respectively, in the initial state and the angles contained by them as h, it holds from Eq. (4.26) that jjdxjjjjdxjj cos h jjdXjjjjdXjj cos h0 ¼ 2EN0 N00 jjdXjjjjdXjj
ð4:37Þ
i.e. jjdxjj jjdxjj cos h cos h0 ¼ 2EN0 N00 ¼ 2Eij Nj0 Nj00 jjdXjj jjdXjj
ð4:38Þ
where h0 is the initial value of h. Here, assuming that the infinitesimal line-elements PP0 and PP00 were mutually perpendicular before a deformation, i.e. h0 ¼ p=2 leading to cos h0 ¼ 0; and making their directions coincide to the X1 -and X2 -axes,
4.2 Strain Tensors
123
i.e. ðN10 ; N20 ; N30 Þ ¼ ð1; 0; 0Þ; ðN100 ; N200 ; N300 Þ ¼ ð0; 1; 0Þ leading to Eij Ni0 Nj00 ¼ E12 , it follows that E12 ¼
1 jjdxjj jjdxjj 1 jjdxjj jjdxjj cos h ¼ sinðp=2 hÞ 2 jjdXjj jjdXjj 2 jjdXjj jjdXjj
ð4:39Þ
Further, for the particular deformation in which the lengths of the line-elements PP′ and PP″ do not change, i.e., jjdxjj=jjdXjj; jjdxjj=jjdXjj, Eq. (4.39) leads to E12 ¼ ðp=2 hÞ=2
ð4:40Þ
Consequently, E12 describes half of the decrease in the angle contained by the two line-elements which were perpendicular before deformation. In addition to the Lagrangian and Eulerian strain tensors defined above, we can define various strain tensors in terms of U or V, fulfilling the condition that they are zero when U ¼ V ¼ I as follows (Seth 1964; Hill 1968): EðmÞ
8 < 1 ðU2m IÞ for m 6¼ 0 2m ¼ fðUÞ ¼ : ln U for m ¼ 0
ð4:41Þ
eðmÞ
8 < 1 ðV2m IÞ for m 6¼ 0 2m ¼ fðVÞ ¼ : ln V for m ¼ 0
ð4:42Þ
where 2m is the integer (positive or negative). The Green strain tensor is obtained by choosing m ¼ 1 in Eq. (4.41), i.e. E ¼ Eð1Þ and the Almansi strain tensor is obtained by choosing m ¼ 1 in Eq. (4.42), e ¼ eð1Þ . The Biot strain tensor (Biot 1965) is given by choosing m ¼ 1=2 in Eq. (4.41), i.e. BUI
ð4:43Þ
The generalized strain tensors in Eqs. (4.41) and (4.42) are mutually related by virtue of Eq. (4.5) as follows. EðmÞ ¼ RT eðmÞ R
ð4:44Þ
The strain tensors in Eqs. (4.41) and (4.42) are coaxial with U and V, respectively, and their principal values are given by the function of the principal stretch ka ða ¼ 1; 2; 3Þ as follows: 8 < 1 ðk2m 1Þ for m 6¼ 0 2m a f ðka Þ ¼ : ln ka for m ¼ 0
ð4:45Þ
124
4
Deformation/Rotation Tensors
The function f ðka Þ fulfills f ð1Þ ¼ 0; f 0 ð1Þ ¼ 1
ð4:46Þ
f 0 ðsÞ [ 0
ð4:47Þ
and
where s is an arbitrary positive scalar quantity. The function f ðka Þ for several values of m is shown in Fig. 4.2. (Note) Eq. (4.45)2 for m ¼ 0 is derived as follows: lim
1
m!0 m
ðkm a 1Þ ¼ lim
m!0
expðm ln ka Þ 1 expðm ln ka Þ ln ka ¼ ln ka ðno sumÞ ¼ lim m!0 1 m
by the aid of l’H^ ospital’s rule. Further, adopt the second-order tensor function fðUÞ which is coaxial with the right stretch tensor U and has the principal values f ðka Þ. Therefore, we can define the general strain tensor in the spectral decomposition as follows: 3 P
EðmÞ ¼
a¼1
f ðka ÞNðaÞ NðaÞ ¼
3 P a¼1
2m 1 2m ðka
1ÞNðaÞ NðaÞ
ð4:48Þ
In addition, for the left stretch tensor V, we can define the following strain tensor.
f(
m =1
)
m =1/2
1.5
1
m=0 m = 1/2 m= 1
0.5
0
0.5
1
2
0.5
1
Fig. 4.2 Relation of principal strain to principal stretch in generalized strain tensor
4.2 Strain Tensors
125
eðmÞ ¼
3 P
f ðka ÞnðaÞ nðaÞ ¼
a¼1
¼
3 X
3 P a¼1
2m 1 2m ðka
1ÞnðaÞ nðaÞ
f ðka ÞRNðaÞ RNðaÞ ¼ RfðUÞRT
ð4:49Þ
a¼1
noting Eq. (1.108). In the particular case of m ¼ 0, noting ka [ 0, the strains defined by the following equation are called the logarithmic or Hencky strain tensor. Lagrangian-logarithmic strain tensor: Eð0Þ ¼
3 P a¼1
Eulerian-logarithmic strain tensor: eð0Þ ¼
3 P a¼1
lnka NðaÞ NðaÞ ln U ¼ 12 ln C
lnka nðaÞ nðaÞ ln V ¼ 12 ln b
ð4:50Þ where ka ¼ Ua ¼ Va ¼
pffiffiffiffiffiffi pffiffiffiffiffi Ca ¼ ba
ð4:51Þ
Ua ; Ca ; Va and ba being the principal values of U; C; V and b, respectively. When the principal directions of U and V are fixed, the following equations hold in these directions. ka ¼
ð0Þ
ðE Þa ¼ ð e
ð0Þ
Þa ¼
@xa @Xa
@xa ðno sumÞ @Xa
@xa @Xa
ð4:52Þ
¼
@ xa ¼ ðln ka Þ ¼ daa @xa
ðno sumÞ ð4:53Þ
where ln ka in Eq. (4.52) is the logarithmic (or Hencky or natural or true) strain and daa (no sum) is the principal component of the strain rate tensor defined in the next section. It follows from Eq. (4.50) that 8 3 3 3 X X > 1X ð0Þ > > ¼ ln k ¼ ln C ¼ ln Ua ¼ lnðU1 U2 U3 Þ trE a a > < 2 a¼1 a¼1 a¼1 3 3 3 > X X > 1X > ð0Þ > ¼ ln k ¼ ln b ¼ ln Va ¼ lnðV1 V2 V3 Þ tre : a a 2 a¼1 a¼1 a¼1
ð4:54Þ
126
4
Deformation/Rotation Tensors
which is nothing but the logarithmic volumetric strain trEð0Þ ¼ treð0Þ ¼
3 P a¼1
4.3
ln
@xa v ¼ ln J ¼ ln ¼ ev @Xa V
ð4:55Þ
Volumetric and Isochoric Parts of Deformation Gradient Tensor
The deformation gradient tensor F is defined as the ratio of the length of the infinitesimal-line element in the current configuration to that of the reference configuration. Therefore, it may not be additively decomposed but it is obliged to be multiplicatively decomposed. The multiplicative decomposition of the deformation gradient into the volumetric part Fvol and the volume preserving, i.e. isochoric (distortional) part F will be described in the following, where F ¼ gFvol F; Fvol ðdetFÞ1=3 g ¼ FTvol ; F ðdetFÞ1=3 F
ð4:56Þ
noting F ¼ gF ¼ ½ðdetFÞ1=3 g ½ðdetFÞ1=3 F ¼ Fvol F F ¼ gi Gi ¼ gi dij Gj ¼ gi gi gj Gj ¼ gF based on Eq. (3.9), where it holds that detFvol ¼ detF ¼ J; detF ¼ 1
ð4:57Þ
noting detFvol ¼ det½ðdetFÞ1=3 g ¼ ½ðdetFÞ1=3 3 detg detF ¼ det½ðdetFÞ1=3 F ¼ ½ðdetFÞ1=3 3 detF ¼ ðdetFÞ1 detF by Eq. (1.143)2 with s ¼ ðdet FÞ1=3 . Then, the tensor F for the isochoric part is called a unimodular tensor. Then, let the following decomposition of the right Cauchy-Green deformation tensor be defined.
4.3 Volumetric and Isochoric Parts of Deformation Gradient Tensor
127
C ¼ FT Cvol F Cvol FTvol Fvol ¼ ðdetCÞ1=3 g;
ð4:58Þ
C FT F ¼ ðdetCÞ1=3 C satisfying detCvol ¼ detC ¼ J 2 ; detC ¼ 1
ð4:59Þ
F ¼ g; trC ¼ 3; C0 ¼ O for F ¼ Fvol
ð4:60Þ
F ¼ ðdetFvol Þ1=3 Fvol ¼ ðdetFvol Þ1=3 ðdetFvol Þ1=3 g ¼ g
ð4:61Þ
by Eq. (1.143)2 . One has
noting
ð4:62Þ
The following partial derivatives hold for the volumetric and the isochoric tensors.
ð4:63Þ
noting @Cvol @ðdetCÞ1=3 g 1 @detC ¼ ¼ g ðdetCÞ2=3 @C @C 3 @C 1 1 2=3 ðdetCÞC1 ¼ g ðdetCÞ1=3 C1 ¼ g ðdetCÞ 3 3
128
4
@C ¼ @C
h i @ ðdetCÞ1=3 C @C
Deformation/Rotation Tensors
1 ¼ C ðdetCÞ4=3 ðdetCÞC1 þ ðdetCÞ1=3 S 3
pffiffiffiffiffiffiffiffiffiffi @ detC 1 @detC 1 ¼ pffiffiffiffiffiffiffiffiffiffi ¼ pffiffiffiffiffiffiffiffiffiffi ðdetCÞC1 @C @C 2 detC 2 detC pffiffiffiffiffiffiffiffiffiffi @ ln detC 1 1 ¼ C @C 2
h i 1=3 trC @trC @ ðdetCÞ 1 ¼ ðdetCÞ4=3 ðdetCÞC1 trC þ ðdetCÞ1=3 G ¼ @C @C 3 by virtue of Eqs. (1.355), (1.357), (1.369) and (4.58). The expression in terms of the principal stretches ka ða ¼ 1; 2; 3Þ will be shown below. trF ¼ k1 þ k2 þ k3 ; detF ¼ k1 k2 k3
ð4:64Þ
trC ¼ k21 þ k22 þ k23 ; detC ¼ ðk1 k2 k3 Þ2
ð4:65Þ
The principal stretch ka is described in terms of the volumetric part kvol and the isochoric part ka as follows: ka ¼ kvol ka
ð4:66Þ
kvol ðk1 k2 k3 Þ1=3 ¼ J 1=3 ; ka ka =J 1=3
ð4:67Þ
where
trF ¼ k1 þ k2 þ k3 ¼ ðk1 þ k2 þ k3 Þ=J 1=3 ; detF ¼ k1 k2 k3 ¼ k1 k2 k3 Þ=J ¼ 1 ð4:68Þ trC ¼ k21 þ k22 þ k23 ¼ ðk21 þ k22 þ k23 Þ=J 3=3 ; detC ¼ ðk1 k2 k3 Þ2 ¼ 1
ð4:69Þ
The invariants of C are defined by 9 IC trC ¼ k21 þ k22 þ k23 > > > = 1 2 2 2 2 2 2 2 2 IIC ðtr C trC Þ ¼ k1 k2 þ k2 k3 þ k3 k1 > 2 > > ; 2 2 2 IIIC detC ¼ k1 k2 k3
ð4:70Þ
4.3 Volumetric and Isochoric Parts of Deformation Gradient Tensor
with
9 @IC @ trC > > ¼ ¼ I > > @C @C > > = 2 1 2 @ 2 ðtr C tr C Þ @IIC ¼ ¼ IC I C > @C @C > > > > @IIIC @detC 2 1 > ¼ ¼ IIC G IC C þ C ¼ IIIC C ; @C @C
129
ð4:71Þ
noting Eqs. (1.354) and (1.355). The invariants of the isochoric part are given by 9 > IC trC ¼ k21 þ k22 þ k23 > > = 1 2 2 2 2 2 2 2 2 IIC ðtr C trC Þ ¼ k1 k2 þ k2 k3 þ k3 k1 > 2 > > ; 2 2 2 IIIC detC ¼ k1 k2 k3
ð4:72Þ
noting tr2 C ¼ ðk21 þ k22 þ k23 Þ2 ; trC2 ¼ k41 þ k42 þ k43 , and their derivatives are given by 9 @IC @trC > > ¼ ¼I > > > @C @C > > > > = 2 1 2 @IIC @ 2 ðtr C trC Þ ð4:73Þ ¼ ¼ IC I C > @C @C > > > > > > @IIIC @detC 2 1 > ; ¼ ¼ II C G IC C þ C ¼ III C C > @C @C
4.4
Strain Rate and Spin Tensors
The idealized deformation process in which the deformation is uniquely determined by the state of stress independent of the loading path is called the elastic deformation process. To describe it, it is required to introduce the strain tensor describing the deformation from the initial state and relate it to the stress. Here, since the superposition rule does not hold in the strain tensor, the null stress state is chosen usually as the reference state of strain. On the other hand, the deformation is not determined uniquely by the state of stress depending on the loading path and thus it cannot be related to the stress in the irreversible deformation process, e.g. the viscoelastic, the plastic and the viscoplastic loading processes. Therefore, it is obligatory to relate the infinitesimal changes of stress and deformation and to integrate them along the loading path in order to capture the current states of stress and deformation.
130
4
Deformation/Rotation Tensors
Here, introduce the velocity gradient tensor in Eq. (2.30), i.e. l
@v @vi ; lij @ j vi @xj @x
ð4:74Þ
Noting F ¼ @ x =@X ¼ @v=@X ðdv ¼ F dXÞ and the chain rule of derivative, Eq. (4.74) can be rewritten as Eq. (2.30), i.e l¼
@v @X 1 @vi @XA ¼ F F ðF ¼ lFÞ; lij ¼ @XA @xj @X @x ðdxÞ ¼ ldx
ð4:75Þ ð4:76Þ
noting ðdxÞ ¼ dv ¼
@v dx @x
ðdxÞ , i.e. dv describes the rate of the infinitesimal line element, i.e. the relative velocity between the velocities of material particles in both sides of infinitesimal line element dx. Here, we can choose the time s ð tÞ to be arbitrary, resulting in
l ¼ s FðtÞs F1 ðtÞ because the velocity gradient tensor l is substantially independent of the reference infinitesimal line element dX but dependent only on rates of deformation and rotation relative to a deformed configuration. Now, choosing the current state for the reference state leading to t F1 ðtÞ ¼ I, the velocity gradient tensor l can be expressed in the updated Lagrangian description as follows:
l ¼ t FðtÞ
ð4:77Þ
Further, taking the time-derivative of Eq. (4.24) and noting t RðtÞ ¼ t UðtÞ ¼ ¼ I, it follows that
t VðtÞ
t F ðtÞ
¼ t UðtÞ þ t RðtÞ ¼ tV þ t RðtÞ
ð4:78Þ
Decomposing l additively into the symmetric and the skew-symmetric parts and noting Eqs. (4.75)–(4.78), it is obtained that l ¼ dþw
ð4:79Þ
4.4 Strain Rate and Spin Tensors
131
where 9 " T # > 1 @v @v > d 12 ðl þ l Þ ¼ þ ¼ t UðtÞ ¼ t VðtÞ > > > = 2 @x @x T
> > > > > ;
1 @vi @vj dij þ 2 @xj @xi 9 " T # > 1 @v @v > w 12 ðl l Þ ¼ ¼ t RðtÞ > > > = 2 @x @x
ð4:80Þ
T
wij
1 @vi @vj 2 @xj @xi
> > > > > ;
ð4:81Þ
where d is called the strain rate tensor or deformation rate tensor or stretching and w is called the (continuum) rotation rate tensor or continuum spin tensor. Here, note that d is not a time-derivative of any strain tensor but is defined independently as the rate variable although it is called the strain rate tensor. In addition, note that the time-integration of d cannot lead to any deformation measure in general, because it concerns with different material line-elements which rotate with material. Only the time-integration of axial component of d coincides with the axial component of the Hencky strain in Eq. (4.50) if the axial direction is fixed. Substituting Eqs. (4.1) and (4.75) into Eqs. (4.80) and (4.81), d and w are described by U; R as follows: 9 1 1 1 1 T T > 1 T d ¼ ½ F F þ ðF F Þ ¼ fðRUÞ ðRUÞ þ ðRUÞ ½ðRUÞ g > > > > 2 2 > > > 1 > 1 1 T > = ¼ RðU U þ U U ÞR 2 1 1 > > w ¼ ½F F1 ðF F1 ÞT ¼ fðRUÞ ðRUÞ1 ðRUÞT ½ðRUÞ T g > > > 2 2 > > > > 1 > T 1 1 T ; ¼ RR þ RðUU U U ÞR 2
ð4:82Þ
Consequently, we obtain
~ RT d ¼ Rsym½U
9 > =
> ~ RT ; w ¼ X þ Rant½U R
ð4:83Þ
132
4
Deformation/Rotation Tensors
where
e U U1 U
XR R R T
ð4:84Þ ð4:85Þ
XR is the relative spin appeared in Eq. (3.38). The following relation holds in the pure rotation ðFð¼ @x=@XÞ ¼ R; dx ¼ RdXÞ.
ðdxÞ ¼ Rd X ¼ R RT dx ¼ XR dx ¼ x dx
ð4:86Þ
Therefore, XR designates the spin of the line element during the pure rotation, while x is the axial vector to XR . Further, d and w are described by V; R as follows: 9 1 > 1 T T > d ¼ fðVRÞ ðVRÞ þ ðVRÞ ½ðVRÞ g > > > 2 > > > 1 1 1 T T T T 1 T T > = ¼ ðV V þ V V Þ þ ðV R R V V R R V Þ > 2 2 1 > > > w ¼ fðVRÞ ðVRÞ1 ðVRÞT ½ðVRÞ T g > > 2 > > > 1 1 T 1 T T T 1 T T > ; ¼ ðV R R V þ V R R V Þ þ ðV V V V Þ > 2 2 and thus
9 R > ~ ~ d ¼ sym½V þ sym½X = > ~ R þ ant½V ~ ; w ¼ ant½X
ð4:87Þ
ð4:88Þ
where
~ V V1 V
ð4:89Þ
~ R VXR V1 X
ð4:90Þ
Equations (4.83) and (4.88) mean that d is not a pure rate of deformation and w is not a pure rate of rotation. It follows from Eq. (4.12) that L
XL R RLT ¼
3 X a;b¼1
ðNðaÞ ea Þ ðNðbÞ eb ÞT
4.4 Strain Rate and Spin Tensors
¼
3 X
133
ðN ðaÞ eðaÞ þ NðaÞ eðaÞ Þ eðbÞ NðbÞ
a;b¼1
¼
3 X
ðaÞ
N
NðaÞ
ð4:91Þ
a¼1
and thus one has
ðaÞ
N
¼ XL NðaÞ
ð4:92Þ
ðaÞ
noting e ¼ 0 since feðaÞ g is the fixed base. Therefore, XL describes the spin of the Lagrangian principal triad fNðaÞ g of the right stretch tensor U and is called the Lagrangian spin tensor. The components of XL in the Lagrangian triad fNðaÞ g are described as ðbÞ
XLab ¼ NðaÞ XL NðbÞ ¼ NðaÞ N
ð4:93Þ
On the other hand, it follows from Eq. (4.12) that E
XE R RET ¼
3 X
T nðaÞ eðaÞ nðbÞ eðbÞ
a;b¼1
¼
3 X nðaÞ nðaÞ
ð4:94Þ
a¼1
and thus one has ðaÞ
n
¼ XE nðaÞ
ð4:95Þ
Therefore, XE describes the spin of the Eulerian principal triad fnðaÞ g of the right stretch tensor V and is called the Eulerian spin tensor. The components of XE in the Eulerian triad fnðaÞ g are described as
XEab ¼ nðaÞ XE nðbÞ ¼ nðaÞ n ðbÞ
ð4:96Þ
It follows from Eq. (4.9) that
F ¼
3 X a¼1
(
) a a a ðaÞ k a n N þ ka n N þ n N
a
a
ð4:97Þ
134
4
Deformation/Rotation Tensors
which is rewritten by Eqs. (4.92) and (4.95) as
F ¼
3 X
ka nðaÞ NðaÞ þ XE F FXL
ð4:98Þ
a¼1
or by Eqs. (4.93) and (4.96), Eq. (4.97) leads to
F ¼
3 X
ka nðaÞ NðaÞ þ
a¼1
3 X
kb XEab ka XLab nðaÞ NðbÞ
ð4:99Þ
a;b¼1
Here, it holds that
R RT ¼
3 X nðaÞ NðaÞ
T nðbÞ NðbÞ
a;b¼1
¼
3 X ðaÞ NðbÞ nðbÞ n ðaÞ NðaÞ þ nðaÞ N a;b¼1
3 X ðcÞ ðaÞ ðaÞ ðaÞ ðaÞ ðcÞ ðbÞ ðbÞ N n ¼ n n þn N N N a;b;c¼1
¼
ðcÞ 3 X ðaÞ n nðaÞ nðaÞ NðaÞ N NðcÞ NðbÞ nðbÞ a;b;c¼1
and thus the following relations hold. XR ¼ XE RXL RT ; XE ¼ XR þ RXL RT ;
XL ¼ R T XE XR R
ð4:100Þ
The velocity gradient is described noting Eq. (4.9) as l¼
3 X a¼1
¼
! ðaÞ
ka n
N
ðaÞ
3 X 1 ðbÞ N nðbÞ k b¼1 b
3 3 X ðaÞ X 1 ðbÞ ðaÞ ka nðaÞ NðaÞ þ ka n NðaÞ þ ka nðaÞ N N nðbÞ k b a¼1 b¼1
0 1 3 X ka @ka nðaÞ nðaÞ þ nðaÞ nðaÞ A þ ¼ N ðaÞ NðbÞ nðaÞ nðbÞ ka k a¼1 a;b¼1 b 3 X
ð4:101Þ
4.4 Strain Rate and Spin Tensors
135
which is rewritten using Eqs. (4.93) and (4.96) as
l¼
3 X ka a¼1
ka
n
ðaÞ
ðaÞ
n
3 X ka L E þ Xa;b Xab nðaÞ nðbÞ kb a;b¼1
ð4:102Þ
from which the strain rate and the continuum spin are represented as
3 X k
d¼
a¼1
w¼
a ðaÞ
ka
3 X a;b¼1
nðaÞ þ
3 k2 k2 X a b
XLab nðaÞ nðbÞ
ð4:103Þ
! 2 2 k þ k a b XEab XL nðaÞ nðbÞ ða 6¼ bÞ 2ka kb ab
ð4:104Þ
n
a;b1
2ka kb
noting
d¼
3 X ka a¼1
2 sym
ka
ðaÞ
n
" 3 X
n
ðaÞ
3 k 1X ka ðaÞ ðbÞ b ðbÞ ðaÞ nðaÞ nðbÞ þ N N N N þ ka 2 a;b¼1 kb
n ðaÞ nðaÞ ¼
a¼1
w¼
3 1X 2 a¼1
þ
3 X n ðaÞ nðaÞ þ nðaÞ nðaÞ ¼
3 X
a¼1
a¼1
! nðaÞ nðaÞ
¼I¼O
!
n ðaÞ nðaÞ nðaÞ n ðaÞ 3 1X ka ðaÞ kb ðbÞ ðaÞ nðaÞ nðbÞ N NðbÞ N N ka 2 a;b¼1 kb
From the relation dab ¼
k2a k2b L X ða 6¼ bÞ 2ka kb ab
ð4:105Þ
obtained from the anti-symmetric part in Eq. (4.103), the Lagrangian spin is written as XLab ¼
2ka kb dab ða 6¼ bÞ k2b k2a
ð4:106Þ
The Eulerian spin is given from Eq. (4.104) as follows: XEa;b ¼ wab
k2a þ k2b k2a k2b
dab ða 6¼ bÞ
ð4:107Þ
136
4
Deformation/Rotation Tensors ðaÞ
~ ¼V¼V ~ ¼ O; N In the rigid-body rotation (U ¼ U Eqs. (4.79), (4.83), (4.91) and (4.100) that l ¼ w ¼ XR ¼ X E ;
¼ O), it follows from
XL ¼ O
ð4:108Þ
In what follows, we consider the physical meanings of d and w. The relative velocity of the particle points P and P0 , the current position vectors of which are x and x þ dx, respectively, is given by dv ¼ ldx
ð4:109Þ
from Eq. (4.76) and it is additively decomposed as dv ¼ dvd þ dvw
ð4:110Þ
dvd ddx
ð4:111Þ
dvw wdx
ð4:112Þ
where
The following equation is obtained for the infinitesimal line-element dx ¼ dxi ei ðno sumÞ. dj€i ¼
dvdj (no sum) dxi
ð4:113Þ
noting dji ¼ ej dei ¼ ej d
dx dvd dvdj ¼ ej ¼ dxi dxi dxi
ð4:114Þ
with the aid of Eqs. (1.107) and (4.111). Therefore, dji is the ej -component of the relative velocity dvd of the unit line element ðdxi ¼ 1Þ in the ei -direction. Consequently, the infinitesimal line-element dx ¼ dxi ei ðno sumÞ rotates in the velocity given by the tangential component dji ðj 6¼ iÞ of the strain rate d. On the other hand, denoting the axial vector described in Eqs. (1.187) and (1.188) _ for the skew-symmetric tensor w by x, it holds that 1 _ xi ¼ eirs wrs ; 2
_
wij ¼ eijr xr
and thus Eq. (4.112) is rewritten by Eq. (1.190) as
ð4:115Þ
4.4 Strain Rate and Spin Tensors
137
_ _ dvwi ¼ wis dxs ¼ eisr xr dxs ¼ eirs xr dxs
_
dvw ¼ x dx;
ð4:116Þ
Therefore, the arbitrary line-element dx rotates in the peripheral velocity dvw _ and angular velocity x, termed often the spin vector, by the continuum spin w, _ whereas 2x is called the vorticity. Noting Eqs. (4.110)–(4.116), it follows for the material line element dx1 e1 that dv ¼ lij ei ej dx1 e1 ¼ li1 dx1 ei ¼ ðd21 þ w21 Þdx1 e2 þ d11 dx1 e1 ¼ ðd12 w12 Þdx1 e2 þ d11 dx1 e1 ¼ -3 dx1 e2 þ d11 dx1 e1
ð4:117Þ
as shown in Fig. 4.3, where we set ^ 3 þ d12 -3 w12 þ d12 ¼ x
ð4:118Þ _
-3 designates the clock-wise angular velocity of the line-element. x 3 denotes the average angular velocity of the line-elements in the plane, which coincides with the angular velocity of the line element in the principal directions of strain rate fulfilling d12 ¼ 0. By choosing dn ; dt for Tn ; Tt described in Sect. 1.14, the relation of the rate of elongation and the rate of rotation is shown in Fig. 4.4. It is depicted by the qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 centering in circle of relative velocity with the radius ½ðd11 þ d22 Þ=22 þ d12 _
ððd11 þ d22 Þ=2; x3 Þ in the two-dimensional plane ðdn ; -3 Þ. The parallelepiped in the _ principal directions of the strain rate d rotates by the angular velocity x as shown in Fig. 4.5. The rate of the scalar product of the vectors dx and dx of the infinitesimal elements connecting the three points P; P0 ; P00 with the position vectors x; x þ dx; x þ dx, respectively, is given noting Eq. (1.139) as follows:
(
dv w = 3 dx1e2 = w12 dx1e2
x2 dv (
3
dv d
d12
e2
0
e1
dx = dx1e1
Fig. 4.3 Extension and rotation of the line-element
dv2d = d12dx1e 2
x1 dv1d = d11dx1e1
138
4 3
=
3
Deformation/Rotation Tensors
dt = w12 dt (d 11,
3
3
d12)
= w12
(d 22 ,
3
0
d12 )
d2
d1
d11 d 22 2
dn
Fig. 4.4 Circle of relative velocity
(= )
d 3 dx3 x3
d 2 dx2 x2
d1dx1
x1 Fig. 4.5 Deformation and rotation for principal directions of strain rate
ðdx dxÞ ¼ dv dx þ dx dv @v @v dx dx þ dx dx ¼ ¼ @x @x
("
T # ) @v @v þ dx dx @x @x
leading to ðdx dxÞ ¼ 2ddx dx
ð4:119Þ
If the vicinity of the particle P undergoes the rigid-body rotation, the quantity in Eq. (4.119) for an arbitrary scalar quantity dx dx is zero and thus d ¼ O has to
4.4 Strain Rate and Spin Tensors
139
hold. Inversely, if d ¼ O, the quantity in Eq. (4.119) for the scalar quantity dx dx of arbitrary line-element vectors becomes zero and thus it can be stated that the vicinity of the particle P does not undergo a deformation. Then, d ¼ O is the necessary and the sufficient condition for the situation that a deformation is not induced, allowing only a rigid-body rotation. Denoting the lengths of the line-elements PP0 and PP00 as ds and ds, respectively, and the angle contained by them as h, it holds that ðdx dxÞ ¼ ðdsdscos hÞ ¼
ðdsÞ ðdsÞ þ ds ds
cos h hsin h dsds
ð4:120Þ
Further, denoting the unit vectors in the directions of the line-elements PP0 and PP as l and m, respectively, and noting dx ¼ lds, dx ¼ mds, it holds from Eqs. (4.119) and (4.120) that 00
ðdsÞ ðdsÞ þ ds ds
cos h hsin h ¼ 2dl v ¼ 2dij li vj
ð4:121Þ
If the particles P0 and P00 chosen in same direction ðh ¼ 0Þ, it follows from Eq. (4.121) that ðdsÞ ¼ dl l ds
ð4:122Þ
The left-hand side of Eq. (4.122) designates the rate of elongation of the line-element. Therefore, the rate of elongation is given by the normal component of d in the relevant direction, noting Eq. (1.107). On the other hand, choosing the line-element PP00 to be perpendicular to the line element PP0 ðh ¼ p=2Þ, it follows from Eq. (4.121) that h ¼ 2dl m
ðl m ¼ 0Þ
ð4:123Þ
The left-hand side of Eq. (4.123) designates the decreasing rate of the angle contained by the two line-elements mutually perpendicular instantaneously and is called the shear strain rate. Next, the relations of the rate E of Green strain tensor E and the rate e of the Almansi strain tensor e to the strain rate tensor d are formulated below. The material-time derivative of Eq. (4.26) is given by
ðdx dxÞ ¼ 2 E dX dX
ð4:124Þ
140
4
It follows from Eqs. (4.119) and (4.124) that
Deformation/Rotation Tensors
:
ddx dx ¼ dFdX FdX ¼ E dX dX
ð4:125Þ
from which, noting Eq. (1.139), we have the relation of Green strain tensor to the strain rate tensor as follows:
1
E ¼ FT dF; d ¼ FT E F
ð4:126Þ
which is obtained also from
ð4:127Þ
Next, the time-differentiation of Eq. (4.27)2 leads to ð4:128Þ
Here, it follows from ðFF1 Þ ¼ FðF1 Þ þ F F1 ¼ O with Eq. (4.75) that
F1
¼ F1 l
ð4:129Þ
which is derived also by 1 @F1 @F1 @ @X @ @X þ v¼ F ¼ þ v @t @x @t @x @x @x @ @X @ @X @X @v @ @X @X @X @v ¼ þ v ¼ þ v @x @t @x @x @x @x @x @t @x @x @x ¼
@ @X @v @X @v X ¼ @x @x @x @x @x
ð4:130Þ
noting that the inside of the bracket ½ in Eq. (4.130) is the material-time derivative of the initial position vector X and thus it is zero.
4.4 Strain Rate and Spin Tensors
141
Substituting Eq. (4.129) into Eq. (4.128), one has
i 1 h 1 T 1 F l F þ FT F1 l 2 1 1 1 T 1 T 1 T 1 I IF F I IF F l ¼l þ 2 2 2 2 1 1 ¼ lT Ie þ Ie l 2 2
e ¼
ð4:131Þ
from which one has the relation of the rate of the Almansi strain tensor to the strain rate tensor:
e ¼ d lT e el
ð4:132Þ
Equation (4.132) is rewritten as
e ¼d
1 1 l þ lT l lT e e l þ lT þ l lT 2 2
and thus we obtain
w
e ¼ d de ed
ð4:133Þ
where
ew e we þ ew
ð4:134Þ
is called the Zaremba-Jaumann rate of Almansi strain tensor, while the Zaremba-Jaumann rate will be explained in Sect. 4.4.
E ¼ e ¼ e w ¼ d holds in the initial state ðF ¼ I; E ¼ e ¼ OÞ and thus all the strain rates mutually coincide by Eqs. (4.126), (4.127), (4.132) and (4.134).
4.5
Logarithmic (True) and Infinitesimal (Nominal) Strains
In what follows, we introduce the following notation.
e d; de ddt
ð4:135Þ
142
4
Deformation/Rotation Tensors
If the direction of the material line-element always coincides with the xi -axis, the principal strain rate in this direction is given by
di ¼ e i ¼
@vi @ui @ui @ui @ui (no sum) ¼ ¼ ¼ = 1þ @xi @xi @ ðXi þ ui Þ @Xi @Xi
ð4:136Þ
The time-integration of Eq. (4.136) leads to @ui @xi ð0Þ ð0Þ ¼ ln ei ¼ ln 1 þ ¼ ln ki ¼ Ei ¼ ei ¼ lnð1 þ ei Þ (no sum) @Xi @Xi ð4:137Þ Therefore, the time-integration ei of principal strain rate di does not coincide with the principal infinitesimal strain ei in Eq. (4.32). Setting @Xi ! l0 ; @xi ! l; @ui ! l l0 , where l0 and l are the lengths of the line-element in the initial and the current states, respectively, it follows that 9 l > ¼ ln ð 1 þ e Þ i = l0 l l0 > ; ei ¼ 0 l ei ¼ ln
ð4:138Þ
Note that the logarithimic axial strain and the infinitesimal strain are denoted by the italic letter ei and the roman letter ei , respectively, while it needs the caution to distinguish them. Consequently, the time-integration of principal component of strain rate tensor d and the Hencky strain tensor Eð0Þ and eð0Þ coincide with ei which is called the logarithmic (or natural) strain, provided that their principal directions are fixed. On the other hand, the principal value of infinitesimal strain tensor e does not coincide with them and it is called the nominal strain. It follows from Eq. (4.138) that ei ¼ 1ne 2 ðffi 0:693Þ for l ¼ 2l0 and ei ¼ 1 for l ¼ 0 ei ¼ þ 1 for l ¼ 2l0 and ei ¼ 1 for l ¼ 0
) ð4:139Þ
Therefore, the magnitude of nominal strain in the deformation that the material length becomes zero, i.e. the material diminishes is identical with that in the deformation that the material length becomes only twice. As a practical example, about 5% error is induced in the nominal strain for 10% elongation as known from ei =ei ¼ 0:1= lnð1:1Þ ¼ 1:049 for l ¼ 1:1 l0 . This property would cause the inconvenience for the adoption in constitutive equation for the wide range of deformation.
4.5 Logarithmic (True) and Infinitesimal (Nominal) Strains
143
Further, one has Z
ln l0
dl ¼ l
Z
l1 l0
dl þ l
Z
l2 l1
dl þ þ l
Z
ln
dl ln1 l
i.e. ln
ln l1 l2 ln ¼ ln þ ln þ þ ln l0 l0 l1 ln1
On the other hand, one sees ln l0 l1 l0 l2 l1 ln ln1 ¼ 6 þ þ þ l0 l0 l1 ln1 Thus, it follows for the superposition of strains that n e0i n ¼ e0i 1 þ e1i 2 þ þ en1 i
) ð4:140Þ
n e0i n 6¼ e0i 1 þ e0i 2 þ þ en1 i
while eai b and eai b designate the longitudinal strain in the xi -direction when the length of the line-element changes from la to lb , provided that the principal direction of strains are fixed. Consequently, the superposition rule holds in the logarithmic strain but it does not hold in the nominal strain. Furthermore, one has ln
l1 l2 l3 l1 l2 l3 ¼ ln 0 þ ln 0 þ ln 0 l01 l02 l03 l1 l2 l3
l1 l2 l3 l01 l02 l03 l1 l01 l2 l02 l3 l03 6¼ þ þ l01 l02 l03 l01 l02 l03 where l1 ; l2 ; l3 are the lengths of line-elements in the directions of three fixed principal strains. Thus, it follows for the sum of the principal strains that ev ¼
ln Vv
9 > > ¼ ln J ¼ ln F ¼ ei ¼ lnð1 þ ev Þ > > = i¼1
3 vV X 6¼ ei ev ¼ V i¼1
3 P
> > > > ;
ð4:141Þ
where V and v are the initial and the current volumes, respectively, of material. Therefore, the sum of logarithmic strains in orthogonal directions coincides with the logarithmic volumetric strain but the sum of nominal strains in orthogonal
144
4
Deformation/Rotation Tensors
directions does not coincide with the nominal volumetric strain. Equation (2.21) is exploited in Eq. (4.141). Further, it follows from Eqs. (4.136) and (4.141) that
ev ¼ dv ¼ trd ¼
v v @vr J ¼ ¼ 6¼ e v ¼ @xr v J V
ð4:142Þ
Therefore, the volumetric strain rate trd coincides with the material-time derivative of the logarithmic volumetric strain em but it does not coincide with that of the nominal volumetric strain em . Consequently, the nominal strain is applicable only to the description of infinitesimal deformation, while the logarithmic strain is applicable to constitutive equations for finite deformation.
Chapter 5
Stress Tensors and Conservation Laws
Conservation laws of mass, momentum, angular momentum, etc. must be fulfilled during a deformation. These laws are described first in detail. Then, the Cauchy stress tensor is defined and further, based on it, various stress tensors are derived from the Cauchy stress tensor. Introducing the stress tensor, the equilibrium equations of force and moment are formulated from the conservation laws. The virtual work principle required for the analyses of boundary value problems are also described. Further, the work-conjugate pairs of the stresses and the deformation rate tensors are explained, which must be adopted in the formulation of constitutive equations.
5.1
Stress Tensor
When the infinitesimal force vector df applies to the surface with the infinitesimal area da and the unit normal vector n, the stress vector, i.e. traction t is given as t
df da
ð5:1Þ
Now, introduce the following second-order tensor r fulfilling the relation t ¼ rn; ti ¼ rij nj
ð5:2Þ
by the quotient law described in Sect. 1.3.2. The following relation holds from Eqs. (5.1) and (5.2) df ¼ rnda; dfi ¼ rij nj da
ð5:3Þ
Equation (5.2) is called the Cauchy’s fundamental theorem or Cauchy’s stress principle. © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_5
145
146
5
Stress Tensors and Conservation Laws
The components of the tensor r are given by Eq. (1.107) as rij ¼ ei rej ; r ¼ rij ei ej
ð5:4Þ
As will be verified by the equilibrium of the angular moment in Sect. 5.6 that r is the symmetric tensor, and thus one has rij ¼ rei ej
ð5:5Þ
from Eq. (5.4). Here, when we choose ei to the unit normal vector n of the surface on which t applies, the following equation holds by substituting Eq. (5.2) with n ¼ ei , i.e. t ¼ rn ¼ rei into Eq. (5.5). rij ¼ ti ej
ð5:6Þ
where the expression ti specifies the stress vector applying to the unit surface possessing the unit normal vector ei ð¼ nÞ. Therefore, rij can be interpreted as the component in the direction ej of the stress vector ti applying to the unit surface element having the outward-normal vector ei . The tensor r is called the Cauchy stress tensor. The six independent components of r exist for the orthonormal surface base vectors ðe1 ; e2 ; e3 Þ by the symmetry, which is described in the matrix form as 2 4
r11
3 r12 r13 r22 r23 5 Sym: r33
The Cauchy’s fundamental theorem and the components of the Cauchy stress tensor are illustrated in Fig. 5.1. The physical meaning of the stress tensor in the general coordinate system is explained comprehensively in Hashiguchi (2020). Various stress tensors are defined from the Cauchy stress tensor described above. Some of them, which are often used in continuum mechanics, are presented below. The tensor s defined by the following equation is called the Kirchhoff stress tensor. s ¼ Jr
ð5:7Þ
The vector T defined by the following equation is called the nominal stress vector. T
df dA
ð5:8Þ
5.1 Stress Tensor
147
n = e3 33
df = tda
n
t3
da 31
σda
13
32
23
t2
t1 22 12
11
n = e2
21
n = e1 t = σn
ij
(a) Cauchy’s fundamental theorem
ti e j
ji
(b) Components of Cauchy stress
Fig. 5.1 Cauchy’s fundamental theorem and components of Cauchy stress tensor
The stress tensor P, which is related to T by the following equation, is called the first Piola-Kirchhoff stress tensor which is the Eulerian-Lagrangian two-point tensor. T PN ðTi PiA NA Þ
ð5:9Þ
where N is the unit outward-normal vector on the surface in the reference configuration. Here, substituting Eqs. (2.24) and (5.9) into Eq. (5.8), we have 1 1 df PNdA ¼ PFT nda; dfi ¼ PiA NA dA ¼ PiA FrA nr da J J
ð5:10Þ
On the other hand, the substitution of Eq. (5.2) into Eq. (5.1) yields df ¼ rnda
ð5:11Þ
It follows from Eqs. (5.10) and (5.11) that 1 df ¼ rnda ¼ PFT nda J
ð5:12Þ
The relation of r and P is given from Eq. (5.12) as follows: 9 1 1 1 1 r ¼ FPT ð¼ PFT Þ; rij ¼ FiA PjA ¼ PiA FjA = J J J J ; T T T P ¼ JrF ¼ sF ð6¼ P Þ; PiA ¼ Jrir ðF1 ÞAr ^g P ¼ Pij gi Gj sij gi Gj ¼ s ^gG ¼ s G ¼ sFT
ð5:13Þ
148
5
df = (σda) n
df = (P dA)N N
n
F
dA P (= FS)
Stress Tensors and Conservation Laws
σ da
dX dx = F dX
dv
dV
Fig. 5.2 Cauchy stress and first Piola-Kirchhoff stress
Equation (5.12) is illustrated in Fig. 5.2. The stress tensor P is called the first Piola-Kirchhoff stress tensor, noting PN ¼ df=dA applied to the reference unit area vector. Incidentally, it is the pull-back of the Cauchy stress tensor only from the right in Eq. (3.25)2 and thus it is the contravariant two-point tensor with the covariant current and reference base vectors. The first Piola-Kirchhoff stress is called also the nominal stress while the Cauchy stress is called the true stress, since the former designates the force per the initial (reference) unit surface while the latter designates the force per the current unit surface. It is used when the rate form of the equilibrium equation in the current configuration is derived as will be shown in Sect. 5.5. Further, the stress tensor S defined by the following equation is called the second Piola-Kirchhoff stress tensor which is the Lagrangian tensor. T G ¼ SN
ð5:14Þ
where TG
F1 df ¼ F1 T dA
ð5:15Þ
is the contravariant pull-back of the nominal stress vector T. Using Eq. (2.24) into these equations, one has the following expression. 1 df FSNdA ¼ FSFT nda J
ð5:16Þ
Comparing Eqs. (5.11) and (5.16), it follows that 1 1 r ¼ FSFT ; rij ¼ FiA SAB FBj J J 1 1 S ¼ F P ¼ JF rF T ¼ F1 sF T ð¼ST Þ;
9 = SAB ¼ JðF1 ÞAi rij ðF 1 ÞBj
;
ð5:17Þ
5.1 Stress Tensor
149
S ¼ Sij Gi Gj ¼F1 P ¼ F1 Pij gi Gj ¼ sij Gi Gj ¼ F1 sFT ¼ s GG
S is the full pull-back of the Cauchy stress tensor, and thus is based in the covariant reference base vectors. (Note) F1 df ¼ ð@XA =@xi ÞeA ei dfj ej ¼ ð@XA =@xi Þdfi eA is the pull-back of df from the current to the reference configurations. One can pull-back df also by FT df as shown in Eq. (3.17). However, the stress tensor R ¼ FT rFT =J obtained by setting FT df=dA ¼ RN instead of Eq. (5.14) does not satisfy the symmetry, i.e. R 6¼ RT and thus it is inconvenient for the practical use of deformation analysis. The substitution of the decomposition of the Cauchy stress into the deviatoric and the spherical part r ¼ r0 þ rm I; rm ð1=3Þtrr
ð5:18Þ
to Eq. (5.17)2 for the second Piola–Kirchhoff stress leads to S ¼ JF1 rFT ¼ JF1 r0 FT þ JF1 rm IFT i.e. S ¼ S0 þ JC1 rm
ð5:19Þ
S0 ¼ JF1 r0 FT
ð5:20Þ
trS0 6¼ 0; S0 :Cð¼ tr½ðJF1 rFT ÞðFT FÞT ¼ Jtrr0 Þ ¼ 0
ð5:21Þ
where
Here, note that
Now, introduce the Mandel stress M ¼ CS ¼ JFT rFT ¼ FT sFT ð6¼ MT Þ
ð5:22Þ
G M ¼ Mij Gi Gj si j Gi Gj ¼ FT sFT ¼ s G i.e. M ¼ M0 þ M m I
ð5:23Þ
150
5
Stress Tensors and Conservation Laws
where M0 ¼ JFT r0 FT ;
Mm ¼ Jrm
ð5:24Þ
Therefore, one has trM0 ¼ 0
ð5:25Þ
The Mandel stress is adopted in the multiplicative hyperelastic-based plasticity for the finite deformation as will be described in Chap. 17. The following stress is called the covariant convected stress tensor. cr
FT rF ð¼ c rT Þ;
cs
FT sF ð¼ c sT Þ
ð5:26Þ
The relations between the above-mentioned stress tensors as the transformations between the Lagrangian and the Eulerian tensors will be described in Sect. 5.9. The relations of various stress tensors defined above are summarized in Table 5.1. (Note) t
df df ;T ; da dA
TG
F1 df ¼ F1 T dA
t ¼ rn; Jt ¼ sn; T ¼ PN; T G ¼ SN
Table 5.1 Relations of various stress tensors Names, notations Cauchy r Kirchhoff s 1st Piola-Kirchhoff (Nominal) P 2nd Piola-Kirchhoff S Mandel M
rð¼ rT Þ
sð¼ sT Þ 1 s J
Jr
P ¼ 6 PT
S ¼ ST
M ¼ 6 MT
1 T FP J PFT
1 FSFT J FSFT
1 T F MFT J FT MFT
FS
FT M
JrFT
sFT
JF1 rFT
F1 sFT
F1 P
JFT rFT
FT sFT
FT P
C1 M CS
5.1 Stress Tensor
5.2
151
Conservation Law of Mass
Denoting the field of material density as qðx; tÞ, the mass in a current volume v is R given as m ¼ v qðx; tÞdv which is kept constant because the mass neither flow into the volume element nor flow out from it. Therefore, the following conservation law of mass must hold. Z
m¼
¼0
qðx; tÞdv
ð5:27Þ
v
from which, noting Eq. (2.50), one has the continuity equation.
q þ qdivv ¼ 0 ;
q þq
@vr ¼0 @xr
ð5:28Þ
Further, setting Tðx; tÞ q/, where / is a physical quantity per unit mass, one has Z
Z Z @vr dv ¼ q/ þ q/ þ q/ q /dv @xr v v
q/dv
¼
v
ð5:29Þ
noting the Reynolds’ transportation theorem in Eq. (2.48) and Eq. (5.28).
5.3
Conservation Law of Linear Momentum
R The linear momentum in a current volume v is given by v qvdv. On the other hand, R the traction applied to the surface of the region is given as a tda and the body force R applied to the region is given by v qbdv, denoting the body force per unit mass as b. The rate of momentum has to be equivalent to the sum of the traction and the body force applied to the region. Therefore, the Euler’s first law of motion (or conservation law of momentum) is given as Z
qvdv
Z
Z
¼
tda þ
v
qbdv
a
v
or R
v
by virtue of Eq. (5.29).
q v dv ¼
R a
tda þ
R v
qbdv
ð5:30Þ
152
5
5.4
Stress Tensors and Conservation Laws
Conservation Law of Angular Momentum
R The angular momentum in a current volume v is given as v qðx vÞdv. On the other hand, since the angular momentum caused by the tractionR and the angular momentum caused by the body force are described by a ðx tÞda and R v qðx bÞdv, respectively, the Euler’s second law of motion, i.e. conservation law of angular momentum is described as Z
qx vdv
Z
Z
¼
x tda þ
v
qx bdv
a
Z
qeijk xj vk dv
v
Z
Z
¼
eijk xj tk da þ
v
qeijk xj bk dv
a
v
which is reduced to R
R
v
qx v dv ¼
Z
a
Z
k
R v
qx bdv
Z
qeijk xj v dv ¼ v
x tda þ
eijk xj tk da þ a
ð5:31Þ
qeijk xj bk dv v
noting ðx vÞ ¼ v v þ x v ¼ x v and Eq. (5.29).
5.5
Equilibrium Equation
Substituting Eq. (5.2) into Eq. (5.30) for the conservation law of momentum and noting Eq. (5.45), the following equation is obtained. Z
Z
Z
q v dv ¼ v
rT nda þ a
Z
Z
q vi dv ¼
qbdv; v
v
Z rir nr da þ
a
qbi dv
ð5:32Þ
v
The first term in the right-hand side of Eq. (5.32) is given by the Gauss’ divergence theorem in Eq. (1.392) as Z Z @rir rir nr da ¼ dv a v @xr By this equation the local form of Eq. (5.32) is given as
rx r þ qb ¼ q v ;
@rij þ qbi ¼ q vi @xj
ð5:33Þ
5.5 Equilibrium Equation
153
This equation is called the Cauchy’s first law of motion, i.e. the equilibrium equation. On the other hand, substituting Eqs. (2.21) and (5.12) into Eq. (5.32), one has Z Z Z q0 v dV ¼ PNdA þ q0 bdV ð5:34Þ V
A
V
which is rewritten by the Gauss’ divergence theorem as follows: Z
Z
Z
q0 v dV ¼ V
PrX dV þ V
ð5:35Þ
q0 bdV V
where rX ð@=@XA ÞeA ¼ @=@X. The local form of this equation is given as
rX P þ q0 b ¼ q0 v ;
@PiA þ q 0 bi ¼ q 0 v i @XA
ð5:36Þ
The equilibrium equation in a rate form is required in constitutive equations for irreversible deformation including elastoplastic deformation. The time-differentiation of Eq. (5.36) engenders the following rate-type (or incremental-type) equilibrium equation, provided that the acceleration does not change, i.e. v ¼ 0.
P rX þ q0 b ¼ 0 ;
@ PiA þ q0 bi ¼ 0 @XA
ð5:37Þ
In order to describe Eq. (5.37) by the Cauchy stress, specifying
D 1J PFT Pr
D 1 6¼P r T ¼ FPT J
ð5:38Þ
and noting Eq. (5.13)2 , we have
By substituting
(due to Eq. (4.75)) and Eqs. (4.79) and (2.34) to this equation, we obtain D
P D
w
r ¼ r þ rtrl rlT ¼ r þ rtrd rd þ wr w
ð5:39Þ
Therein, P r is referred to as the nominal stress rate, whereas r r w r þ r w is the Zaremba-Jaumann stress rate, nothing Eq. (3.40).
154
5
Stress Tensors and Conservation Laws
It follows that
D
@P r 1 @ P ¼ @x J @X
ð5:40Þ
D @ FT =J @p r @ P FT =J 1@P T 1@ P ¼ F þP ¼ ¼ @x @x @x J @x J @X
ð5:41Þ
by virtue of
0
1 @ FjA =J 1 @PiA @xj 1 @PiA C B ij 1 @Pi;A ¼ Fj;A þ PiA ¼ ¼ @ A @xj @xj J @xj J @xj @XA J @XA D
@P r
with @ FT =J ¼O @x
ð5:42Þ
where O is the third-order zero tensor, noting " T # @ FT =J 1 @FT 1 @FT T @J T @ det F @F JF JF ¼ 2 ¼ 2 @x @x @x J @x J @F @x " # 1 @FT @F T J FT JFT ¼O ¼ 2 @x J @x Substitution of Eq. (5.40) into Eq. (5.37) yields the rate-type equilibrium defined in the current configuration:
D
P r rx þ q b ¼ 0 ;
D
@P rij þ qbi ¼ 0 @xj
ð5:43Þ
This equation is derived also by the following manner. From Eq. (5.32) with v ¼ 0, one has Z
Z Z T q vdv ¼ r nda þ qbdv Z v Za Zv T q0 v dV ¼ r nda þ q0 bdV V a V |fflfflfflfflfflfflfflfflfflffl ffl{zfflfflfflfflfflfflfflfflfflfflffl}
0
5.5 Equilibrium Equation
155
is Z
Z
rT ðndaÞ þ
r_ T nda þ
0¼
Z
q0 b dV a V Za Z Z
¼ r_ T nda þ rT ðtrlÞI lT nda þ q b dv a v Z Za T T T T q b dv r_ þ r trd r l $x dv þ ¼ v
v
which results in Eq. (5.43), noting Eqs. (1.392), (2.38) and r ¼ rT which will be verified in the next section. The equilibrium equation in the rate form, e.g. Eq. (5.43) is not needed necessarily by checking the equilibrium for the current variables pushing-forwarded to the ones in the reference configuration calculated by the constitutive equation, noting that the boundary condition is given by the stress tensor in the current configuration.
5.6
Equilibrium Equation of Angular Moment
Substituting Eq. (5.2) into Eq. (5.31) of the conservation law of angular momentum and noting Eq. (5.45), one has Z
Z
Z
qeijk xj vk dv ¼ v
eijk xj rkr nr da þ a
qeijk xj bk dv
ð5:44Þ
v
Because the first term in the right-hand side of this equation is rewritten as Z
Z eijk xj rkr nr da ¼ a
a
@xj rkr eijk dv ¼ @xr
Z @rkr dv eijk rkj þ eijk xj @xr v
Equation (5.44) leads to Z eijk rkj þ eijk xj v
@rkr dv ¼ 0 þ qbk qvk @xr
Noting the equilibrium Eq. (5.33) to this equation, it holds that eijk rkj ¼ 0 (e.g. e123 r23 þ e132 r32 ¼ r23 r32 ¼ 0) from which we have the symmetry of Cauchy stress tensor, i.e. r ¼ rT ;
rij ¼ rji
ð5:45Þ
156
5.7
5
Stress Tensors and Conservation Laws
Virtual Work Principle
Giving the scalar product of the arbitrary velocity increment dv to the equilibrium Eq. (5.33), we find
dv rx r þ qdv b ¼ qdv v
ð5:46Þ
Noting
divðdvrÞ ¼ r : graddv þ dvdivr r : graddv ¼ r : dl ¼ r : dd
ð5:47Þ
Equation (5.46) leads to divðdvrÞ r : dd þ qdv b ¼ qd
1 vv 2
i.e. r : dd ¼ divðdvrÞ þ þ qdv b qd
1 vv 2
ð5:48Þ
which is the local form of the virtual work equation in the current configuration. Now, consider the integration of Eq. (5.48) over the current volume. The integration of the first term in the right hand side in Eq. (5.48) yields Z
Z
Z
div ðdvrÞdv ¼
Z
ðdvrÞ nda ¼
v
a
dv ðrnÞda ¼ a
dv tda
ð5:49Þ
a
by virtue of Eqs. (1.391), (5.2) and (5.45). Further, noting Eq. (5.49), the integration of Eq. (5.48) leads to Z
Z r : dddv ¼ v
Z t dvda þ
a
Z qb dvdv
v
qd v
1 vv 2
dv
ð5:50Þ
Incidentally, applying the velocity v instead of dv, one has the following equation analogously to Eq. (5.50). Z
Z r : ddv ¼ v
t vda þ a
Z
Z qb vdv v
v
1 qv vdv 2
ð5:51Þ
The quantity in the left-hand side means the stress power, while the first, the second and the third terms in the right-hand side mean the powers done by the surface force (traction), the body force and the rate of the kinetic energy. Therefore, r : d designates the stress power done in the unit current volume.
5.8 Conservation Law of Energy
5.8
157
Conservation Law of Energy
The internal energy U is defined by Z U
ð5:52Þ
qedv v
where e is the internal energy per unit mass. The heat input caused by the internal heat source r per unit current unit mass and the heat flux vector q per unit area is given by Z Z Z Z qrdv qnda ¼ qrdv divqdv ð5:53Þ Q v
a
v
v
The stress power is given by Z Pr ¼
r : ddv
ð5:54Þ
v
The following energy equilibrium equation must hold.
U ¼ Pr þ Q
ð5:55Þ
i.e. Z
qedv
Z
Z
¼
r : ddv þ
v
v
Z qrdv
v
divqdv
ð5:56Þ
v
from which the conservation law of energy in the local form is given as
qe ¼ r : d þ qr divq
5.9
ð5:57Þ
Work Conjugacy
The work rate done for the unit volume in the current configuration is given by r : d ¼ rij dij which is the basic variable governing the conservation law of energy in Eq. (5.57). Designating the infinitesimal volumes in a specific region of material in the reference (initial) and the current configurations as dV and dvð¼JdVÞ, respec tively, the work rate w0 done per the unit reference volume is given from Eqs. (4.27), (4.75), (4.80), (4.126), (5.7), (5.13) and (5.17) as follows:
158
5
Stress Tensors and Conservation Laws
w0 ¼ r : ddv=dV ¼ trðrdÞJ ¼ trðslÞ ¼ trðsdÞ T 8 h i T T > > ðlFÞ ¼ tr PF tr sF > > > > > > < 1 T T
tr F sF dF ¼ trðSE Þ ¼ trðSC =2Þ F ¼ h i > T > > ¼ tr MLT tr FT sFT F1 lF > > > > > : trðBBÞ leading to
w0 ¼ Jr : d ¼ s : d ¼ P : F ¼ S : E ¼ S : C =2 ¼ M : L ¼ B : B
ð5:58Þ
where L ¼ F1 lF ¼ ~lG
ð5:59Þ
B ðSU þ USÞ
ð5:60Þ
G
which is called the Biot stress tensor. Equation (5.58)8 is derived also as follows: 1 1 trðS EÞ ¼ tr S ðU U þ U UÞ ¼ trðSU U þ US UÞ 2 2 i 1 h ¼ tr ðSU þ USÞ U ¼ trðB B Þ 2
ð5:61Þ
noting Eqs. (4.27)1 , (4.43) and (5.60). The work-conjugate pairs are described noting Eq. (3.24) as 8 G ^g j i T > P ¼ sF ¼ s ¼ s g G ; F ¼ IF ¼ l ^g ¼ lij gi G j > j i G < S ¼ F1 sFT ¼ s GG ¼ sij Gi Gj ; E ¼ FT dF ¼ d GG ¼ dij Gi G j > > G : G M ¼ FT sFT ¼ s G ¼ si j Gi Gj ; L ¼ F1 IF ¼ l G ¼ lij Gi G j leading to 8 j i j r i s j i r s > > > P : F ¼si g Gj :ls gr G ¼ si ls dr dj ¼ si lj <
S : E ¼sij Gi Gj :drs Gr Gs ¼ sij drs dri dsj ¼ sij dij ¼ sij lij > > > : M : L¼si j Gi Gj :lrs Gr Gs ¼si j lrs dir dsj ¼si j lij
5.9 Work Conjugacy
159
Therefore, the work-conjugate pair tensors possess the opposite (covariant and contravariant) variants to each other so that their scalar product is represented only by their components, while, needless to say, all tensors can be represented in any variants. The work rate reflecting the constitutive property is not concerned with a current unit volume but is concerned with a reference unit volume, noting that the mass in the current unit volume is variable but the mass in the reference unit volume is invariable. All the solids are composed of the solid particles, e.g. the crystals in metals and the soil particles in soils and thus they exhibit the hyperelastic deformation behavior in a low stress level and the inelastic deformations due to the mutual slips or the change of mutual positions between solid particles is induced in addition to the elastic deformation when a stress increases. Then, the stress tensor and the strain (rate) tensor used in a constitutive equation for solids must be the pair causing the work rate done to the unit reference volume. This requirement was proposed by Hill (1968, 1978) and Macvean (1968) through Truesdell (1952), Ziegler and Macvean (1967). This concept is called the work-conjugacy and the pair of the stress and the strain (rate) fulfilling the work-conjugacy is called the work-conjugate pair. The elastic constitutive equation describing the relation between the stress and the deformation is formulated by the second Piola-Kirchhoff stress tensor S and the right Cauchy-Green deformation tensor C. On the other hand, the plastic constitutive equation is formulated by the Mandel stress tensor M and the plastic part of the velocity gradient tensor L as will be shown Chap. 17. Here, note that the rate of C cannot be decomposed into the elastic and the plastic parts, while L can be decomposed into these parts exactly within the framework of the multiplicative decomposition of the deformation gradient tenor. Now, consider the unit reference orthogonal cell with the side vectors ðN1 ; N2 ; N3 Þ ðjjNi jj ¼ 1Þ which changes to the current cell with the side vectors ðn1 ; n2 ; n3 Þ by the relation ni ¼ FNi due to dx ¼ FdX as shown in Fig. 5.3. By taking account of the force f i ¼ PNi due to Eqs. (5.8) and (5.9) with dA ¼ 1 and (5.58)4 into the relation, it follows that
f 2 = PN 2 N2 f1 = PN1
N2 N3
N1 f 3 = PN3 (|| Ni || = 1)
f2
F
n2 n2
f1
N1
n1
n3 f3
n1
(||ni || 1)
Fig. 5.3 Current cell deformed from reference orthogonal unit cell to which First Piola-Kirchhoff stress applies
160
5
P : F ¼ trðPFT Þ ¼ ðF T PÞii ¼
Stress Tensors and Conservation Laws 3 X
Ni F T PNi
i¼1
¼
3 X
PNi ðFNi Þ ¼
i¼1
3 X
f i ni ¼ w0
ð5:62Þ
i¼1
Therefore, we can confirm the fact that P : F designates the work rate (power) done to the current cell which was the orthogonal unit cell in the reference state. The
equation P : C ¼ w0 can be derived by a similar way. The pairs of stresses and strain rates (or rates of deformation gradient) shown in Eq. (5.58) are called the work-conjugate pair. Stress and strain rate tensors in the work-conjugate pair have to be used for the formulation of constitutive equation. The stress vs. strain relations based on the typical work-conjugate pairs in the uniaxial loading process will be shown below.
(1) P F pair The relation df ¼ PdA in Eq. (5.10) is reduced to f ¼ PA, where the uniaxial components of f, P and A are designated by f , P and A, respectively. The time-integration of the work-conjugate deformation measure to the first Piola-Kirchhoff stress P is the deformation gradient Fð¼ @x=@XÞ which is
reduced to F ¼ l=L ¼ 1 þ u=L ¼ 1 þ e leading to F ¼ e. Consequently, designating the Young’s modulus E, the following relation holds. P ¼ Ee
ð5:aÞ
(2) S E pair The relation df ¼ FSdA in Eq. (5.16) is reduced to f ¼ FSA¼ ð1 þ eÞSA leading to S ¼ ðf =AÞ=ð1 þ eÞ ¼ P=ð1 þ eÞ ¼ Ee=ð1 þ eÞ, where the uniaxial component of S is designated by S. The work-conjugate deformation measure: Green strain E to the second Piola-Kirchhoff stress S is represented as pffiffiffiffiffiffiffiffiffiffiffiffiffiffi E ¼ u=L þ ðu=LÞ2 =2 ¼ e þ e2 =2 leading to e ¼ 1 þ 1 þ 2E. Consequently, noting Eq. (5.a), one has the relation ! 1 S ¼ E 1 pffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 þ 2E
ð5:bÞ
(3) rd pair The relation df ¼ sda in Eq. (5.3) leads to df ¼ rda under the volume-preserving deformation J ¼ 1, which is reduced to r ¼ f =a ¼ f =ðAL=lÞ¼ ðf =AÞðl=LÞ¼ ðf =AÞ½ðL þ uÞ=L ¼ Pð1 þ eÞ ¼ Eeð1 þ eÞ in the uniaxial loading process. The work-conjugate deformation rate measure to the Cauchy stress tensor r is the strain rate tensor d as shown in Eq. (5.58). The
5.9 Work Conjugacy
161
time-integration of the axial component l =l of d is identical to the axial component e of the Eulerian logarithmic strain eð0Þ in Eq. (4.50), i.e. e ¼ lnðl=LÞ¼ lnð1 þ u=LÞ ¼ lnð1 þ Þ leading to e ¼ exp e 1 in the uniaxial loading process. Consequently, noting Eq. (5.a), one has the relation r ¼ Eðexp e 1Þ exp e
ð5:cÞ
The stress versus strain relations in Eqs. (5.a), (5.b) and (5.c) are shown in Table 5.2 and depicted choosing the Young’s modulus E ¼ 100; 000N=mm2 in Fig. 5.4 referring to Ishikawa (2015). It is known from this figure that the stress versus strain relations differ significantly from each other for the strain greater than 0.1. Table 5.2 Various stress vs. strain relations by various work-conjugate pairs in uniaxial loading Classification of deformation
Strains and stresses
Stress versus strain relations
Infinitesimal deformation
Nominal strain 1st Piola-Kirchhoff (nominal) stress Green strain
e ¼ u=L F P¼ A
Finite deformations
Lagrangian (reference) description
Eulerian (current) description
2nd Piola-Kirchhoff stress Logarithmic (true) strain Cauchy (true) stress
E¼ S¼
u þ L ð1 þ
1 u 2 1 ð Þ ð¼ e þ e2 Þ 2 L 2 u 1 Þ F P L ð¼ Þ A 1þe
l u e ¼ ln ¼ lnð1 þ Þ ð¼ lnð1 þ eÞÞ L L F F u r ¼ ¼ ð1 þ Þð¼ Pð1 þ eÞÞ a A L
E(exp e 1)exp e
P
S
E E (1
1 ) 1 2E
Fig. 5.4 Various work-conjugate stress versus strain relations premised on linear relation of P ¼ Ee in uniaxial loading ðE ¼ 100; 000 N/mm2 Þ
162
5.10
5
Stress Tensors and Conservation Laws
Various Simple Deformations
Let various strain (rate) and stress (rate) described in the foregoing be shown explicitly and let their relation be described for various simple deformations. These deformations are often observed in experiments for measurement of material properties. Homogeneous and isotropic deformation is assumed therein.
5.10.1
Uniaxial Loading
For a cylindrical specimen with the initial length L and the initial radius R, suppose that the length and the radius changes to l and r (Khan and Huang 1995). Choosing the X1 -axis to the axial direction of cylinder, it holds that x 1 ¼ k1 X 1 ; x 2 ¼ k2 X 2 ; x 3 ¼ k3 X 3
ð5:63Þ
where k1 ¼
l r ; k2 ¼ k 3 ¼ L R
ð5:64Þ
from which one has 2
k1 6 F ¼U ¼ V ¼ 4 0 0
0 k2 0
3 0 7 0 5; k2
2
F1
k1 1 6 ¼4 0 0
0 k1 2 0
R ¼I
3 0 7 0 5; k1 2
9 > > > 2> J ¼ detF ¼ k1 k2 = > > > > ;
ð5:65Þ The aforementioned measures of deformation (rate) are given as 2
C ¼ U2 ;
k21 T 4 F F¼ 0 0
0 k22 0
3 2 2 0 k1 0 5; b1 ¼ V2 ¼ FT F1 ¼ 4 0 k22 0
0 k2 2 0
3 0 0 5 k2 2 ð5:66Þ
2
2 1 1 4 k1 1 E ¼ ðC IÞ ¼ 0 2 2 0
0 k22 1 0
3 0 0 5 k22 1
5.10
Various Simple Deformations
163
2 2 1 1 4 1 k1 1 ¼ e¼ Ib 0 2 2 0 2
Eð0Þ ¼ eð0Þ
l ¼ F F1
ln k1 ¼4 0 0
0 2 ll1 6 ¼4 0 0
3 2 lnðl=LÞ 0 0 5¼4 0 0 ln k2
0 ln k2 0
2 k1 0 6 6 ¼¼ 4 0 k2
0 1 k2 2 0
3 0 0 5 1 k2 2
0 lnðr=RÞ 0
ð5:67Þ
3 0 0 5 ð5:68Þ lnðr=RÞ
32 3 3 2 1 k k 0 0 1 k1 0 0 1 76 1 7 7 6 1 0 5¼6 k2 k1 0 7 07 4 0 54 0 k2 5 2 1 0 0 k 1 2 0 k2 0 0 k2 k2 3 0 0 ð0Þ ð0Þ 7 1 rr 0 5¼d¼E ¼e 0
rr 1
0
ð5:69Þ w¼O
ð5:70Þ
The infinitesimal strain in Eq. (4.31) is given by 2
ðl LÞ=L e¼4 0 0
0 ðr RÞ=R 0
3 2 0 0 k1 1 5¼4 0 k2 1 0 ðr RÞ=R 0 0
3 0 0 5¼UI k2 1
ð5:71Þ 2 l =L 6 e¼4 0 0
3 0 r =R 0
0 7 0 5 r =R
ð5:72Þ
Denoting the axial load as F, various stresses are shown as follows: 2
F 6 pr2 r¼4 0 0
3
2
F 0 0 2 7 6 6 k2 pR2 0 05 ¼ 4 0 0 0 0
3
0 0 0
2
F 07 6 2 7 ¼ 6 k2 A0 05 4 0 0 0
3 0 0 0
07 7 05 0
ð5:73Þ
164
5
2
3
F 2 6 k2 A 0 s ¼ Jr ¼ k1 k22 6 4 0 0 2
k1 1
P ¼ JF1 r ¼ k1 k22 4 0 0
0 k1 2 0
0 0 0
2F 07 k1 A0 7¼6 4 0 05 0 0
2 3 F 0 6 2 k2 A0 0 56 4 0 1 k2 0
S ¼ JF1 rFT
5.10.2
Stress Tensors and Conservation Laws
2 F 6 ¼ 4 k10A0 0
0 0 0
3 0 0 0
0
7 05 0
3 2 F 07 6 A 7¼4 0 0 05 0 0
ð5:74Þ 3 0 0
7 0 0 5 ð5:75Þ 0 0
3 0 0
7 0 05 0 0
ð5:76Þ
Simple Shear
Consider the simple shear in which the shear deformation is induced in parallel to the x1 -axis as shown in Fig. 5.5. x1 ¼ X1 þ cX2 ; x2 ¼ X2 ; x3 ¼ X3
ð5:77Þ
where c is the engineering shear strain. Denoting the shear angle by h, it holds that c ¼ 2tanh;
2 cos h ¼ pffiffiffiffiffiffiffiffiffiffiffiffi ; 4 þ c2
/2
/2
1
X 2 = x2 e2
X
X2 x
e1 0 Fig. 5.5 Simple shear
X1
x1
c sin h ¼ pffiffiffiffiffiffiffiffiffiffiffiffi 4 þ c2
ð5:78Þ
5.10
Various Simple Deformations
165
2c ¼ c cos2 h; h h¼ c ¼ 2 h sec2 2 4 þ c2
ð5:79Þ
It holds in this situation that 2
1 F ¼ 40 0
2
3 0 0 5; 1
c 1 0
F1
3 0 0 5; 1
c 1 0
1 ¼ 40 0
J ¼ detF ¼ 1
ð5:80Þ
where the inverse tensor F1 is derived using Eq. (1.161). The components in the third line and those in the third row are zero except for unity in the third line and the third row in all tensors appearing hereinafter for the simple shear deformation. Then, for simplicity, let them be expressed by the matrix with two lines and two rows.
c ; 1
1 F¼ 0
F
1
1 ¼ 0
c 1
ð5:81Þ
from which it is obtained that
l ¼ FF
1
d¼
¼ 0 c 0 0 0 1
1 c ; 0 2
0 0
c ¼ 0 1 0
w¼
0 1
c 0
ð5:82Þ
1 c 0 2
ð5:83Þ
Further, it follows noting Eq. (1.161) that
1 c 1 þ c2 1 2 T 2 1 T C ¼U ¼F F ¼ C ¼U ¼F F ¼ 2 ; c 1þ c c
2 1 1 2 c 2 þ c c U1 ¼ pffiffiffiffiffiffiffiffiffiffiffiffi U ¼ pffiffiffiffiffiffiffiffiffiffiffiffi 2 ; c 2 4 þ c2 c 2 þ c 4 þ c2
1 1 þ c2 c b ¼ V2 ¼ FFT ¼ ; b1 ¼ V2 ¼ FT F1 ¼ c c 1
1 1 2 c 2 þ c2 c ; V1 ¼ pffiffiffiffiffiffiffiffiffiffiffiffi V ¼ pffiffiffiffiffiffiffiffiffiffiffiffi 2 c 2 4 þ c2 4 þ c2 c c þ 2
9 c > > 1 = > > ;
ð5:84Þ 9 c > > 1 þ c2 = > > ; ð5:85Þ
b1 ¼ V2 ¼ FT F1 ¼
1 c
c 1 þ c2
ð5:86Þ
166
5
1 2 R ¼ FU ¼V F ¼ pffiffiffiffiffiffiffiffiffiffiffiffi 2 c 4þc 1
c cos h ¼ 2 sin h
1
sin h R¼h cos h
Stress Tensors and Conservation Laws
sin h cos h
2c c cos h ¼ 3=2 2 2 sin h ð4 þ c Þ
2 c
ð5:87Þ
ð5:88Þ
2c 1 0 1 ¼ 0 4 þ c2 1 0
1 0 1 1 0 c 1 c 1 E ¼ ðC GÞ ¼ ;e ¼ g b ¼ 2 2 c c2 2 2 c c2
1 1 0 c T e ¼ FþF I ¼ 2 2 c 0
0 X ¼ RR ¼ h 1 R
ð5:89Þ
T
ð5:90Þ ð5:91Þ
noting FU1 ¼
R¼h
1 1 c 2 þ c2 pffiffiffiffiffiffiffiffiffiffiffiffi 0 1 c 4 þ c2
2
c 2
c ffi pffiffiffiffiffiffiffiffi 2
2 ffi pffiffiffiffiffiffiffiffi
2c 4 4þc sin h cos h 5 ¼ c ffi 2 ffi pffiffiffiffiffiffiffiffi cos h sin h 4 þ c2 pffiffiffiffiffiffiffiffi 2 2 4þc
sin h cos h RR ¼ h cos h sin h
3
4 þ c2
T
cos h sin h
4þc
sin h cos h
Various stress tensors are described using the notation s ¼ r12 as follows:
r¼P¼ P ¼ JrFT ¼
r11 s
s r22
1 c
r11 s
s r22
0 r11 cs ¼ s cr22 1
ð5:92Þ s r22
1 c r11 cs s S ¼ JF1 rFT ¼ F1 P ¼ 0 1 s cr22 r22
r11 c2 r22 2cs s cr22 ¼ s cr22 r22
ð5:93Þ
ð5:94Þ
5.10
Various Simple Deformations
5.10.3
167
Combination of Tension and Distortion
Consider a thin cylindrical specimen subjected to the combination of tension and distortion described by the following equation in the polar coordinate system r ¼ aR; h ¼ H þ xZ;
z ¼ kZ
ð5:95Þ
where ðR; H; ZÞ signifies the initial configuration, and a, x and k denote the proportionality factors depending on the deformation, while x is described by the relative distortion angle / between both ends as follows: x /=L
ð5:96Þ
L being the length of the specimen. The explanation in the following is referred to Khan and Huang (1995). Variables describing the deformation are given as follows: 2
@r 6 @R FrZ 6 7 6 6 r@h 7 FhZ 5 ¼ 6 6 @R 6 FzZ 4 @z @R
@r R@H FrR FrH 6 r@h F¼6 4 FhR FhH R@H FzR FzH @z R@H 2 3 @aR @aR @aR 6 7 2 @R R@H @Z a 6 7 6 7 6 r@ðH þ xZÞ r@ðH þ xZÞ r@ðH þ xZÞ 7 6 6 ¼6 7 ¼ 40 6 7 @R R@H @Z 6 7 0 4 5 @kZ @kZ @kZ @R R@H @Z 2
3
3 @r @Z 7 7 7 r@h 7 7 @Z 7 7 @z 5 @Z
0 a 0
0
3
7 xaR 7 5
ð5:97Þ
k
from which we have 21 F1
6a 6 6 ¼6 60 6 4 0
3 0
0
7 7 1 xR 7 7 a k 7 7 5 1 0 k
ð5:98Þ
168
5
Stress Tensors and Conservation Laws
and 2
a V ¼ 40 0
3 0 ksin/ 5 k cos /
ð5:99Þ
3 0 5 a sin / axR sin / þ k cos /
ð5:100Þ
0 a cos / þ axR sin / k sin /
2
a U ¼ 40 0
0 a cos / a sin /
where kþa cos / ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ; ða þ kÞ2 þ ðaxRÞ2
axR sin / ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ða þ kÞ2 þ ðaxRÞ2
ð5:101Þ
R being the mean radius ðR ffi RÞ of the thin cylindrical specimen in the initial state. Further, one has 2
a2 C ¼ FT F ¼ 4 0 0
3 0 5 xa2 R 2 k2 þ x 2 a 2 R
0 a2 xa2 R
2
b1 ¼ FT F1
1 6 a2 6 6 ¼6 60 6 4 0
l ¼ FF1
ð5:102Þ 3
0 1 a2 xR ak
0
7 7 xR 7 7 ak 7 7 2 x2 R 15 þ 2 k2 k
2 a 3 xZ 0 6 a 7 6 7 6 7 a xaR 6 7 ¼ 6xZ a k 7 6 7 4 5 k 0 0 k
ð5:103Þ
ð5:104Þ
from which we have 2 aa 6 6 0 d¼6 6 4 0
0 a a x aR 2k
3 0 x aR 7 7 7; 2k 7 5 k k
2
0
6 6xZ w¼6 6 4 0
xZ 0
x aR 2k
3 0 x aR 7 7 2k 7 7 5 0
ð5:105Þ
5.10
Various Simple Deformations
169
Stresses in various definitions are described by the following equations, designating the normal stress rzz and the shear stress rrh applied to the traverse section of the cylinder by r and s, respectively. 2
3 0 0 0 s5 s r
0 r ¼ 40 0
ð5:106Þ
It holds from Eqs. (5.97), (5.98) and (5.106) that 2
3
1 6a 6 6 P ¼ JF1 r ¼ a2 k6 0 6 4 0 2
0 1 a 0
0 0 ¼ 4 0 a2 xsR 0 a2 s 2
S ¼ PFT
0 ¼ 40 0 2
0 6 6 ¼ 60 4 0
0 a2 xsR a2 s
0
72 7 0 xR 74 7 0 k 7 1 5 0 k 0
0 0 s
3 0 s5 r
3 0 aks a2 xrR 5 a2 r
2 1 36 0 6a 6 2 aks a xrR 56 0 6 2 ar 4 0
2
a2 x2 rR 2axsR þ k a2 xrR as k
ð5:107Þ 3 0 1 a xR k 3
0 2 7 7 as a xrR k 7 5 2 a k
07 7 7 07 7 15 k
ð5:108Þ
Chapter 6
Objectivity and Objective (Rate) Tensors
Constitutive property of material is independent of observers. Therefore, constitutive equation has to be described by variables obeying the common objective transformation rule described in Sect. 1.3.1. State variables, e.g. stress, strain and back stress tensors in the same configuration obey the common coordinate transformation rule. However, the material-time derivatives of tensors in the current configuration do not obey the objective transformation rule, since they are influenced by the rigid-body rotation. Then, instead of the material-time derivative of tensors, particular time-derivatives of tensors obeying the objective transformation rule, i.e. the convective time-derivative have to be adopted in constitutive equations. The consideration on the fulfillment of objectivity is of great importance for the hypoelastic-based constitutive equation formulated in the current configuration which is influenced directly by the rigid-body rotation. Then, the objectivity and the formulation of constitutive relations fulfilling the objectivity will be explained in this chapter.
6.1
Objectivity
Physical quantities except for scalar ones are observed to be different depending on the state, e.g. position, direction, velocity of observers. On the other hand, mechanical property of material is observed identically independent of the state of observers. In particular, it is observed identically independent of the rigid-body rotation of material. Therefore, a constitutive equation describing material property must be expressed in a common form independent of coordinate systems. Then, it must be described so as not to be influenced by the rigid-body rotation of material. This fact was not so obvious in the olden time and was advocated by Oldroyd (1950) in the middle of the last century. It is referred to as the principle of material-frame indifference (Oldroyd 1950) or principle of objectivity or simply © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_6
171
172
6
Objectivity and Objective (Rate) Tensors
objectivity. This would be regarded as one of the starting points of the modern continuum mechanics which is called sometimes as the rational mechanics (Truesdell and Toupin 1960; Truesdell and Noll 1965) in order to describe the finite elastoplastic deformation and has been studied over the last half century prior to the recent formulation by the hyperelastic-based plasticity which will be comprehensively described in Chap. 17. Here, note that components of tensor describing mechanical state of material, e.g. stress, strain and anisotropic internal variables are observed to be changed by the fixed coordinate system if the material rotates, even when the components are observed to be unchanged by the coordinate system rotating concurrently with the material itself. Therefore, the material-time derivative of tensor describing mechanical state is observed to be non-zero, by the fixed coordinate system when the material rotates even when it is observed to be zero by the observer rotating concurrently with the material itself. It is caused by the fact that the material-time derivative of tensor designates the rate of tensor observed by the coordinate system moving in parallel with material but without rotation. Then, the material-time derivative cannot be adopted for the description of constitutive equations in a current rate form. Machine elements are often subjected not only to deformation but also to rigid-body rotation, as seen in metal forming, gears, wheels, etc. Soils near the side edges of footings, at the bottom ends of piles, etc. undergo a large rigid-body rotation. Therefore, formulations of constitutive equations which is not influenced by the rigid-body rotation are of great importance in practical engineering problems.
6.2
Influence of Rigid-Body Rotation on Various Mechanical Quantities
In order to check whether or not a constitutive equation is formulated so as to satisfy the objectivity principle, it is expedient to examine the influence of rigid-body rotation on the tensor variables used in constitutive equations. Instead, one may examine how the components of these variables areobserved by the coordinate systems with the fixed base fei g and the rotating base ei ðtÞ which are related as ei ðtÞ ¼ QT ðtÞei ;
ei ð0Þ ¼ ei ;
ei ðtÞ ¼ QT ðtÞei
ð6:1Þ
provided that the rotating base ei ðtÞ coincides with the fixed base fei g at the beginning of deforming/rotation ðt ¼ 0Þ, noting Eq. (1.111), where one has QðtÞ ¼ ei ei ðtÞ;
Qð0Þ ¼ I
These bases are illustrated in Fig. 6.1 for the two-dimensional state.
ð6:2Þ
6.2 Influence of Rigid-Body Rotation on Various Mechanical Quantities
173
Fig. 6.1 Coordinate systems with the fixed base fei g and the rotating base ei ðtÞ which coincides with the base fei g at the beginning of deformation/rotation (illustrated in two-dimensional state for e3 ¼ e3 )
The components of the initial infinitesimal line element dX in the initial state (t ¼ 0) is observed to be identical by these bases, since ei ðtÞ coincides with fei g in the initial state. On the other hand, the components of the current infinitesimal line-element dxðtÞ is observed to be different by these bases as the rotating base ei ðtÞ differs from the fixed base fei g for t [ 0. Here, noting dxðtÞ ¼ FðtÞdX, we have dX ¼ dX
ei ð0Þ ¼ fei g ;
dx ðtÞ ¼ QðtÞdxðtÞ ¼ QðtÞFðtÞdX
ð6:3Þ
and dx ðtÞ ¼ F ðtÞdX ð0Þ ¼ F ðtÞdX
ð6:4Þ
from which it follows that F ðtÞ ¼ QðtÞFðtÞ
ð6:5Þ
It is known from Eq. (6.5) that the deformation gradient FðtÞ is the second-order tensor but it obeys the transformation rule of the first-order tensor. This is based on the fact that the deformation gradient is the two-point tensor as specified in Eq. (3.9). Substituting Eq. (6.5) into Eqs. (4.2)–(4.5), (4.16), (4.17) and (4.27), the following relations are obtained for various quantities describing a deformation.
174
6
U ¼ U;
ðF TF ¼ ðQFÞT QF ¼ FT QT QF ¼ FT FÞ
C ¼ C
V ¼ QVQT ;
Objectivity and Objective (Rate) Tensors
b ¼ QbQT
ð6:6Þ
ðF FT ¼ QFðQFÞT ¼ QFFT QT Þ
ð6:7Þ
ðF ¼ QF ¼ QRU ¼ QRU ¼ R U Þ
R ¼ QR
ð6:8Þ
E ¼ E ð ¼ FT eFÞ; e ¼ QeQT
ð6:9Þ
Noting the relation
F F1 ¼ ðQFÞ ðQFÞ1 ¼ ðQF þ QFÞF1 Q1 ¼ QðFF1 QT QÞQT it holds for the velocity gradient in Eq. (4.75) that l ¼ Qðl XÞQT ¼ QlQT þ X
ð6:10Þ
where X and X are defined by
X QT Q;
X Qri Qrj ei ej
X QQT ;
X Qir Qjr ei ej
)
ð6:11Þ
and they are related by X ¼ QXQT ;
X ¼ QT XQ
ð6:12Þ
where Q is given by
Q ¼ er er
ð6:13Þ
from Eq. (6.2) because of ei ¼ 0. Substituting Eqs. (6.2) and (6.13) into Eq. (6.11), we have
X ¼ er er ;
X ¼ ðei ej Þei ej
ð6:14Þ
from which it follows that
ei ¼ Xei
ð6:15Þ
It is known from Eq. (6.15) that X is the spin of the base ei . The substitution of Eq. (6.10) into Eqs. (4.80) and (4.81) yields the following transformation rules.
6.2 Influence of Rigid-Body Rotation on Various Mechanical Quantities
175
d ¼ QdQT
ð6:16Þ
w ¼ QðwXÞQT ¼ QwQT þ X
ð6:17Þ
The relative spin in Eq. (4.85), i.e.
XR R RT
ð6:18Þ
obeys the transformation identical to that of w as follows:
XR ¼ QðXR XÞQT ¼ QXR QT þ X
ð6:19Þ
noting
R RT ¼ ðQRÞ ðQRÞT ¼ ðQR þ QRÞRT QT
¼ QRRT QT þ QQT ¼ QXR QT þ QQT QQT ¼ QXR QT QQT QQT We obtain the following conclusions for the influence of rigid-body rotation from Eqs. (6.6)–(6.19). (1) The right Cauchy-Green deformation tensor C and the Green strain tensor E are based in the reference configuration and thus they are observed to be unchangeable, i.e. invariant, obeying the transformation rule of scalar quantities independent of the rigid-body rotation. On the other hand, the left Cauchy-Green deformation tensor b and the Almansi strain tensor e are based in the current configuration and thus obey the transformation rule of second-order tensor. (2) The strain rate tensor d obeys the transformation rule of the second-order tensor. On the other hand, the velocity gradient tensor l, the continuum spin tensor w and the relative spin XR are directly subjected to the influence of rate of rigid-body rotation, lacking the objectivity. The following transformations hold for stress tensors described in Chap. 5. r ¼ QrQT
ðtt ¼ Qtt ¼ Qrn ¼ QrQT Qn ¼ r n Þ s ¼ QsQT
T
ðP ¼ Jr F
ð6:20Þ ð6:21Þ
P ¼ QP ¼ JQrQ ðQFÞ ¼ JQrQT QT FT ¼ QJrFT ¼ QPÞ T
T
ð6:22Þ
176
6
Objectivity and Objective (Rate) Tensors
S ¼ S ¼ ðQFÞ1 ðQsQT ÞðQFÞT S ¼F s F ¼ F1 QT QsQT QT FT ¼ F1 sFT ¼ S
1 T
ð6:23Þ
Then, the Cauchy stress tensor r and the Kirchhoff stress tensor s obeys the transformation rule of the second-order tensor. On the other hand, the first Piola-Kirchhoff stress tensor P obeys the transformation rule of the first-order tensor so that it is the two-point tensor. The second Piola-Kirchhoff stress tensor S is the invariant under the superposition of rigid-body rotation. The consideration of objectivity is of great importance in the formulation of hypoelastic-based plastic constitutive equations since the rates of stress and anisotropic internal state variables in the current configuration are influenced directly by the rigid-body rotation as described above. Then, the time-derivatives of state variables will be further considered in the subsequent sections.
6.3
Material-Time Derivative of Tensor
The material-time derivatives of state variables is the rates of them observed by the coordinate system moving in parallel with material particle as explained in Sect. 2.2. However, it will be mathematically verified in this section that the material-time derivative of tensor does not obey the objective transformation rule and thus it cannot be used in constitutive equations. Consider the transformation of the material-time derivative of a state variable obeying the objective transformation in Eqs. (1.90) and (1.92). The material-time derivative of the tensor t in the current configuration reads (Hashiguchi 2007a):
t p1 p2
pm
¼ Qp1 q1 Qp2 q2
Qpm qm tq1 q2
qm
þ Qp1 q1 Qp2 q2 Qpm qm tq1 q2
þ Qp1 q1 Qp2 q2 Qpm qm tq1 q2
qm
þ
qm
þ Qp1 q1 Qp2 q2 Qpm qm tq1 q2
qm
ð6:24Þ
tp1 p2
pm
¼ Qq1 p1 Qq2 p2 Qqm pm tq1 q2
qm
þ Qq1 p1 Qq2 p2 Qqm pm tq1 q2
qm
þ
þ Qq1 p1 Qq2 p2 Qqm pm tq1 q2
qm
þ Qq1 p1 Qq2 p2 Qqm pm tq1 q2
qm
ð6:25Þ
Noting the relation Qpi qi ¼ dpi s Qsqi ¼ Qpi t Qst Qsqi ¼ Qpi t Qst Qsqi ¼ Qpi t Xtqi and replacing t ! qi , qi ! ri , then Eqs. (6.28) and (6.29) can be rewritten as follows:
6.3 Material-Time Derivative of Tensor
tp1 p2
pm
¼ Qp1 q1 Qp2 q2 Qpm qm ðtq1 q2 Xq1 r1 tr1 q2
qm
qm
Xq2 r2 tq1 r2
qm
Xq1 rm tq1 q2
Xqm rm tq1 q2
rm Þ
ð6:26Þ
t p1 p2
177
pm
¼ Qq1 p1 Qq2 p2 Qqm pm ðtq1 q2
Xq1 r1 tr1 q2 qm
qm Xq2 r2 tq1 r2 qm
rm Þ
ð6:27Þ
It is known from Eqs. (6.26) and (6.27) that the material-time derivative does not
obey the objective transformation rule, noting that the components tp1 p2
pm
in the
t p1 p2 pm
in the fixed coordinate system is not zero even when the components coordinate system rotating with the material is zero. Equations (6.26) and (6.27) are expressed for the vector v and the second-order tensor t in symbolic notation as follows:
v ¼ Qðv XvÞ; v ¼ QT ðv Xv Þ
ð6:28Þ
t ¼ Qðt Xt þ tXÞQT ; t ¼ QT ðt Xt þ t XÞQ
ð6:29Þ
Consequently, the material-time derivative cannot be adopted in constitutive equations. In order to see the irrationality for using the material-time derivative of tensor, consider the hypoelastic constitutive equation in which the material-time derivative of Cauchy stress tensor is related linearly to the strain rate tensor as follows:
r¼H:d
ð6:30Þ
where the tangent modulus tensor H (fourth-order tensor) is the function of stress and anisotropic internal variables in general. It follows from this equation with Eq. (6.29) that
d ¼ H1 : QT ðr Xr þ r XÞQ
ð6¼ H1 : QT r Q ¼ H1 : QT ðQrQT Þ Q ¼ H1 : rÞ
ð6:31Þ
This leads to the irrational result that the deformation is induced, i.e. d 6¼ O even if the stress observed by the coordinate system rotating ðX ¼ QXQT 6¼ OÞ with the material itself does not change, i.e. r ¼ O. This is caused by the non-objectivity of material-time derivative of tensor, while the strain rate d is not a material-time derivative of tensor but is the original tensor defined so as to obey the objective transformation by excluding the continuum spin tensor w from the velocity gradient tensor l. In the next section, the objective time-derivative of tensor will be introduced which is based on the rate of tensor observed by the coordinate system deforming/rotating with material itself, satisfying the objectivity.
178
6
Objectivity and Objective (Rate) Tensors
The Eulerian tensor changes even when the state of physical quantity observed from the material itself does not change under a material rotation, since the base vectors change even by a rotation of the material. On the other hand, the Lagrangian tensor pulled-back to the reference configuration with the fixed base vectors does not change in the material rotation and thus it called the rotation-free tensor, while the Lagrangian tensor inherits the components in the Eulerian tensor. Then, the constitutive relation described by the Lagrangian tensors is used in the deformation analysis. The Eulerian base vectors are calculated by the push-forward operation of the Lagrangian (reference) base vectors through the deformation gradient tensor, which is required to capture the Eulerian tensor in the current configuration from the Lagrangian tensor and vice versa. The physical and geometrical interpretations for the relations between the above-mentioned Eulerian and Lagrangian tensors can be referred to Hashiguchi and Yamakawa (2012).
6.4
Objectivity of Convective Time-Derivative and Corotational Rate
The convected derivatives in Eq. (3.34) satisfy the objectivity obviously because they are based on the rates of tensor observed by a material itself. In addition, this fact can be mathematically confirmed as shown below for Eq. (3.34)1 as an example, noting Eq. (6.10). !
t
gg
¼ F ðF1 t FT Þ FT ¼ QF½ðQFÞ1 QtQT ðQFÞT ðQFÞT ¼ QF½F1 QT QtQT QFT FT QT ¼ Q½FðF1 tFT Þ FT QT
ð6:32Þ
or !
t
gg
¼ t l t t lT ¼ ðQtQT Þ Qðl XÞQT QtQT QtQT ½Qðl XÞQT T
¼ QtQT þ QtQT þ QtQT Qðl XÞtQT QtQT QðlT XT ÞQT
T
¼ QtQT þ Q tQT þ QtQT Qðl Q QÞtQT QtQT QðlT QT QÞQT
T
¼ QtQT þ QtQT þ QtQT QltQT QQT Q tQT QtlT QT QtQT QQT
¼ Qðt lt tlT ÞQT ð6:33Þ Hereinafter, we consider the general corotational rate tensor and show its objectivity in the following.
6.4 Objectivity of Convective Time-Derivative and Corotational Rate
179
The general corotational rate tensor is represented as t ¼ R ðRT ½½tÞ
ð6:34Þ
by the symbolic notation defined in Eq. (1.103). Equation (6.34) is expressed in the component description as follows (Hashiguchi 2003): t p1 p2 pn
¼ Rp1 q1 Rp2 q2 Rpn qn ðRs1 q1 Rs2 q2 Rsn qn ts1 s2
sn Þ
¼ Rp1 q1 Rp2 q2 Rpn qn ðRs1 q1 Rs2 q2 Rsn qn ts1 s2
þ Rs1 q1 Rs2 q2 Rsn qn ts1 s2 þ
sn
þ Rs1 q1 Rs2 q2 Rsn qn ts1 s2
sn
þ Rs1 q1 Rs2 q2 Rsn qn ts1 s2
sn
sn Þ
¼ Rp1 q1 Rs1 q1 Rp2 q2 Rs2 q2 Rpn qn Rsn qn ts1 s2
sn
þ Rp1 q1 Rs1 q1 Rp2 q2 Rs2 q2 Rpn qn Rsn qn ts1 s2
sn
sn
þ Rp1 q1 Rs1 q1 Rp2 q2 Rs2 q2 Rpn qn Rsn qn ts1 s2 þ
þ Rp1 q1 Rs1 q1 Rp2 q2 Rs2 q2 Rpn qn Rsn qn ts1 s2
¼ dp1 s1 dp2 s2 dpn sn ts1 s2
¼ t p1 p2
pn
Xp1 s1 ts1 p2
sn
Xp1 s1 dp2 s2 dpn sn ts1 s2
sn
þ dp1 s1 ðXp2 s2 Þ dpn sn ts1 s2
pn
sn
þ
Xp2 s2 tp1 s2
sn
þ dp1 s1 dp2 s2 ðXpn sn Þts1 s2 pn
X p n s n tp 1 p 2
sn
sn
ð6:35Þ
noting Eq. (6.18) with Rp1 q1 Rs1 q1 ¼ ðRRT Þp1 s1 ¼ ðRRT Þp1 s1 ¼ Xp1 s1 . The objectivity of the corotational rate is verified as follows: t p1 p2 pn
¼ Rp1 q1 Rp2 q2 Rpn qn ðRs1 q1 Rs2 q2 Rsn qn ts1 s2
sn Þ
¼ Qp1 t1 Rt1 q1 Qp2 t2 Rt2 q2 Qpn tn Rtn qn ½Qs1 r1 Rrn qn Qs2 r2 Rr2 q2 Qsn rn Rrn qn ðQs1 u1 Qs2 u2 Qsn un tu1 u2 ¼ Qp1 t1 Qp2 t2 Qpn tn Rt1 q1 Rt2 q2 Rtn qn ½Rrn qn Rr2 q2 Rrn qn ðQs1 r1 Qs1 u1 Qs2 r2 Qs2 u2 Qsn rn Qsn un tu1 u2
un Þ
un Þ
¼ Qp1 t1 Qp2 t2 Qpn tn Rt1 q1 Rt2 q2 Rtn qn ½Rrn qn Rr2 q2 Rrn qn ðdr1 u1 dr2 u2 drn un tu1 u2 un Þ ¼ Qp1 t1 Qp2 t2 Qpn tn Rt1 q1 Rt2 q2 Rtn qn ½Rr1 q1 Rr2 q2 Rrn qn tr1 r2
¼ Qp1 t1 Qp2 t2 Qpn tn t t1 t2
tn
rn
ð6:36Þ
180
6
Objectivity and Objective (Rate) Tensors
i.e.
t ¼ Q½½t
ð6:37Þ
The fourth-order tensor as the special case of Eq. (6.36) was used to the anisotropic tensor for the directional distortional yield surface by Feigenbaum and Dafalias (2014) referring to Hashiguchi (2003).
6.5
Various Objective Stress Rate Tensors
Various rates of the Cauchy stress r and the Kirchhoff stress s ð¼ JrÞ can be obtained from the aforementioned convected and corotational time-derivatives as will be shown in this section. Corotational time-derivative with a spin tensor is designated by the symbol ð Þ as shown in the last section. On the other and, the time-derivatives other than corotational time derivatives are designated by the symbol ðD Þ. (a) Contravariant convected rates Based on Eq. (3.34)1, the contravariant convected rate of the Cauchy stress r is given by ! gg
D
rOl r
D
¼ FðF1 rFT Þ FT ¼ FS=J FT ¼ r lr rlT ð¼ rOl Þ T
ð6:38Þ
which is termed the Oldroyd rate of Cauchy stress (Oldroyd 1950). Likewise, it holds for the Kirchhoff stress that DOl
s
! gg
s
D
¼ FðF1 sFT Þ FT ¼ FSFT ¼ s ls slT ð¼ sOl Þ T
ð6:39Þ
which is termed the Oldroyd rate of Kirchhoff stress. Further, D
rTr
J
1 DOl
s
¼J
1
!
gg
D
ðJrÞ ¼ J 1 FðF1 ðJrÞFT Þ FT ¼ J 1 FS FT ¼ rOl þ rtrd
¼ r lr rlT þ rtrdð¼ rTrT Þ ð6:40Þ is termed the Truesdell rate of Cauchy stress.
6.5 Various Objective Stress Rate Tensors
181
(b) Covariant-contravariant convected rates The covariant-contravariant convected rate of the Kirchhoff stress s is given from Eq. (3.34)3 as ð6:41Þ t t The particular case of the rate in Eq. (6.41) is given by D
!
D
T T T T T P s s ^g ¼ ðsF Þ F ¼ P F ¼ s sl ð6¼ P s Þ g
ð6:42Þ
which is termed the relative 1st Piola-Kirchhoff stress rate. The following stress rate is defined as the nominal stress rate in Eqs. (5.38) and (5.39). D Pr
1 D T P s ¼ r rl þ rtrd ð6¼P rT Þ J
ð6:43Þ
which is the nominal stress rate and used for the equilibrium equation of rate-form in the current configuration as described in Eq. (5.43). (c) Covariant convected rates The covariant convected rate of Cauchy stress is given from Eq. (3.34)4 as !
D
D
rCR r gg ¼ FT ðFT rFÞ F1 ¼ r þ lT r þ rl ð¼ rCR Þ T
ð6:44Þ
which is termed the Cotter-Rivlin rate of Cauchy stress (Cotter and Rivlin 1995). Likewise, the covariant convected rate of Kirchhoff stress is given by DCR
s
!
D
s gg ¼ FT ðFT sFÞ F1 ¼ s þ lT s þ sl ð¼ sCR Þ T
ð6:45Þ
(d) Corotational rates The following stress rate based on Eq. (3.39) is termed the Green-Naghdi rate of Cauchy stress (Green and Naghdi 1965).
!
rR R r ¼ RðRT rRÞ RT ¼ r XR r þ rXR ð¼ rRT Þ
ð6:46Þ
Similarly, the Green-Naghdi rate of Kirchhoff stress is given by
!
sR R s ¼ RðRT sRÞ RT ¼ s XR s þ sXR ð¼ sRT Þ
ð6:47Þ
182
6
Objectivity and Objective (Rate) Tensors
The stress rate based on Eq. (3.40) is given by
T
rw r wr þ rwð¼ rw Þ
ð6:48Þ
which is termed the Zaremba-Jaumann rate of Cauchy stress (Zaremba 1903; Jaumann 1911). Likewise, it follows for the Kirchhoff stress that
T
sw s ws þ sw ð¼ sw Þ
ð6:49Þ
The stress rate tensors described above are collectively shown below.
Oldroyd rate of Cauchy stress : rOl r lr rlT ð¼ rOl Þ
T
Truesdell rate of Cauchy stress : rTr rOl þ r trd ¼ r lr rlT þ r tr d ð¼ rTR Þ Covariant-contravariant D
T
D
convected rate of the Kirchhoff stress : t s þ lT s slT ð6¼ t T Þ 1 D D D Nominal stress rate : P r P s ¼ r rlT þ r tr d ð6¼ PrT Þ J D
D
Cotter-Rivlin rate of Cauchy stress : rCR r þ lT r þ rl ð¼ rCRT Þ
T
Green-Naghdi rate of Cauchy stress : rR r XR r þ rXR ð¼ rR Þ
T
Zaremba-Jaumann rate of Cauchy stres : rw r wr þ rw ð¼ rw Þ
ð6:50Þ Here, it should be noted that a corotational rate must be used also for evolution equations of all internal variables in addition to the stress. The above-mentioned rate tensors are used for hypoelastic constitutive equations. In particular, the hypoelastic constitutive equation in terms of the Oldroyd rate is derived directly from the hyperelastic equation as will be described in Sect. 7.4. However, it should be noticed that the hypoelastic-based plastic constitutive equation is limited to the infinitesimal elastic deformation as will be revealed in Sect. 17.3. Incidentally, the particular spin, called the logarithmic spin, by which the corotational rate of the Eulerian-logarithmic (Hencky) strain ln V in Eq. (4.50)2 coincides with the strain rate d, was proposed and the hyperelastic constitutive equation was derived from the hypoelastic constitutive equation in terms of the logarithmic rate of Cauchy stress and the strain rate d by Xiao and his colleagues (Xiao 1995; Xiao et al. 1997, 1999) as explained in Appendix B. However, it would be limited to the isotropy.
6.6 Jaumann Rate with Plastic Spin
6.6
183
Jaumann Rate with Plastic Spin
The physically meaningful rotation of material is the rotation of the substructure of material (material fibers) which cannot be observed from the outside appearance of material. It is not caused by the plastic deformation which causes only the parallel slip of material particles. Therefore, the spin of substructure xs is given by subtracting the plastic spin wp from the continuum spin w, resulting in xs ¼ wwp . Here, the plastic spin wp is determined depending on the constitutive property of plastic anisotropy, while the continuum spin w is determined only by the geometrical variation of outside appearance. Consequently, xs ¼ we holds, i.e.
xs ¼ we ¼ Re Re ¼ w wp T
ð6:51Þ
(Mandel 1971, 1985a; Kratochvil 1971; Dafalias 1983; Loret 1983). The explicit formulation of the plastic spin wp will be given in Sect. 8.8.
6.7
Time-Derivative of Scalar-Valued Tensor Function
Scalar-valued tensor functions of stress and internal variables appear often in continuum mechanics as seen in the strain energy function and the yield function. Then, the time-derivative of scalar-valued tensor function is required in order to derive the rate-type relation of variables, e.g. the consistency condition of yield condition. The time-derivative of scalar function is independent of rigid-body rotation and thus it can be given primarily by its material-time derivative. Here, it should be noticed that the internal variables are formulated by the objective time-derivatives and thus the consistency condition must be transformed to the objective time-derivative. It can be proved that the material-time derivative of scalar-valued tensor function is transformed only to its corotational time-derivative. This fact would seem physically obvious but it must be proved mathematically. To this end, its mathematically exact proof for scalar valued function of general tensor will be given below, referring to the previous studies by Dafalias (1985a, 1998, 2011) for vector and second-order tensor and Hashiguchi (2007b) for general tensor. The corotational rate of the general tensor is defined based on Eq. (3.34) for the second-order tenor as follows: hh _ ii _ t ¼ R ðRT ½½tÞ
ð6:52Þ
_
denoting the general orthogonal tensor by R, where use is made of the symbol ½½ _
for general objective transformation in Eq. (1.103) with the replacement Q ! RT .
184
6
Objectivity and Objective (Rate) Tensors
_
_
Here, noting f ðtÞ ¼ f ðRT ½½tÞ because of the requirement f ðtÞ ¼ f ðRT ½½tÞ for scalar variable, one has
_
_
f ðtÞ ¼ f ðR ½½tÞ ¼ T
@f ðRT ½½tÞ _
@ðR ½½t Þ h h ii @f ðtÞ _ _ T ¼ R ðR ½½tÞ @t
f ðtw1 w2 Þ ¼
T
_
_
_
_
_
_
ðR ½½tÞ ¼ R T
@f ðRu1 w1 Ru2 w2 tu1 u2 Þ
T
_
@f ðtÞ @t
_
ðRT ½½tÞ ð6:53Þ
_
ðRv1 p1 Rv2 p2 tv1 v2
Þ
@ðRs1 p1 Rs2 p2 ts1 s2 Þ _ _ @f ðtw1 w2 Þ _ _ ¼ Rs1 p1 Rs2 p2 ðRv1 p1 Rv2 p2 tv1 v2 Þ @ts1 s2
ð6:54Þ
where the symbol designates the full contraction between derivative components in order between derivative components, i.e. t s ¼ tp1 p2 pm sp1 p2 pm . The derivation of Eq. (6.53) is shown for vector and second-order tensor as follows:
_
_
f ðvÞ ¼ f ðRT vÞ ¼
@f ðRT vÞ _
@ðRT vÞ
_
ðR
T
_
vÞ ¼ RT
@f ðvÞ _ T @f ðvÞ _ _ T ðR vÞ ¼ RðR vÞ @v @v ð6:55Þ
_
_
_
f ðtÞ ¼ f ðR tRÞ ¼ T
_
@f ðRT tRÞ _
_
@ðRT tRÞ
_
_
_
: ðRT tRÞ ¼ RT
@f ðtÞ _ T _ T _ R :ðR tRÞ @t
ð6:56Þ
@f ðtÞ _ _ T _ _ T : RðR tRÞ R ¼ @t Equations (6.53), (6.55) and (6.56) can be satisfied by the corotational rate in Eq. (6.52) amongst objective rates. Then, we have the following relation.
f ðtÞ ¼
f ðtq1 q2
@f ðtÞ @f ðtÞ t ¼ t @t @t @f ðtq1 q2 qm Þ tp1 p2 qm Þ ¼ @tp1 p2 pm
pm
¼
@f ðtq1 q2 @tp1 p2
ð6:57Þ
qm Þ
pm
t p1 p2
pm
which is described for vector and second-order tensor as follows:
f ðvÞ ¼
@f ðvÞ v; @v
f ðvr Þ ¼
@f ðvr Þ vi @vi
ð6:58Þ
6.7 Time-Derivative of Scalar-Valued Tensor Function
f ðtÞ ¼
@f ðtÞ : t; @t
f ðtrs Þ ¼
185
@f ðtrs Þ t ij @tij
ð6:59Þ
It follows from Eq. (6.54) that
_ _ _ _ _ _ @f ðtw1 w2 w3 Þ _ _ _ Rs1 p1 Rs2 p2 Rs3 p3 ðRv1 p1 Rv2 p2 Rv3 p3 þ Rp1 v1 Rp2 v2 Rp3 v3 þ @ts1 s2 s3
Þtv1 v2 v3
_ _ @f ðtw1 w2 w3 Þ _ _ ðRs1 p1 Rv1 p1 ds2 v2 ds3 v3 þ ds1 v1 Rs2 p2 Rv2 p2 ds3 v3 þ @ts1 s2 s3 @f ðtw1 w2 w3 Þ ¼ ðxv1 s1 tv1 s2 s3 þ xs2 v2 ts1 v2 s3 þ Þ ¼ 0 @ts1 s2 s3
¼
Þtv1 v2 v3
ð6:60Þ
which is reduced for vector and second-order tensor as follows:
@f ðvÞ @f ðvu Þ @f ðvu Þ @f ðvÞ Þx ¼ 0 ð6:61Þ xv ¼ xri vi ¼ vi xri ¼ 0; i.e. tr ðv @v @vr @vr @v tr½ð
@f ðtÞ T @f ðtÞ t tT Þx ¼ 0 @t @t
ð6:62Þ
The fulfillments of Eqs. (6.61) and (6.62) require for the tensors in the brackets ð Þ to be zero or symmetric tensor, while Dafalias (1998) has required for the latter to be zero tensor. The fulfillment of Eq. (6.61) is easily known for f ðvÞ ¼ svv because of @f ðvÞ=@v ¼ 2sv with Eq. (1.176)3, and that of Eq. (6.62) for the second-order symmetric tensor t because of ð@f ðtÞ=@tÞtT tT ð@f ðtÞ=@tÞ ¼ O. Equation (6.57) is extended for plural variables as follows:
f ðt1 ; t2 ;
@f ðt1 ; t2 ; @t1 @f ðt1 ; t2 ; ¼ @t1
Þ¼
@f ðt1 ; t2 ; Þ t2 þ @t2 Þ @f ðt1 ; t2 ; Þ t1 þ t2 þ ¼ f ðt1 ; t2 ; @t2
Þ
t1 þ
Þ ð6:63Þ
which is shown for the function of two tensors in Belytschko et al. (2014). Here, it should be noted that the mathematical property does not hold for each term, i.e. @f ðt1 ; t2 ; ; tm Þ @f ðt1 ; t2 ; ; tm Þ ti 6¼ ti @ti @ti
(no sum)
ð6:64Þ
Scalar-valued functions must be independent of rigid-body rotation so that their material-time derivative possess a unique value which coincides with their corotational time-derivative as can be confirmed by the above-mentioned proof.
186
6
Objectivity and Objective (Rate) Tensors
However, they do not lead to the other convected time-derivatives which depend on the rate of deformation, i.e. velocity gradient. Therefore, corotational timederivatives can be adopted in the time-derivatives of scalar functions in constitutive relations, e.g. a strain energy function and a yield function of tensors in the current configuration but convected time-derivatives other than corotational rates cannot be adopted in them. The most popular scalar-valued tensor functions are the principal invariants of tensor. Principal invariants of tensor are described by three independent principal invariants of tensor. Then, the material-time derivatives of the principal invariants are transformed to the corotational time-derivatives merely by replacing all the material time-derivatives of tensor to the corotational time-derivatives of tensor as will be written below. It follows from Eqs. (6.59) and (6.63) that
9 > > > > =
I ¼ ðtrtÞ ¼ ðtrtÞ ¼ I : t ¼ tr t
II ¼ ðtrt2 Þ ¼ ðtrt2 Þ ¼ 2tT : t ¼ 2trðt tÞ
III ¼ ðtrt3 Þ ¼ ðtrt3 Þ ¼ 3t2 : t ¼ 3trðt2 tÞ T
> > > > ; 9 > > > > =
I ¼ ðtrtÞ ¼ tr t o 1n ðtrtÞ2 trt2 ¼ ðtrtÞtr t trðt tÞ II ¼ 2
III ¼ ðdettÞ ¼ ðdettÞtT : t ¼ ðdettÞtrðt1 tÞ
> > > > ;
ð6:65Þ
ð6:66Þ
for the principal invariants in Eqs. (1.222) and (1.217)–(1.219), noting Eqs. (1.354) and (1.355). Further, it holds that
ðt1 : t2 Þ ¼ t1 : t2 þ t1 : t2 ¼ trðtT2 t1 Þ þ trðtT1 t2 Þ
ð6:67Þ
for the two tensor variables. If t1 and t2 are commutative (possessing same principal directions) leading to t1 t2 ¼ t2 t1 ¼ t1 tT2 ¼ tT2 t1 , one has
t1 : t2 ¼ t1 : t2 ;
t1 : t2 ¼ t 1 : t2
ð6:68Þ
noting t1 : ðt2 x xt2 Þ ¼ tr½t1 ðt2 xÞT tr½t1 ðxt2 ÞT ¼ trðtT2 t1 xÞ þ trðt1 tT2 xÞ ¼ 0 ð6:69Þ
6.7 Time-Derivative of Scalar-Valued Tensor Function
187
All the equations in Eqs. (6.65)–(6.67) hold for arbitrary corotational tensors as proved here, although they are written explicitly for the Zaremba-Jaumann rate in some literatures (e.g. Prager 1961b, Belytschko et al. 2014).
Chapter 7
Elastic Constitutive Equations
Elastic deformation is induced by the reversible deformation of material particles themselves without a mutual slip between them. They therefore exhibit higher stiffness compared to the stiffness associated with plastic deformation. Elastic constitutive equations are classifiable into the three types depending on the exactness in the description of reversibility, i.e. the hyperelasticity (or Green elasticity) possessing the strain energy function, the Cauchy elasticity possessing the one-to-one correspondence between stress and strain and the hypoelasticity possessing the linear relation between stress rate and strain rate. As preparation for the study of elastoplasticity in the subsequent chapters, they will be explained previously in this chapter.
7.1
Definition of Hyperelasticity
The material in which a work done per unit reference volume during a certain deformation state to other deformation state is invariant independent of a deformation path is defined as a hyperelastic material or Green elastic material. Therefore, the hyperelastic material possesses the following mechanical properties. (1) An energy is not dissipated/produced during a stress or strain cycle. (2) There exists the one-to-one correspondence between the stress and the deformation. Therefore, the current stress is determined only by the current deformation, and vice versa independently of loading path.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_7
189
190
7.2
7
Elastic Constitutive Equations
Hyperelastic Equations
R The work done per unit reference volume is designated as w0 ¼ w0 dt where w0 is described by the scalar product of the work-conjugate pair of stress and deformation rate in Eq. (5.58). Now, let the right Cauchy-Green deformation tensor C ¼ U2 in Eq. (4.16) be adopted as the deformation measure which designates the pure deformation independent of the rigid-body rotation R, although the deformation gradient tensor F is influenced by the rigid-body rotation superimposed on the deformed configuration. Therefore, the single-valued scalar function wðCÞ for C must exist and the following relation must hold. Z w0 ¼
C
Z dw0 ¼ wðCÞ wðC0 Þ ¼
C0
C
C0
@wðCÞ : dC @C
ð7:1Þ
where wðCÞ is called the elastic strain (or potential) energy function or stored energy function or Helmholtz free energy function. (Note) Substituting DU ¼ Q þ W ¼ TDS þ W (U: internal energy, T: absolute temperature, Q: heat, W: work, S: entropy) into the Helmholtz free energy function F ¼ U TS ! DF ¼ DU TDS under the isothermal process (DT ¼ 0), one has W ¼ DF. Therefore, the work increment done by the external force is identical to the increment of the Helmholtz free energy. Then, it follows that dw0 ¼
@wðCÞ @wðCÞ : dC ¼ 2 : dC=2 @C @C
ð7:2Þ
On the other hand, the work increment done to the unit reference volume is given by dw0 ¼ S : dC=2 in Eq. (5.58). Therefore, the second Piola-Kirchhoff stress is given by S¼2
@wðCÞ @C
ð7:3Þ
where the one-to-one correspondence between the stress S and the deformation variable C exists. Here, note that the function wðCÞ must satisfy the complete integrability condition @Sij @Skl ¼ @Ckl @Cij
ð7:4Þ
which is imposed by the exchangeability of the order of partial derivatives, i.e.
7.2 Hyperelastic Equations
191
@2w @2w ¼ @Cij @Ckl @Ckl @Cij
ð7:5Þ
which holds for an arbitrary scalar continuous function wðCÞ. Now, the strain energy function is chosen to be positive and zero in the reference configuration for convenience, i.e. wðCÞ 0;
wðIÞ ¼ 0
ð7:6Þ
leading to S¼
@w ¼O @C C¼I
ð7:7Þ
Then, the strain energy function wðCÞ takes the minimal-minimum (global minimum) stationary value in the reference configuration C ¼ I (no deformation) and the stress S is zero in the reference configuration C ¼ I (no deformation) as shown in Fig. 7.1. In the above, the hyperelastic constitutive equation is defined in terms of the right Cauchy-Green deformation tensor C which is independent of the rotation of material but depends only on the pure deformation. On the other hand, it is defined first in terms of the deformation gradient tensor F which depends not only on the deformation also on the rotation of material and after that it is reduced to the function of the right Cauchy-Green deformation tensor C or the Green strain tensor E in order to fulfill the objectivity by Holzapfel (2000), Bonet and Wood (2008), de Souza Neto et al. (2008), Gurtin et al. (2010), which would be roundabout explanation compared with the explanation in terms of the right Cauchy-Green deformation tensor C as mentioned above.
S
ψ (C) C=I
C ij
C =I
Fig. 7.1 Rough illustrations of strain energy function and Piola-Kirchhoff stress
C ij
192
7
Elastic Constitutive Equations
On the other hand, the constitutive equation which describes the one-to-one correspondence of stress and strain but does not possess an elastic strain energy function and thus does not fulfill the integrability condition described by the partial derivative of the stress tensor by the strain tensor, resulting in the energy production or dissipation during a loading cycle, is defined as the Cauchy elasticity. Furthermore, a linear relation between a stress rate and a strain rate which does not lead to the one-to-one correspondence between a stress and a strain and in which the strain energy function does not exist and thus the energy is produced or dissipated during a loading cycle is called the hypoelasticity (Truesdell 1955). Here, note that the elastoplastic constitutive equation must be formulated in a rate form because it exhibits the loading-path dependency. Concurrently, the hyperelastic-based plastic constitutive equation describing both the elastic and the plastic deformations exactly must be formulated in terms of the work-conjugate pair
inducing the work rate identical to S : C=2. Here, note that the elastoplastic constative equation described in the intermediate configuration based on the isoclinic concept must be described in terms of the Mandel stress tensor M and the velocity gradient tensor L in the intermediate configuration which are work-conjugate pair fulfilling
s : l ¼ S : C=2 ¼ M : L which is reduced to the hyperelastic equation w0 ¼
e
wðCe Þ ¼ s : le ¼ S : Ce=2 ¼ M : L in the elastic deformation process as will be described in Sect. 17.6. Now, substituting Eq. (7.3) into the stress tensors listed in Table 5.1, we obtain various expressions of the hyperelasticity as follows: 1 @wðCÞ T 1 @wðEÞ T @wðCÞ T @wðEÞ T F ¼ F F ; s ¼ 2F F ¼F F r¼2 F J @C J @E @C @E @wðCÞ @wðEÞ ¼F ; P ¼ 2F @C @E @wðCÞ @wðEÞ @wðCÞ @wðEÞ S¼2 ¼ ; M ¼ 2C ¼ ð2E þ IÞ @C @E @C @E
ð7:8Þ
The tangent stiffness modulus tensor CS 2
@S @ 2 wðCÞ ¼4 @C @C @C
ð7:9Þ
in the hyperelastic relation in terms of the second Piola-Kirchhoff stress tensor S satisfies both the minor symmetry and the major symmetry, i.e. ð7:10Þ
7.2 Hyperelastic Equations
193
Furthermore, noting @ @ @CPQ @ @FrP FrQ @ ¼ ¼ ¼ ðdri dPA FrQ þ FrP dri dQA Þ @FiA @CPQ @FiA @CPQ @FiA @CPQ @ @ @ ¼ FiQ þ FiP ¼ 2FiP @CAQ @CPA @CPA
ð7:11Þ
@ @ @bpq @ @FpR FqR @ ¼ ¼ ¼ ðdip dAR FqR þ FpR diq dAR Þ @FiA @bpq @FiA @bpq @FiA @bpq @ @ @ FqA þ FpA ¼ 2 FqA ¼ @biq @bpi @biq
ð7:12Þ
one has
and
9 @ @ @ > > ¼ 2F ¼F > > @F @C @E > > > > @ 1 1 @ 1 @ > > > > ¼ F ¼ @C 2 @F 2 @E = > @ > > F ¼ F1 > > @b > > > > > @ @ > > 1 @ ; ¼F ¼2 @E @F @C
ð7:13Þ
9 @ @ > ¼2 F > > > > @F @b > = @ 1 @ 1 ¼ F > @b 2 @F > > > @ @ 1 > > ; ¼ F @e @F
ð7:14Þ
by which Eq. (7.8) is described in terms of the deformation measures in the current configuration as follows: 1 @wðbÞ @wðbÞ b; s ¼ 2 b J @b @b @wðbÞ P¼2 F; @b @wðbÞ @wðbÞ F; M ¼ 2FT F S ¼ 2F1 @b @b
r¼2
ð7:15Þ
194
7
Elastic Constitutive Equations
noting @wðCÞ 1 @wðCÞ @wðbÞ ¼ 2 F1 ¼ 2F1 F @C 2 @F @b @wðbÞ T @wðbÞ FF ¼ 2 b s ¼ FSFT ¼ 2FF1 @b @b @wðbÞ @wðbÞ P ¼ FS ¼ 2FF1 F¼2 F @b @b @wðbÞ @wðbÞ F ¼ 2FT F M ¼ CS ¼ 2FT FF1 @b @b S¼2
Noting @ 1 @ 1 ¼ V @b 2 @V
ð7:16Þ
@ @ @b @ @V2 @ ¼2 V ¼ ¼ @V @b @V @b @V @b
ð7:17Þ
derived from
the stress tensors in Eq. (7.15) are expressed in terms of V as follows: @wðVÞ @wðVÞ ; s¼V @V @V @wðVÞ 1 V F; P¼ @V @wðVÞ 1 @wðVÞ 1 V F; M ¼ FT V S ¼ F1 @V @V r ¼ 1J V
ð7:18Þ
If the strain energy function w is represented by the principal stretches ka ða ¼ 1; 2; 3Þ, the principal values of stress tensors are given as follows (Holzapfel, 2000) @w @w ; s a ¼ ka @ka @ka @w ða ¼ 1; 2; 3Þ P a ¼ Ma ¼ @ka @w Sa ¼ k1a @ka ra ¼ 1J ka
ð7:19Þ
by Eq. (7.18) with the replacements of F and V to ka . These principal stresses are expressed by the principal values ra of the Cauchy stress as follows: Sa ¼ J
1 1 ra ; Pa ¼ Ma ¼ J ra ka k2a
ða ¼ 1; 2; 3Þ
ð7:20Þ
7.2 Hyperelastic Equations
195
The stress-strain relation of the isotropic material is given from Eq. (1.304) by S ¼ /E0 I þ /E1 E þ /E2 E2
ð7:21Þ
where /E0 ; /E1 ; /E2 are the functions of invariants of E, while it is not necessary hyperelastic equation. By eliminating the third term, Eq. (7.21) is reduced to the isotropic linear hyperelastic equation: S ¼ aðtrEÞI þ 2bE
ð7:22Þ
possessing the elastic strain energy function 1 wðEÞ ¼ aðtrEÞ2 þ bE2 2
ð7:23Þ
where a and b are the material constants. Now, consider the strain energy function in terms of the invariants of C, i.e. wðCÞ ¼ wðIC ; IIC ; IIIC Þ
ð7:24Þ
The substituting Eq. (7.24) into Eq. (7.8), it follows that 1 @wðCÞ @wðCÞ @wðCÞ 2 Iþ ðIC I CÞ þ ðIIC I IC C þ C Þ FT r ¼ 2 pffiffiffiffiffiffiffiffi F @IC @IIC @IIIC IIIC ð7:25Þ noting Eq. (4.71). It follows by virtue of Eq. (7.3) that @wðCÞ : C¼0 S2 @C
ð7:26Þ
On the other hand, the incompressibility requires 1 J ¼ Jtrd ¼ JtrðFT E F1 Þ ¼ JtrðC1 EÞ ¼ JC1 : E ¼ JC1 : C ¼ 0 ð7:27Þ 2
The following relation must hold from Eqs. (7.26) and (7.27) for the incompressible material. S2
@wðCÞ 1 ¼ rm JC1 @C 2
ð7:28Þ
where rm is the hydrostatic stress and can be determined by the additional equation based on the incompressibility condition J ¼ 1. Eventually, the constitutive equation of the incompressible material is given from Eq. (7.28) by
196
S¼2
7
Elastic Constitutive Equations
@wðCÞ 1 þ rm JC1 @C 2
ð7:29Þ
which is represented by the other stress tensors M ¼ CS ¼ 2C
@wðCÞ þ rm JI @C
ð7:30Þ
1 1 @wðCÞ T r ¼ FSFT ¼ 2 F F þ rm I J J @C
7.3
ð7:31Þ
Explicit Hyperelastic Models
Various explicit hyperelastic constitutive equations will be shown in this section.
7.3.1
St.Venant-Kirchhoff Model
The simplest hyperelastic material model is the St. Venant-Kirchhoff model which is just an extension of the geometrically linear elastic material model to the geometrically nonlinear regime. It possesses the following strain energy function. 1 wðEÞ ¼ kðtrEÞ2 þ ltrE2 ; 2
1 wðCÞ ¼ kftr½ðC IÞ=2g2 þ ltr½ðC IÞ=22 2 ð7:32Þ 0
where the material constants k and l correspond to the Lame constants in the infinitesimal linear elasticity theory. Equation (7.32) is identical to Eq. (7.23) with the replacement of a ! k and b ! l. The substitution of Eq. (7.32) into Eq. (7.8) reads: S ¼ kðtrEÞI þ 2lE ¼ kftr½ðC IÞ=2gI þ lðC IÞ
ð7:33Þ
It is followed that
where
ð7:34Þ
ð7:35Þ
7.3 Explicit Hyperelastic Models
197
Further, the Mandel stress in Eq. (5.22) is given for Eq. (7.33) as M ¼ ktr½ðC IÞ=2C þ lCðC IÞ
ð7:36Þ
The Kirchhoff stress is given from Eq. (7.33) as s ¼ ktr½ðb IÞ=2b þ lbðb IÞ
ð7:37Þ
noting s ¼ FSFT ¼ F½kftr½ðC IÞ=2gg þ lðC IÞFT 1 1 ¼ ktr ðFT F IÞ FFT þ 2lF ðFT FIÞ FT 2 2 The multiplications of C to S, C and I from the left lead to M, CC and C. On the other hand, the contravariant push-forward operation, i.e. the multiplications of F from the left and FT from the right to S, C and I lead to FSFT ¼ s, FCFT ¼ FðFT FÞFT ¼ bb and FIFT ¼ b. Therefore, the hyperelastic equations in terms of ðM; CC; CÞ and ðs; bb; bÞ possess the identical form. It is unfortunate that the St. Venant-Kirchhoff equation is practically useless except for the small deformation range.
7.3.2
Neo-Hookean Model
The neo-Hookean elastic constitutive equation is derived by Rivlin (1948) based on the statistical thermodynamics of cross-linked polymer chains, which is usable for incompressible plastics and rubber-like substances. It is given as 1 1 wðCÞ ¼ lðIC 3Þ ¼ lðk21 þ k22 þ k23 3Þ 2 2
ð7:38Þ
noting Eq. (4.70)1 , where l is the martial constant. The substitution of Eq. (7.38) into Eq. (7.8) with Eqs. (4.71) and (5.22) leads to S ¼ lI
ð7:39Þ
M ¼ lC
ð7:40Þ
s ¼ lb
ð7:41Þ
The neo-Hookean model can be derived through the statistical thermodynamics of cross-linked polymer chains. The model is also inadequate for biaxial states of stress and incompressible materials and has been superseded by the Mooney-Rivlin model and the others described in the following.
198
7
Elastic Constitutive Equations
The hyperelastic constitutive models other than the Neo-Hookean model have been formulated by the statistical thermodynamics of cross-linked polymer. The well-known models are the Arruda-Boyce model (Arruda and Boyce 1993), the Gent model (Gent 1996) and the micro-sphere model (Miehe et al. 2004). The neo-Hookean model was extended to the following various compressible hyperelastic models. (a) Modified Neo-Hookean Model (1) The following strain energy function for the Neo-Hookean model is extended to the compressible material as follows: 1 1 wðCÞ ¼ lðIc 3Þ l ln J þ kðln JÞ2 2 2
ð7:42Þ
where k and l are the material constants, which is adopted for the multiplicative elastoplasticity model by Ulm and Celigoj (2021). It follows from Eq. (7.42) that 1 1 1 1 S ¼ 2 lI l C1 þ k2ðln JÞ C1 2 2 2 2 leading to S ¼ lðI C1 Þ þ kðln JÞC1
ð7:43Þ
pffiffiffiffiffiffiffiffiffiffiffi 8 @J @ det C 1 1 > > ¼ ¼ JC < @C @C 2 pffiffiffiffiffiffiffiffiffiffiffi > @ ln J @ ln det C 1 > : ¼ ¼ C1 @C @C 2
ð7:44Þ
noting
with the aid of Eq. (4.71). Further, the Mandel stress M is given by M ¼ lðC IÞ þ kðln JÞI
ð7:45Þ
Equation (7.43) is described by the Kirchhoff stress as s ¼ FSFT ¼ F½lðI C1 Þ þ kðln JÞC1 FT ¼ F½lðI F1 FT Þ þ kðln JÞF1 FT FT leading to s ¼ lðb IÞ þ kðln JÞI
ð7:46Þ
7.3 Explicit Hyperelastic Models
199
(b) Modified Neo-Hookean Model (2) The following strain energy function is adopted by Simo and Pister (1984), Ciarlet (1988), etc. 1 1 1 wðCÞ ¼ kðJ 2 1Þ ð k þ lÞ ln J þ lðIc 3Þ 4 2 2
ð7:47Þ
from which, noting Eq. (7.44), we have 1 S ¼ kðJ 2 1ÞC1 þ lðI C1 Þ 2
ð7:48Þ
1 M ¼ kðJ 2 1ÞI þ lðC IÞ 2
ð7:49Þ
1 s ¼ kðJ 2 1ÞI þ lðb IÞ 2
ð7:50Þ
Equation (7.47) is reduced to the Neo-Hookean model in Eq. (7.38) for the incompressible material ðJ ¼ 1Þ. (c) Modified Neo-Hookean Elasticity (3) The following strain energy function is used for the finite elastoplastic constitutive relation by Vladimirov et al. (2008, 2010). 1 l wðCÞ ¼ KðJ 2 1 2lnJÞ þ ðtrC 3 2lnJÞ 4 2
ð7:51Þ
where K and l are material constants, from which we have K 2 ðJ 1ÞC1 þ lðI C1 Þ 2
ð7:52Þ
M¼
K 2 ðJ 1ÞI þ lðC IÞ 2
ð7:53Þ
s¼
K 2 ðJ 1ÞI þ lðb IÞ 2
ð7:54Þ
S¼
noting Eq. (7.44). Equation (7.51) will be adopted for the formulation of the finite elastoplastic constitutive equation in Sect. 17.11.1 for the multiplicativehyperelastic-based plasticity for metals. (d) Modified Neo-Hookean Elasticity (4) The following the strain energy function is assumed by Simo and Pister (1984), Ciarlet (1988), Simo and Miehe (1992), etc., which has the separated form into two terms related to the volume change J and the isochoric deformation C.
200
7
Elastic Constitutive Equations
K J2 1 l wðCÞ ¼ lnJÞ þ ðtrC 3 2 2 2
ð7:55Þ
where K and l are the material constants, from which we have 1 1 2 1 2=3 1 G ðtrCÞC S ¼ KðJ 1ÞC þ lJ 2 3
ð7:56Þ
1 1 M ¼ KðJ 2 1ÞG þ lJ 2=3 C' ¼ KðJ 2 1ÞG þ lC' 2 2
ð7:57Þ
1 s ¼ KðJ 2 1Þg þ lb' 2
ð7:58Þ
exploiting Eq. (4.63). Except for the St. Venant-Kirchhoff elastic constitutive equation in Sect. 7.3.1, the hyperelastic equations are described in complex forms in terms of the second Piola-Kirchhoff stress tensor S but it can be described concisely in terms of the Mandel stress tensor M.
7.3.3
Mooney Model
The strain energy function in the Mooney model (Mooney 1940) is given by wðCÞ ¼
C1 ðk21
þ k22
þ k23
3Þ þ C2
1 1 1 þ 2 þ 23 2 k1 k2 k 3
! ð7:59Þ
where C1 and C2 are the material constants. Equation (7.59) is regarded as the particular case of the Ogden model which will be shown in the next subsection. For the incompressible case ðk1 k2 k3 ¼ 1Þ, noting Eq. (4.70), Eq. (7.59) is reduced to the Mooney-Rivlin model: wðCÞ ¼ C1 ðIc 3Þ þ C2 ðIIc 3Þ ¼ C1 ðk21 þ k22 þ k23 3Þ þ C2 ðk21 k22 þ k22 k23 þ k23 k21 3Þ
ð7:60Þ for which one has S ¼ 2C1 I þ 2C2 ðIC I CÞ
ð7:61Þ
M ¼ 2C1 C þ 2C2 ðIC C C2 Þ
ð7:62Þ
7.3 Explicit Hyperelastic Models
201
s ¼ 2C1 b þ 2C2 ðIC b b2 Þ
7.3.4
ð7:63Þ
Ogden Model
The generalized hyperelastic model was proposed by Ogden (1972) referring to the generalized strain tensor in Eq. (4.48) or (4.49) as follows: w¼
N X bn
a n¼1 n
ðka1n þ ka2n þ ka3n 3Þ
ð7:64Þ
where N, bn and an are the material constants. The Ogden model in Eq. (7.64) is reduced to the Mooney model in Eq. (7.59) by choosing N ¼ 2; a1 ¼ 2; b1 ¼ C1 and a2 ¼ 2; b2 ¼ C2 . Equation (7.64) is rewritten as follows: w¼
3 X
xðka Þ
ð7:65Þ
a¼1
with xðka Þ ¼
7.4
N X b
n
a n¼1 n
ðkaan 1Þ
ð7:66Þ
Rate Forms of Hyperelastic Equation
The time-differentiation of Eq. (7.8)4 leads to
S¼
@ 2 wðEÞ :E @E @E
ð7:67Þ
Substituting Eqs. (4.126) and (7.67) into Eq. (6.39), the Truesdell rate of Ol
Kirchhoff stress s
is rewritten as
202
7 DOl
s
¼F
Elastic Constitutive Equations
@ 2 wðEÞ @ 2 wðEÞ DOl dkl : ðFT dFÞ FT sij ¼ FiA FjB FkC FlD @E @E @EAB @ECD
ð7:68Þ DOl
which is the rate of hyperelastic equation in the current configuration. Here, s related to the Zaremba-Jaumann rate of Cauchy stress in Eq. (6.48) as DOl
s
w
¼ Jðr dr rd þ rtrdÞ
is
ð7:69Þ
The Zaremba-Jaumann rate of Cauchy stress is related to the strain rate from these equations as 2 1 @ wðEÞ T r ¼ F ðF dFÞ FT þ dr þ rd rtrd detF @E @E w
ð7:70Þ
which is expressed as w
~ :d r ¼E
ð7:71Þ
~ in the current configuration is where the hyperelastic tangent modulus tensor E given by ð7:72Þ
with 1 Rijkl ðrik djl þ ril djk þ rjk dil þ rjl dik Þ ðRijkl ¼ Rklij ¼ Rjikl ¼ Rijlk Þ 2
7.5
ð7:73Þ
Infinitesimal Strain-Based Elastic Equation
For the infinitesimal deformation, the hyperelastic constitutive equation can be given as r¼ e¼
@wðeÞ ; @e
@/ðrÞ ; @r
r¼
e¼
@w2 ðeÞ : e ¼ E : e; @e @e
@/2 ðrÞ : r ¼ E : r; @r @r
E
@r @ 2 wðeÞ ¼ @e @e @e
E1
@e @ 2 /ðrÞ ¼ @r @r @r
ð7:74Þ ð7:75Þ
7.5 Infinitesimal Strain-Based Elastic Equation
203
where e is the infinitesimal strain in Eq. (4.32). The work w done during the elastic deformation is uniquely determined by the elastic strain energy function before and after the deformation as follows: Z W¼
e
e0
Z r : de ¼
e
e0
@wðeÞ : de ¼ ½wðeÞee0 ¼ wðeÞ wðe0 Þ @e
ð7:76Þ
Adopting the elastic strain energy function 1 wðeÞ ¼ LðtreÞ2 þ Gtre2 2
ð7:77Þ
the stress is given by the linear relation to the elastic strain e as r ¼ Lev I þ 2Ge
ð7:78Þ
i.e. r ¼ ðL þ
2 GÞev I þ 2Ge' 3
ð7:79Þ
which is to be the general isotropic linear elastic equation in Eq. (1.307) or (1.308) with Eq. (1.309) and is referred to as the Hooke’s law, where L and G are called the 0
Lame constants. It follows by taking the trace and the deviatoric part of Eq. (7.78) that rm ¼ Kev ; r0 ¼ 2Ge0
ð7:79Þ
1 1 0 rm ; e 0 ¼ r K 2G
ð7:80Þ
ev ¼ where
2 2 K L þ GðL ¼ K GÞ 3 3
ð7:81Þ
rm ð ðtr rÞ=3Þ is the spherical stress. K and G are called the bulk elastic modulus and the shear elastic modulus, respectively. It follows from Eqs. (7.79) and (7.80) that r ¼ Kev I þ 2Ge' ¼ ðK 23 GÞev I þ 2Ge 1 rm I þ e ¼ 3K
1 1 1 1 r' ¼ ð Þrm I þ r 2G 3K 2G 2G
ð7:82Þ
204
7
Elastic Constitutive Equations
Equation (7.82) are represented as r¼E:e
ð7:83Þ
using the elastic modulus tensor E given by
ð7:84Þ
where is the fourth-order tracing tensor and is the fourth-order deviatoric projection tensor defined in Eqs. (1.191) and (1.194), respectively. The inverse relation between the two equation in Eq. (7.84) is confirmed by ( KT + 2GI ' ) : [(1/ 9) KT + (1/ 2G )I ' ] = (1/ 3)T + I ' = I , noting Eq. (1.194). It follows from Eq. (7.80) for the uniaxial loading process (rij ¼ 0 for i ¼ j 6¼ 1and i 6¼ j) that e11 ¼
1 r11 ; E
m e22 e33 e22 ¼ e33 ¼ r11 ! ¼ ¼ m e11 e11 E
ð7:85Þ
where E¼
9KG ; 3K þ G
m¼
3K 2G 2ð3K þ GÞ
ð7:86Þ
noting 9 h 1 i 1 9K þ 3G 1 1 1 r11 þ ðr11 r11 Þ ¼ r11 þ r11 ¼ r11 > > > > 9K 2G 3 9K 3G 27KG > > = 3K 2G
1 2G 3K > 1 2ð3K þ GÞ > > r11 r11 ¼ r11 ¼ ¼ r11 > > 9KG > 9K 6G 18KG ; 3K þ G ð7:87Þ
e11 ¼
e22
The inverse relations of Eq. (7.86) are given as K
E ; 3ð1 2mÞ
G
E 2ð1 þ mÞ
ðL ¼
vE Þ ð1 þ mÞð1 2mÞ
ð7:88Þ
Here, in the uniaxial loading state, E is the ratio of the axial stress to the axial strain and is called the Young’s modulus, and m is the ratio of lateral strain to axial strain and is called the Poisson’s ratio. The strain energy function in (7.77) with Eq. (7.88) is expressed noting e ¼ ½ðtreÞ=3I þ e0 is expressed as
7.5 Infinitesimal Strain-Based Elastic Equation
vE E ðtreÞ2 þ tre2 2ð1 þ mÞð1 2mÞ 2ð1 þ mÞ vE E 1 2 2 02 ðtreÞ þ ðtre Þ þ tre ¼ 2ð1 þ mÞð1 2mÞ 2ð1 þ mÞ 3 E E 2 ðtreÞ2 þ tre0 ¼ 6ð1 2mÞ 2ð1 þ 2mÞ
205
wðeÞ ¼
ð7:89Þ
Both of the variables ðtre2 Þ for the volumetric deformation and tre0 2 ð¼ e0rs e0rs Þ for the deviatoricdeformation are positive. Then, the Poisson’s ratio must be limited in the range 1\v\1=2
ð7:90Þ
in order that wðeÞ 0 hold in any deformation, e.g. in the pure volumetric deformationðtre0 2 ¼ 0Þ and the pure deviatoric deformation ðtre ¼ 0Þ. Here, note that the lower limit ðm ¼ 1Þ and the upper limit ðm ¼ 0:5Þ in Eq. (7.90) correspond to keep the similar shape and the constant volume, respectively. The former is seen in artificial structures, e.g. honeycomb. Substituting Eq. (7.88) into Eq. (7.84), the elastic modulus tensor is also described using the Young’s modulus and the Poisson’s ratio as follows: E=
E T+ E E δ δ + E 1( , E ijkl = + δ ilδ jk ) − 1 δ ijδ kl ] 3(1 − 2ν ) 3(1 − 2ν ) ij kl 1 +ν [ 2 δ ik δ jl 1 +ν I ' 3
−1 −1 E = 1 − 2ν T + 1 +ν I ' , E ijkl = 1 − 2ν δ ij δ kl + 1 +ν [ 1 (δ ik δ jl + δ ilδ jk ) − 1 δ ijδ kl ] 3E 3E E E 2 3
ð7:91Þ
or E = E (S + ν T ), E ijkl = E [ 1 (δ ik δ jl + δ ilδ jk ) + ν δ ijδ kl ] 1 +ν 1 − 2ν 1 +ν 2 1 − 2ν −1 −1 E = 1 [(1 +ν )S −νT ], E ijkl = 1 [ 1 (1 +ν )(δ ik δ jl + δ ilδ jk ) −νδ ijδ kl ] E E 2
ð7:92Þ
noting Eqs. (1.191)–(1.195), which is confirmed by −1 E : E = E (S + ν T ) : 1 [(1 + ν )S −νT ] E 1 +ν 1 − 2ν
(1 +ν ) ) S :T − 1 ν− 22ν T :T ] = 1 [(1 +ν )S + (−ν + ν 1 +ν 1 − 2ν −ν (1 − 2ν ) +ν (1 +ν ) − 3ν 2 T =S =S + (1 +ν )(1 − 2ν )
ð7:93Þ
206
7
Elastic Constitutive Equations
noting Eqs. (1.1.91) and (1.1.92) Incidentally, by the matrix notation 2
3 T1111 T1122 T1133 T1123 T1131 T1112 6 7 6 T2211 T2222 T2233 T2223 T2231 T2212 7 6 7 6T 7 6 3311 T3322 T3333 T3323 T3331 T3312 7 T ¼ ½Tijkl ¼ 6 7 6 T2311 T2322 T2333 T2323 T2231 T2312 7 6 7 6T 7 4 3111 T3122 T3133 T3123 T3131 T3112 5 T1211 T1222 T1233 T1223 T1231 T1212
ð7:94Þ
as explained in Appendix C, where the components ð9 9Þ are abbreviated to ð6 6Þ, Eq. (7.92) is represented as follows: 2 6 6 6 E 6 E¼ ð1 þ mÞð1 2mÞ 6 6 4
1v
2 E1
v v 1v v 1v Sym:
m m 1 Sym:
0 0 0 1þv
0
0
0
1v v 0 1v 0
0 0
0 0
0
0
1 2v
0
1
6 6 16 ¼ 6 E6 6 4
m 1
0 0 0 1 2v
3 0 0 7 0 0 7 7 0 0 7 ð7:95Þ 7 0 0 7 5 1 2v 0 1 2v 3 0 7 0 7 7 0 7 7 0 7 5 0 1þv
0 0 0 0 1þv
ð7:96Þ
which is confirmed by 2
1
E:E
6 6 6 6 E 6 ¼ ð1 þ mÞð1 2mÞ 6 6 6 4 2
1v
v
1
60 6 6 60 6 ¼6 60 6 40 0
0
0
1
0
0
1
0
0
0 0
0 0
1 2v
Sym:
ð1mÞm2 m2
6 mmð1mÞm2 6 6 6 mmð1mÞm2 1 6 ¼ ð1 þ mÞð1 2mÞ 6 60 6 40 2
v
0 3 0 0 0 0 0 07 7 7 0 0 07 7 1 0 07 7 7 0 1 05 0 0
3 2 7 6 7 6 7 6 716 7 6 7E6 7 6 7 6 5 4
1
m
m
0
0
0
1
m 1
0 0
0 0
0 0
1þv
0
0
1þv
0
Sym:
3 7 7 7 7 7 7 7 7 5
mð1mÞ þ mm2
1 2v mð1mÞm2 þ m
0
0
1þv 0
ð1mÞm2 m2
mð1mÞm2 þ m
0
0
0
mmð1mÞm2
ð1mÞm2 m2
0
0
0
0
0
ð1 þ mÞð12mÞ 0
0 0
0 0
0 0
ð1 þ mÞð12mÞ 0
3
0 0 ð1 þ mÞð12mÞ
7 7 7 7 7 7 7 7 5
1
ð7:97Þ
7.5 Infinitesimal Strain-Based Elastic Equation
207
Table 7.1 Relationships between two independent elastic constants E; m
G; m
E; G
E; K
E
E
2ð1 þ mÞG
E
E
G
E 2ð1 þ mÞ E 3ð1 2mÞ
G
G EG 3ð3G EÞ E 2G 2G GðE 2GÞ 3G E
3EK 9K E K
m
m
2ð1 þ mÞG 3ð1 2mÞ m
L
mE ð1 þ mÞð1 2mÞ
2Gm 1m
K
G; K 9KG 3K þ G G K
3K E 6K 3Kð3K EÞ 9K E
3K 2G 2ð3K þ GÞ 2 K G 3
L; G lð3L þ 2GÞ LþG G 2 G 3 L 2ðL þ GÞ L
Lþ
Relationships between two independent elastic constants are listed in Table 7.1. The elastic strain energy function wðeÞ and the complementary energy function /ðrÞ are given for the linear elasticity as ⎧ ψ (ε) = 1 ε : E: ε = 1 E [ε ij εij + E ( ε kk ) 2 ] ⎪ 2 2 1 +ν 1 − 2ν ⎪ ⎨ ⎪φ (σ ) = 1 σ : E −1 : σ = 1 [(1+ν )σ σ −ν (σ ) 2 ] ij ij kk 2 ⎪ 2E ⎩
ð7:98Þ
from which it follows that ⎧ ∂ψ ν E 1 1 ⎪σ = ∂ε = E: ε = 1 + ν (1 − 2ν ε v I + ε ) = E[3(1 − 2ν ) ε v I + 1 + ν ε' ] ⎪ ⎨ 1 ⎪ε = ∂φ = E−1 : σ = 1 [ ν σ − ν σ I ] = E [(1 − 2ν ) σm I + (1 + ν )σ' ] ⎪ ∂σ E (1+ ) 3 m ⎩
ð7:99Þ
The rate forms of the infinitesimal elastic equation is given from Eqs. (7.80) and (7.99) as follows: 8 2 > > > < r ¼ K ev I þ 2G e' ¼ K 3 G ev I þ 2G e ð7:100Þ > 1 1 1 1 1 > >e¼ rm I þ r' ¼ rm I þ r : 3K 2G 3K 2G 2G
r m ¼ K ev ;
r' ¼ 2G e'
8 1 1 E m > > ' r e e ev I þ e ¼ E I þ ¼ < v 3ð1 2mÞ 1þm 1 þ m 1 2m h i > > : e ¼ 1 ð1 2mÞ rm I þ ð1 þ mÞ r ' ¼ 1 ½ð1 þ mÞ r 3m rm I E E
ð7:101Þ
ð7:102Þ
208
7.6
7
Elastic Constitutive Equations
Cauchy Elasticity
The elastic material which does not have a strain energy function but has a one-to-one correspondence between the Cauchy stress and a strain is called the Cauchy elastic material. Here, the stress tensor is given by an equation of strain tensor and thus the equation includes six strain components. The equation of six strain components does not fulfill the condition of complete integration leading to the strain energy function so that it does not result in the hyperelasticity in general. Then, the work done by the stress is generally dependent on the deformation path. For that reason, an energy dissipation/production is induced during the stress or strain cycle. In the above-mentioned definition, the Cauchy elastic material is described as r ¼ fðeÞ
ð7:103Þ
in terms of the Almansi strain tensor e in Eq. (4.27) or (4.29). Equation (7.103) is reduced to the following equation by virtue of Eq. (1.304) for the isotropic material. r ¼ /e0 I þ /e1 e þ /e2 e2
ð7:104Þ
where /e0 ; /e1 ; /e2 are functions of invariants of e. Furthermore, for an isotropic linear elastic material, Eq. (7.104) is reduced r ¼ LðtreÞI þ 2Ge
ð7:105Þ
noting Eq. (7.78). Limiting to the infinitesimal strain leading to e ffi e, Eq. (7.105) results in Eq. (7.76), i.e. r ¼ LðtreÞI þ 2Ge
ð7:106Þ
Here, substituting Eq. (7.106) with Eq. (4.32) into Eq. (5.33), the Navier’s equation is obtained by replacing L and G to a and b, respectively, as follows:
ða þ bÞ$ð$ uÞ þ bDu þ qb ¼ q v ða þ bÞ
@ 2 uj @ 2 ui þb þ qbi ¼ q vi @xj @xi @xj @xj
noting Eqs. (1.374), (1.380) and h
1 @ui k @ a @u d þ 2b ij 2 @xj þ @xk @xj
@uj @xi
i ¼a
@ 2 uj @ 2 ui @ 2 uj þb þb @xj @xi @xj @xj @xj @xi
ð7:107Þ
7.7 Hypoelasticity
7.7
209
Hypoelasticity
The following relation, called the hypoelasticity, was proposed by Truesdell (1955) in order to take the material rotation into account.
r ¼ HðrÞ : d
ð7:108Þ
where HðrÞ is the forth-order tensor-valued isotropic function of the stress tensor r and it is called hypoelastic response function. However, the hypoelastic equation is incapable of describing the constitutive equation exactly, since d does not mean the pure rate of deformation as shown in Eqs. (4.83) and (4.88). In addition, there does not exist a one-to-one correspondence between the stress and the deformation and the work done during a certain variation of stress depends on the loading path. If we adopt the Hookean elastic modulus tensor E for HðrÞ, Eq. (7.108) is reduced to
r¼E:d
ð7:109Þ
which is described as follows, noting Eqs. (7.100) and (7.102). 9 2 > > K G dv I þ 2Gd > = 3 1 1 0 1 1 1 > > rm I þ r rm I þ r> d¼ ; 3K 2G 3K 2G 2G
r ¼ Kdv I þ 2Gd0 ¼
rm ¼ Kdv ; r0 ¼ 2Gd0 9 1 1 0 E m > dv I þ d ¼ dv I þ d > r¼E = 3ð1 2mÞ 1þm 1 þ m 1 2m i> i 1h 1h > d¼ ð1 2mÞ rm I þ ð1 þ mÞ r0 ¼ ð1 þ mÞ r 3m rm I ; E E
ð7:110Þ
ð7:111Þ
ð7:112Þ
Incidentally, the following equation in which the Jaumann rate of Cauchy stress is related nonlinearly to the strain rate is called the hypoplastic material (Kolymbas and Wu 1993) named by Dafalias (1986).
r ¼ fðd; rÞ;
rij ¼ fij ðdkl ; rkl Þ
ð7:113Þ
pffiffiffiffiffiffiffiffiffiffiffi where fij is the nonlinear function of dkl , e.g. drs drs ð¼ jjdjjÞ. However, the hypoplastic constitutive equation would not express the real elastoplastic deformation behavior as will be described in Sect. 8.1.
210
7
Elastic Constitutive Equations
Equation (7.109) is modified to the following rational relation in terms of the
work-conjugate pair ðs; dÞ as shown in Eq. (5.58) in Sect. 5.8. τo = E : d
ð7:114Þ
While the three major types of elastic materials are described in this chapter, the other elastic material, called the Cosserat elastic material, was advocated by Cosserat and Cosserat (1909). The couple stress is related to the rotational strain in this material. It has been applied to the prediction of localized deformation (e.g. cf. Mindlin, 1963; Muhlhaus and Vardoulaskis, 1987).
Chapter 8
Elastoplastic Constitutive Equations
Elastic deformation is induced microscopically by the elastic deformations of the material particles themselves, exhibiting a one-to-one correspondence to the stress as described in Chap. 7. However, when the stress reaches an yield stress, slippages between material particles (e.g. crystal lattice in metals and soil particles in soils) are induced, which do not disappear even if the stress is removed, leading macroscopically to the plastic deformation. Then, the one-to-one correspondence between the stress and the strain, i.e. the stress–strain relation does not hold in the elastoplastic deformation process, exhibiting the loading-path dependence. Therefore, one must formulate the elastoplastic constitutive equation as a relation between the stress rate and the strain rate. This chapter addresses the basic concept and formulation for elastoplastic constitutive equations in the conventional elastoplasticity (Drucker 1988) based on the assumption that the inside of the yield surface is a purely elastic domain as the introduction to elastoplasticity, while the formulations are given within the framework of the infinitesimal strain theory in terms of the material-time derivative. The unconventional elastoplasticity describing the plastic strain rate induced by the rate of stress inside the yield surface will be described in the subsequent chapters, which is required to describe the cyclic loading behavior, and further the exact finite strain theory based on the multiplicative decomposition of the deformation gradient tensor will be given in Chap. 17.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_8
211
212
8.1
8
Elastoplastic Constitutive Equations
Fundamental Requirements for Elastoplastic Constitutive Equations
The elastoplastic constitutive equations must be formulated to satisfy the following requirements. (1) Decomposition of deformation measure into elastic and plastic parts The deformation of solids which are assemblies of sloid particles, e.g. crystal particles in metals and soil particles in sands and clays are induced by the deformations of material particles themselves and their mutual slips. The deformations of material particles themselves cause the macroscopic elastic deformation of the material as far as a large stress such as material particles deforms irreversibly is not applied. On the other hand, when the macroscopic stress applied to the material increases up to the yield stress, the mutual slips between the material particles are not removed even if the stress is removed (unloaded), so that they cause the plastic deformation of the material. Therefore, the decomposition of strain rate into the elastic and the plastic parts is inevitable for the formulations of the evolution rules of internal variables which is irrelevant to the elastic deformation but changes with the plastic deformation history in order to describe the irreversible change of the internal structure. Eventually, the deformation measure of material is decomposed into the elastic part and the plastic part. (2) Rate(incremental) form equation by loading-path dependence The plastic deformation induced during the change from a certain stress to the other stress depends on the stress path during which the stress traces. Therefore, it exhibit the loading path dependence. Then, the constitutive equation describing the deformation including the plastic deformation is obliged to be formulated in the rate (incremental) form. The elastic deformation reflects the variation of the internal structure of material, i.e. the elastic deformational state of material particles. Although the plastic deformation itself is irrelevant to the internal structure of material, the plastic internal variable reflects the variation of the internal structure of material which influences on the plastic deformational characteristics. The examples of the plastic internal variables are the isotopic hardening variable and the kinematic hardening variable, i. e. the backstress which will be described in the subsequent sections.
8.1 Fundamental Requirements for Elastoplastic Constitutive Equations
213
(3) Incorporation of yield surface The stress applied between the material particles needs to increase to overcome the friction resistance in order to induce their mutual slips leading to the macroscopic plastic deformation. Then, the stress which causes the plastic deformation is called the yield stress. Therefore, the yield surface must be incorporated, which is formed by connecting the yield stresses in various directions. In addition, the incorporation of the yield surface is required for the formulation of the plastic strain rate which is associated to the yield surface. Besides, the above-mentioned basic requirements for the constitutive equation of irreversible deformation are ignored in the hypoplasticity in which the rate-nonlinear terms of the strain rate are incorporated. The hypoplasticity was disused in the field of metal plasticity twenty years later after the proposition of the hypoelasticity by Truesdell (1955) but sadly it is still used in the field of soil plasticity.
8.2
Classification of Elastoplastic Constitutive Equations
The elastoplastic constitutive equations can be classified into the three types described below.
8.2.1
Infinitesimal Hyperelastic-Based Plasticity
The infinitesimal strain tensor e in Eq. (4.31) or (4.32) is additively decomposed into the elastic strain tensor ee and the plastic strain tensor ep as follows:
e ¼ e e þ ep ; e ¼ ee þ e p
ð8:1Þ
The Cauchy stress is given by the hyperelastic relation through the strain energy function wðee Þ in Eq. (7.74) as follows: r¼
@wðee Þ @ee
ð8:2Þ
Then, the work We done during the elastic deformation process is uniquely determined by the values of the strain energy function before and after the elastic deformation based on Eq. (7.76), i.e., Z w ¼ e
ee ee0
Z r : de ¼ e
ee ee0
e @wðee Þ : dee ¼ ½wðee Þee0 ¼ wðee Þ wðee0 Þ @ee
214
8
Elastoplastic Constitutive Equations
Further, the strain energy function wðee Þ is given in the quadratic form for the linear elasticity: 1 wðee Þ ¼ ee : E : ee 2
ð8:3Þ
r ¼ E :ee ¼ E :ðe ep Þ; ee ¼ E1 :r; e ¼ E1 :r þ ep
ð8:4Þ
leading to
which further leads to
p
r ¼ E : ee ¼ E : ðe e p Þ; ee ¼ E1 : r; e ¼ E1 : r þ e
ð8:5Þ
for the materials possessing the constant elastic modulus tensor, i.e. E ¼ const The stress r is calculated by substituting the plastic strain ep calculated by the plastic constitutive equation into Eq. (8.4). On the other hand, it can be calculated also by the time-integration by Eq. (8.5). The accuracy and the efficiency in the numerical calculations by these two methods are almost same. This can be stated also for tensor-valued internal variables, e.g. the kinematic hardening variables and the elastic-core which will be described in the subsequent sections and chapters. The stress rate versus elastic strain rate for the Hooke’s law is given from Eqs. (7.82) and (7.89) as follows: 9 2 > e e > ¼ K G ev I þ 2Ge r¼ > = 3 1 1 0 1 1 1 > > ee ¼ rm I þ r ¼ rm I þ r> ; 3K 2G 3K 2G 2G Keev I þ 2Gee0
r0 ¼ 2Gee0
rm ¼ Keev ;
ð8:6Þ
ð8:7Þ 9 e e > ev I þ e > =
1 1 e0 E m eev I þ e ¼ 3ð1 2mÞ 1þm 1 þ m 1 2m 1 1 ee ¼ ½ð1 2mÞrm I þ ð1 þ mÞr0 ¼ ½ð1 þ mÞr 3mrm I E E r¼E
> > ;
ð8:8Þ
where 1 eev tree ; ee0 ee eev I 3
ð8:9Þ
The infinitesimal hyperelastic-based plasticity is limited to the infinitesimal elastic and plastic deformation without a rotation of material. Elastoplastic constitutive equations will be described in the infinitesimal hyperelastic-based plasticity
8.2 Classification of Elastoplastic Constitutive Equations
215
in the subsequent sections, since it is quite simple to be applied to various engineering problems in practice and it can be transformed directly to the multiplicative hyperelastic-based plasticity in Chap. 17. Further, Eq. (8.5) will be used for soils, etc. even if the elastic modulus tensor E is not con-stant but it is a function of stress r as will be described in Chap. 13, in which a strain energy function in Eq. (8.3) does not hold usually so that Eqs. (8.4) and (8.5) cannot lead to the hyperelastic equation in that case.
8.2.2
Hypoelastic-Based Plasticity
A constitutive equation describing the inherent deformation properties of materials which is not influenced by the rigid-body rotation of materials. However, the influence of the rigid-body rotation is not excluded in the infinitesimal strain formulation. Then, the hypoelasticity was proposed by Truesdell (1955) as the modification of the infinitesimal strain formulation so as to exclude the influence of the rigid-body rotation as described in Sect. 7.7. The strain rate d in Eq. (4.62) is additively decomposed into the elastic strain rate de and the plastic strain rate dp , i.e. d ¼ de þ dp
ð8:10Þ
The corotational stress rate is given by
r ¼ E : de ¼ E : ðd dp Þ; de ¼ E1 : r
ð8:11Þ
based on Eq. (7.109). The stress rate versus elastic strain rate is given for the Hooke’s law from Eqs. (7.110) as follows: ) r ¼ Kdve I þ 2Gde0 ¼ K 23 G dve I þ 2Gde 1 1 1 0 1 1 rm I þ 2G de ¼ 3K rm I þ 2G r ¼ 3K 2G r
rm ¼ Kdve ;
0 e0 r ¼ 2Gd
9 1 1 e0 E m > dve I þ d ¼ dve I þ de > > = 3ð1 2mÞ 1þm 1 þ m 1 2m > > 1 1 0 ; de ¼ ½ð1 2mÞ rm I þ ð1 þ mÞ r ¼ ½ð1 þ mÞ r 3m rm I > E E
ð8:12Þ ð8:13Þ
r¼E
ð8:14Þ
216
8
Elastoplastic Constitutive Equations
where dve trde ;
1 de0 de dve I 3
ð8:15Þ
The convected time-derivative of internal variables, e.g. the kinematic hardening must be also given by the convected time-derivative explained in Sect. 3.4. The hypoelastic-based plastic constitutive equation is obtained from the infinitesimal strain-based elastoplasticity by replacing
e
p
r ! r; e ! d; e ! de ; e ! dp ; e ! d The hypoelastic-based plasticity is capable of describing the finite plastic deformation and rotation. It has been studied widely by numerous researchers represented by Rodney Hill and James R. Rice after the proposition of hypoelasticity (Truesdell 1955). However, the hypoelastic-based framework possesses the following various defects. (1) The accurate description of elastic deformation is limited to be infinitesimal, (2) The energy dissipation and its accumulation are caused even within a purely elastic range during a cyclic loading (cf. Kojic and Bathe 1987; Brepols 2014), (3) A cumbersome time-integration scheme is required for the calculation of the stress and internal variables, and then a careful treatment is necessary in the numerical time-integration scheme to guarantee the objectivity of the constitutive relation as was described in detail in Sect. 3.4.4. (4) There exists the arbitrariness or the non-uniqueness regarding the choice of a corotational rate (cf. Peric 1992). Then, the hypoelastic-based plasticity will be disused by the skip from the infinitesimal hyperelastic-based plasticity to the multiplicative hyperelastic-based plasticity in the near feature.
8.2.3
Multiplicative Hyperelastic-Based Plasticity
The exact description of the finite deformation and rotation can be described by the multiplicative hyperelastic-based plasticity, in which the deformation gradient tensor is multiplicatively decomposed into the elastic and the plastic parts as will be explained in detail in Chap. 14. The primary purpose of this monograph is the explanation of the exact finite strain theory within the framework of the multiplicative hyperelastic-based plasticity. The multiplicative hyperelastic-based plasticity cannot be formulated from the hypoelastic-based plasticity but can be derived as the extension of the infinitesimal hyperelastic-based plasticity. Therefore, the formulation of the infinitesimal hyperelastic-based plasticity will be described in detail prior to the extensions.
8.3 Conventional Plastic Constitutive Equation
8.3
217
Conventional Plastic Constitutive Equation
The conventional plastic constitutive equation incorporating the yield surface enclosing a purely-elastic domain is described in this section. Now, consider first the following isotropic yield condition incorporating the isotropic hardening/softening. f ðrÞ ¼ FðHÞ
ð8:16Þ
where Fð 0Þ is the function of the isotropic hardening variable H and is called the hardening function which describes the isotropic hardening or softening, i.e. the expansion or contraction of yield surface. Here, we choose the yield stress function f ðrÞð 0Þ in Eq. (8.16) to be the homogeneous function of r in degree-one. Therefore, it follows that f ðjsjrÞ ¼ jsjf ðrÞ
ð8:17Þ
@f ðrÞ : r ¼ f ðrÞ ¼ FðHÞ @r
ð8:18Þ
for an arbitrary scalar s and
for the sake of Euler’s theorem for homogeneous function in degree-one (see Appendix D). Then, it follows from Eq. (8.18) that @f ðrÞ
@f ðrÞ
: r @f ðrÞ n : r
=
1=
¼ @r
¼ f ðrÞ @r @r F
ð8:19Þ
where n is the normalized outward-normal of the yield surface (see Appendix E). @f ðrÞ n @r
@f ðrÞ
@r ðknk ¼ 1Þ
ð8:20Þ
The material-time derivative of Eq. (8.16) leads to the consistency condition: @f ðrÞ 0 : r F H ¼ 0 @r
ð8:21Þ
which is ttransformed to the following normalized equation by multiplying Eq. (8.19).
0
n : r F H
n:r ¼0 F
ð8:22Þ
218
8
Elastoplastic Constitutive Equations
where p 0 p p p F dF=dH; H r; H; e ¼ fHd r; H; e jje k ke k
ð8:23Þ
p
H has to be the homogeneous function of e in degree-one for the rate-independent deformation behavior because it evolves only in the plastic loading process ðe p 6¼ OÞ and possesses the dimension of time in minus one, while, needless to say, it is a nonlinear equation of the components eijp in general as seen in metals, i.e. ffiffiffiffiffiffiffiffiffiffi ffi q pffiffiffiffiffiffiffiffi p pffiffiffiffiffiffiffiffi p p H ¼ 2=3 jje k ¼ 2=3 ers ers described in Sect. 8.4. Now, assume the associated flow rule in which the plastic potential function is given by the yield function:
p
p k ¼ ke k [ 0
e ¼ kn
ð8:24Þ
where k is the magnitude of plastic strain rate, called often plastic positive proportionality factor or plastic multiplier. The expression of the flow rule in Eq. (8.24) possesses the clear physical meaning as definitely divided into the
magnitude k and the pure direction tensor n. Then, the magnitude of plastic strain
p
rate, i.e. ke k appearing often in the hardening variable can be replaced to k so that the physical interpretation of plastic constitutive relation can be captured clearly.
On the other hand, the magnitude cannot be represented only by k in the flow rule
with the expression e p ¼ k @f ðrÞ=@r except for the particular yield function given by the magnitude of stress jjrjj leading to jj@f ðrÞ=@rjj ¼ jj@jjrjj=@rjj ¼ 1. The physical backgrounds of the associated flow rule will be explained in Sect. 8.6. Substituting the plastic flow rule in Eq. (8.24) into the consistency condition (8.22), one has
n :r
0
F k fHn ðr; H; nÞn : r ¼ 0 F
ð8:25Þ
which leads to
n : r k Mp ¼ 0
ð8:26Þ
where M p is called the plastic modulus and is given by 0
Mp
F fHn ðr; H; nÞn : r F
ð8:27Þ
8.3 Conventional Plastic Constitutive Equation
219
with
fHn ðr; H; nÞ ¼ H = k
ð8:28Þ
noting Eq. (8.23) with Eq. (8.24). Here, the yield surface in Eq. (8.16) retains the similar shape and orientation with respect to the origin of stress space by virtue of homogeneity of function f ðrÞ. It follows from Eq. (8.26) that
n : r p n : r ; e ¼ p n Mp M
k¼
ð8:29Þ
Substituting Eqs. (8.11) and (8.29) into Eq. (8.5), the strain rate is given by
e ¼ E1 : r þ
n :r n n 1 n ¼ E þ :r Mp Mp
ð8:30Þ
The scalar product of n : E to Eq. (8.30) leads to n :r n :r n :r p p n : E : e ¼ n :r þ n : E : n ¼ ðM þ n : E : nÞ p ¼ ðM þ n : E : nÞ p Mp M M
¼ ðM p þ n : E : nÞ k
from which the magnitude of plastic strain rate described in terms of strain rate,
denoted by K instead of k, in the flow rule (8.24) is described as follows:
K¼
n : E :e n : E :e p ; e ¼ p n p M þn : E : n M þn : E : n
ð8:31Þ
Incidentally, Eq. (8.31)1 can be also derived directly from the following relation obtained by substituting Eqs. (8.1) and (8.5) into the consistency condition (8.26).
e
p
n : r K M p ¼ n : E : e K M p ¼ n : E :ðe e Þ K M p
¼ n : E :ðe K nÞ K M p ¼ n : E : e ðn : E : n þ M p Þ K ¼ 0 ð8:32Þ The stress rate is given from Eq. (8.5) with Eq. (8.31) as follows:
e
p
p
r ¼ r þ r ¼ E :ðe e Þ ¼ E : e
n : E :e E:n¼ Mp þ n : E : n
E
E:nn:E :e Mp þ n : E : n
220
8
e
p
Elastoplastic Constitutive Equations
n pq pqkl ε• kl ijrsn rs M p + nab abcd n cd
p
σ• ij = σ• ije + σ• ij = ijkl (ε• kl −ε• kl ) = ij kl ε• kl −
ð8:33Þ
where e
r E : e;
p
p
r E : e ¼ Kpr : e
ð8:34Þ
E ijrs nrs npq E pqkl p :n n K pr ≡ E p ⊗ : E , K ijklr ≡ M + n: E: n M p + nab E abcd ncd
ð8:35Þ
e
r is called the elastic stress rate since it is the stress rate calculated supposing that p a purely elastic deformation is induced, and r is called the plastic relaxation stress pr rate. The fourth-order tensor K is called the plastic relaxation stiffness modulus tensor. Furthermore, using the elastoplastic stiffness modulus tensor
ijrs n n pq pqkl ep pr : n ⊗ n : , ijkl = ijkl − p rs ≡ ijkl − ijkl ep ≡ − pr = − p M + nab abcd ncd M + n: : n ð8:36Þ the stress rate can be described as
r ¼ Kep : e
ð8:37Þ
Here, note that E possesses not only the minor symmetry but also the major symmetry E ijkl = E klij providing ( E : n)ij = ( n : E)ij = E ijkl nkl, so that the symmetries of E : n n : E ¼ ½ðE : nÞ ðn : EÞT and thus Kep ¼ KepT hold. Here, consider the non-associated flow rule
p
e ¼ k m ðjjmjj ¼ 1; m 6¼ nÞ
ð8:38Þ
where m is the normalized second-order function of stress and internal variables.
The magnitude of plastic strain rate K is given instead of Eq. (8.31) as follows:
n : E :e n : re ¼ p K¼ p M þn : E : m M þn : E : m
! ð8:39Þ
noting Eq. (8.34), Then, the elastoplastic stiffness modulus tensor is given by Kep ¼ EKpr ¼ E
E:mn:E Mp þ n : E : m
ð8:40Þ
8.3 Conventional Plastic Constitutive Equation
221
Therefore, the plastic relaxation modulus tensor and the elastoplastic stiffness modulus tensor are not the symmetric tensors, i.e. Kp 6¼ KpT ; Kep 6¼ KepT in the non-associated flow rule, which causes the complexity and the inefficiency in numerical calculation.
8.4
Constitutive Equation of Metals
The general form of the elastoplastic constitutive equations for isotropic materials are described above. The constitutive equation of metals is shown below, which has made an important contribution to the development of elastoplasticity. The following von Mises yield condition (von Mises, 1923) with the associated flow rule can be assumed for metals. rffiffiffi pffiffiffiffiffiffiffi 3 0 jjr jj ¼ 3J2 2 qffiffi qffiffi p ¼ 23jj e jj; fHn ¼ 23
f ðrÞ ¼ req ; req (
H ¼ eeqp
Rqffiffi2
p
3jj e
jjdt; H ¼ e
eqp
ð8:41Þ
0
FðHÞ ¼ F0 f1 þ sr ½1 expðcH HÞg; F dF=dH ¼ sr cH F0 expðcH HÞ ð8:42Þ where F0 ; sr and cH are the material constants. The hardening function F in Eq. (8.42) increases from the initial value F0 with the equivalent plastic strain eeqp and saturates when it reaches the maximum value ð1 þ sr ÞF0 as shown in Fig. 8.1.
F( Fs
eqp)
(1 sr ) F0 1
F0
F ' = sr cH F0 exp( cH
0 Fig. 8.1 Isotropic hardening function in the uniaxial loading process
eqp)
eqp
222
8
Elastoplastic Constitutive Equations
In the monotonic uniaxial loading with the assumption of the plastically-pressure p p p p p independence (rij ¼ 0 except for i ¼ j ¼ 1; e2 ¼ e3 ¼ e1 =2 ðtr e ¼ 0Þ, eij ¼ 0 ði 6¼ jÞ), it holds that 8 pffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi < req 3=2jjr0 jj ¼ 3=2 ðr1 r1 =3Þ2 þ 2ð0 r1 =3Þ2 ¼ r1 ð8:43Þ ffi R pffiffiffiffiffiffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R p p : eeqp R pffiffiffiffiffiffiffi 2 2=3jjep0 jj dt ¼ 2=3 ep2 e1 dt ¼ ep1 1 þ 2ð e1 =2Þ dt ¼
Then, req and eeqp coincide with the axial stress and the axial plastic strain in the uniaxial loading and thus are called the equivalent stress and the equivalent (or accumulated) plastic strain, respectively. Substituting Eq. (8.42) into Eqs. (8.30) and using the relations 9 rffiffiffi @f ðrÞ 3 r0 r0 0 > > ; n ¼ 0 ; n : r ¼ jjr jj > ¼ > > jjr jj @r 2 jjr0 jj > > > r ffiffi ffi > rffiffiffi 0 > 0 > 3 r 3r :r 3 = 0 eq : r n : r ¼ ¼ r ¼ 0 eq ð8:44Þ 2 jjr jj 2 r 2 > > > p p0 > > e ¼e > > rffiffiffi > 0 > > F 2 2 0 > 0 p ; M ¼ jjr jj ¼ F F 3 3 the constitutive equation of the isotropic Mises material is given as follows: qffiffi
n :r e ¼ E1 : r þ 2 0 n ¼ E1 : r þ 3F
2 eq 3r 2 0 3F
0
eq
r 31r ¼ E1 : r þ r0 0 jjr jj 2 F 0 req
ð8:45Þ
which is called the Prandtl-Reuss equation. The plastic work rate of this material is described as
p
r:e ¼ r:k
r0 p eqp eqp ¼ k jjr0 jj ¼ jjr0 jjjj e jj ¼ req e ¼ Fðeeqp Þ e 0 jjr jj
ð8:46Þ
which is the product of the hardening function and the rate of the equivalent plastic strain and thus the hardening attributable to the plastic work is called the work hardening, too. The traction t acting on the octahedral plane is expressed by 1 1 1 t ¼ r em ¼ ðr1 e1 e1 þ r2 e2 e2 þ r3 e3 e3 Þ pffiffiffi e1 þ pffiffiffi e2 þ pffiffiffi e3 3 3 3 1 ¼ pffiffiffi ðr1 e1 þ r2 e2 þ r3 e3 Þ ð8:47Þ 3
8.4 Constitutive Equation of Metals
223
on the principal base, noting Eq. (5.2) with n ¼ em and Eq. (1.135). Then, the normal component roct and the tangential component s oct of the traction t are given as 1 1 1 1 roct t em ¼ pffiffiffi ðr1 e1 þ r2 e2 þ r3 e3 Þ pffiffiffi e1 þ pffiffiffi e2 þ pffiffiffi e3 3 3 3 3 1 ¼ ðr1 þ r2 þ r3 Þ ¼ rm 3 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi soct jjt jj2 r2oct rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 1 ðr þ r22 þ r23 Þ ðr1 þ r2 þ r3 Þ2 ¼ 3 1 9 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 ¼ fr þ r22 þ r23 ðr1 r2 þ r2 r3 þ r3 r1 Þg 9 1 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rffiffiffipffiffiffiffiffi pffiffiffi 2 eq 1 2 r ðr1 r2 Þ2 þ ðr2 r3 Þ2 þ ðr3 r1 Þ2 ¼ J2 ¼ ¼ 3 3 3
ð8:48Þ
ð8:49Þ
where soct is called the octahedral shear stress and J2 is given by choosing the deviatoric Cauchy tensor r0 as the second invariant of the deviatoric tensor T0 in Eq. (1.182) as follows: 1 1 1 1 2 J2 kr0 k ¼ tr r0 2 ¼ r0rs r0sr ¼ req 2 2 2 2 3 1 02 0 2 0 2 0 2 ¼ ðr11 þ r22 þ r33 Þ þ r12 þ r0232 þ r0312 2 1 1 ¼ ðr012 þ r022 þ r032 Þ ¼ fðr1 r2 Þ2 þ ðr2 r3 Þ2 þ ðr3 r1 Þ2 g 2 6
ð8:50Þ
It is interpreted from Eqs. (8.49) and (8.50) for the Mises yield condition that the yielding is induced when the octahedral shear stress reaches a certain value. Equations (8.48) and (8.49) are also derived as the components in the directions pffiffiffi Im ð¼ 3em Þ and t0 ð¼ T0 =jjT0 jjÞ, i.e. Tm and jjT0 jj of Tm and T0 , respectively, regarding T as r in Eqs. (1.314) and (1.318).
8.5
Formulation of General Loading Criterion
The judgment of whether or not the plastic strain rate is induced is required for the elastoplastic deformation analysis. The criterion for this judgment is called the loading criterion. In what follows, it is formulated (Hashiguchi 2000), which is not limited to the associated flow rule but applicable even to the non-associated flow rule described at the end of Sect. 8.3.
224
8
Elastoplastic Constitutive Equations
1. In the loading (plastic deformation) process, it is required noting Eqs. (8.29) and (8.39) that 9 > n :r > = k ¼ p [0 p M for e 6¼ O ð8:51Þ > n : E :e > ; [0 K¼ p M þn : E : m
p
2. In the unloading (elastic deformation) process e ¼ O, it holds that
n :r 0
ð8:52Þ
n :r [ 0 for M p \0 Mp
ð8:53Þ
and thus
k¼
e
e
On the other hand, substituting e ¼ e leading to n : E : e ¼ n : E : e ¼ n : r into
Eq. (8.31), it holds for K that
n :r \0 ð8:54Þ Mp þ n : E : m in this process. Here, note that the elastic modulus tensor E is the positive definite tensor fulfilling n : E : n [ 0 for an arbitrary n as defined in Eq. (1.264) for the second-order tensor, noting (ε e) ε e: E: ε e 0. Here, we may assume
K¼
n : E : m jM p j, provided that m is not far different from n. Eventually, it follows that M p þ n : E : m [ 0:
ð8:55Þ
Also, note that the infinite rate of plastic relaxation is induced so that the stress decreases in the infinite rate if the denominator becomes zero, i.e. M p þ n : E : m ! 0 as known from Eq. (8.33) illustrated in Fig. 8.2 for the unip e axial loading process. Then, in the unloading process ðe ¼ O; e 6¼ OÞ, the following inequalities hold from Eqs. (8.53) and (8.54), depending on the sign of the plastic modulus M p leading to the hardening, perfectly-plastic and softening state. 8 p > > < k \0 and K \0 for M [ 0 k ! Indeterminate and K \0 for M p ¼ 0 > > : k [ 0 and K \0 for M p \0
ð8:56Þ
8.5 Formulation of Loading Criterion
225
Fig. 8.2 Signs of k and K in uniaxial loading process ðr e ¼ E : Þ
Therefore, the sign of k at the moment of unloading from the state M p 0 can be
positive or indeterminate in the unloading process. On the other hand, K is definitely negative in the unloading process. Thus, the distinction between a loading
and an unloading process cannot be judged by the sign of k but it can be done by
that of K. Consequently, the loading criterion is given as either
p 6 O for f ðrÞ ¼ FðHÞ and K [ 0 e ¼ p e ¼ O for others
ð8:57Þ
or
p
e ¼ 6 O for f ðrÞ ¼ FðHÞ and n : E : e [ 0 p e ¼ O for others
ð8:58Þ
in lieu of Eq. (8.55). Limiting to the hardening process with M p [ 0, Eq. (8.58) leads to
p 6 O for f ðrÞ ¼ FðHÞ and n : r [ 0 e ¼ p e ¼ O for others
ð8:59Þ
However, the loading criterion in Eq. (8.59) is inapplicable to the softening state observed in loose soils, rocks, concrete and damage materials, etc. The plastic loading is interpreted as the process that the plastic relaxation stress p rate r is induced, noting Eq. (8.34).
226
8
Elastoplastic Constitutive Equations
dσ e (n: d σ e = n: : dε > 0): loading
n
σ dσ e (n : d σ e = n : : unloading
0
: dε < 0)
dσ e (n : d σ e = n : : dε = 0) : neutral loading ij
Yield surface f ( : (ε ε p )) = F ( H ) Fig. 8.3 Loading criterion by the direction of the elastic stress increment dre E : de. in strain space
The loading criterion in Eq. (8.58) is rewritten by the elastic strain rate r E : e in Eq. (8.34) as
p e e ¼ 6 O for f ðrÞ ¼ FðHÞ and n : r [ 0 p e ¼ O for others
ð8:60Þ
Equation (8.60) is interpreted as follows: The loading, the neutral loading and the
unloading are defined as processes that the elastic stress rate re is directed outward, tangential and inward direction, respectively, of the yield surface as shown in Fig. 8.3. In elastoplastic deformation analysis, suppose to calculate first r and e by either of the elastic or the elastoplastic constitutive equation. Then, check the sign of n : E : e. If the sign conflicts with the loading criterion, it is required to recalculate them using another constitutive equation. Here, it would be efficient to calculate first by the elastoplastic constitutive equation since the monotonic loading process in which the elastoplastic deformation process continues is seen often in practical engineering problems. The loading judgment by the direction of strain rate, i.e. the sign of n : E : e (or n : r for the hardening state) is not required for the numerical calculation exploiting the return-mapping projection for hyperelastic-based constitutive equations.
8.6 Physical Backgrounds of Associated Flow Rule
8.6
227
Physical Backgrounds of Associated Flow Rule
The associated flow rule holds for a wide range of materials, which insists “The direction of the plastic strain rate is the outward-normal of the yield surface in the coordinate space where the each directions of the plastic strain rate components p eij are taken same to the each directions of the stress components rij ”. The adjective “associated” is added because the flow rule is associated to the yield surface. On the other hand, the flow rule which is not associated to the yield surface is called the non-associated flow rule which is observed in highly frictional (pressure-dependent yielding) materials. Several physical interpretations for the associated flow rule are described in this section.
8.6.1
Positiveness of Second-Order Plastic Work Rate: Prager’s Interpretation
Prager (1949) reported that the associated flow rule must hold to fulfill the positivity of the second-order plastic work rate, i.e.
p
r:e 0
ð8:61Þ
The direction of the plastic strain rate must be outward-normal to the yield surface for any stress rate directing outwards the yield surface, assuming that the direction of plastic strain rate depends on the state of stress but independent of the rate of stress as shown in Fig. 8.4. However, Eq. (8.61) holds only for hardening materials in which the stress rate is directed outwards the yield surface but does not hold for softening materials. εp
σ
σy
0
ij ,
Yield surface
Fig. 8.4 Prager’s positive plastic work rate for hardening materials
p
ij
228
8.6.2
8
Elastoplastic Constitutive Equations
Principle of Maximum Plastic Work
The postulate, called the principle of maximum plastic work, was proposed by Mises for rigid-plastic materials and Hill (1948b, 1950) and Mandel (1964) for elastoplastic materials. It insists that the plastic work rate done by the actual stress ry on the yield surface is greater than a plastic work rate done by any mechanically-admissible stress r inside the convex yield surface, leading to
ry : e [ r : e ; i:e:ðry r Þ: e [ 0 p
p
p
ð8:62Þ
as depicted in Fig. 8.5. Then, the plastic strain rate must be directed to the outward-normal of convex yield surface, so that the associated flow rule must hold in order to satisfy the principle of maximum plastic work under the assumption that p the direction of plastic strain rate e depends only on the state of stress but it is independent of the rate of stress. This fact was discussed for the finite strain theory by Lubliner (1984).
8.6.3
Positiveness of Work Done During Stress Cycle: Drucker’s Interpretation
Drucker (1951) proposed the background as will be described in the following. Drucker (1951) postulated “the work done during the stress cycle by the external agency is positive”. It is described mathematically as follows: Z
tðr0 Þ t0 ðr0 Þ
ðr r0 Þ : e dt 0
ð8:63Þ
εp
σy
σy σ*
0
σ* ij ,
Yield surface
Fig. 8.5 Principle of maximum plastic work
p
ij
8.6 Physical Backgrounds of Associated Flow Rule
229
y
(σ y σ 0 ) : ε pdt
0
d
0
p
0 Fig. 8.6 Positive work done by external agency in Drucker’s (1951) postulate (Illustration in uniaxial loading process)
where r0 stands for the initial stress at the initial time t0 ðr0 Þ, and tðr0 Þ designates the time that the stress returns to the initial stress. The following inequality is obtained from Eq. (8.63) under the assumption that the inside of the yield surface is the purely-elastic domain (see Fig. 8.6).
p
ðry r0 Þ: e 0
ð8:64Þ
where ry designates the stress on the yield surface in which the plastic strain rate e is induced. The followings must hold in order to fulfill Eq. (8.64).
p
(1) The plastic strain rate is directed outward-normal of the yield surface in the coordinate system where the components of the stress and the corresponding components of the plastic strain rate are taken to the same directions. Then, the associated flow rule must hold, provided that the direction of plastic strain rate is determined solely by the current stress and internal variable but independent of stress rate. (2) In this occasion the yield surface has to be the convex surface (see Fig. 8.7). The result (1) is called the associated flow rule or the normality rule and the result (2) is called the convexity of yield surface.
8.6.4
Positiveness of Second-Order Plastic Relaxation Work Rate
Ilyushin (1961) postulated that “the work done during the strain cycle is positive”. Limiting to the infinitesimal deformation process. It leads to the postulate “the second-order work increment d 2 w is not larger than the second-order elastic stress
230
8
εp
Yield surface
σy
σ0 σ0
σ0
σ0
εp
σ0
σy
Elastic region
0
Elastoplastic Constitutive Equations
σ0
(σ y σ 0 ) : ε p 0 ij ,
p
ij
Concave yield surface: Violation of Drucker (1951)’s postulate.
Normality rule Fig. 8.7 Associated flow (normality) rule and convexity of yield surface based on the Drucker’s (1951) postulate
work increment d 2 wes calculated by presuming that the strain increment is induced elastically” or “the second-order plastic relaxation work increment d 2 wpr , i.e. the work increment done during the infinitesimal strain cycle is not negative” (Hill 1968; Petryk 1991, 1997; Hashiguchi 1993a). The second-order work increment d 2 w is shown by Dabf, the second-order elastic stress work increment d 2 wes by Dadf and the second-order plastic relaxation work increment d 2 wpr by Dabc noting Dadb ¼ Dabc in Fig. 8.8. It is described mathematically as follows: d 2 w ¼ d 2 wes d 2 wpr ; d 2 w d 2 wes ; d 2 wpr 0
ð8:65Þ
where 9 1 1 > d 2 w dr : de ¼ dee : E : de > > > 2 2 = 1 e 1 2 es d w dr : de ¼ de : E : de > 2 2 > > > 1 p 1 p 1 2 pr d w dr : de ¼ de : E : de ¼ k dtm : E : de ; 2 2 2
ð8:66Þ
with dre E : de; drp E : dep
ð8:67Þ
dre and drp are called the elastic stress increment and the plastic relaxation stress e p increment, respectively, following r and r in Eq. (8.34).
8.6 Physical Backgrounds of Associated Flow Rule
231
d
d d
p
E
e
1
d
a
d 2wpr
e
b f
1
E
c
0
d
p
d
e
d
(a) Hardening process ( M p
0) d
d
e
d
E
p
1
a
f e
d
b
d 2wpr 1 E
c
0
d
d d
(b) Softening process ( M p
e
p
0)
Fig. 8.8 Positiveness of work done during strain cycle, i.e. second-order plastic work rate ðde ¼ dee þ dep Þ (Illustration in uniaxial loading)
232
8
Elastoplastic Constitutive Equations
It should be noted that the associated flow rule with m ¼ n in Eq. (8.38) must hold in order that Eq. (8.65) conforms to the loading condition in Eq. (8.58), 3
although Eq. (8.58) was not restricted to the associated flow rule. The above-mentioned positiveness of second-order plastic relaxation work rate is violated in the non-associated flow which causes i) the physically unacceptable deformation in which the tangent stiffness modulus is larger than the elastic modulus and ii) the asymmetry of the tangent stiffness modulus tensor, iii) the difficulty in the formulation of the variational principle and iv) the necessity for the incorporation of additional material constant(s) as revealed by Hashiguchi (1991).
8.6.5
Comparison of Interpretations for Associated Flow Rule
Prager’s (1949) interpretation of the associated flow rule is concerned only with hardening materials as described previously. Hill’s (1948b, 1950) principle of maximum plastic work is not limited to the hardening behavior but does not provide a clear physical background. On the other hand, the interpretations of the positivity of work done by the external agency, i.e. the additional stress during a stress cycle by Drucker (1951) and of the positivity of the second-order plastic relaxation work rate during a strain cycle by Ilyushin (1961) are based on the conceivable physical postulates of the dissipation energy of materials unlimited to the hardening behavior. Here, Drucker’s (1951) postulate is related to the stress cycle but Ilyushin’s (1961) postulate of the second-order plastic relaxation work rate is related with the infinitesimal strain cycle. Now, compare below the pertinence of these postulates. (1) The strain cycle can be realized always. However, the stress cycle cannot be made in a softening state in which the stress cannot be returned to the initial state if the plastic strain rate is induced. It is based on the fact that any deformation can be given but a stress cannot be given arbitrarily to materials since strength of materials is limited. (2) Limiting to the infinitesimal cycles, consider the stress and the strain cycles. The second-order work increment done during the infinitesimal stress cycle is given by ð1=2Þdr : dep (Dabe in Fig. 8.8a). On the other hand, the additional work increment ð1=2Þdep : E : dep (Daec in Fig. 8.8a) must be done to close the strain cycle, whilst ð1=2Þdep : E : dep 0 holds because of the positive-definiteness of the elastic modulus tensor E. Therefore, the work done during the infinitesimal stress cycle is far smaller than the work during the infinitesimal strain cycle. In other words, Drucker’s (1951) postulate holds for the materials fulfilling a more restricted condition, i.e., more particular materials than the materials fulfilling the positivity of the second-order plastic relaxation work rate. (3) Strain (increment) in any definition is determined uniquely by the displacement (increment) induced in the material. Therefore, the configuration of material
8.6 Physical Backgrounds of Associated Flow Rule
233
returns to the initial configuration only if the strain returns to the initial value. In other words, if a cycle of strain in a certain definition closes, cycle of strain in any other definition (Lagrangian, Almansi, logarithmic, nominal and infinitesimal strains for instance) also closes. On the other hand, the stress is defined by the force per unit area and thus it is influenced by the deformation. The configuration in the end of stress cycle differs from the initial configuration depending on the loading path chosen during the cycle and on the definition of stress (Cauchy, Kirchhoff, nominal and second Piola-Kirchhoff stresses for instance). Then, even if a cycle of stress in a certain definition closes, the cycle in the stress in the other definition does not close. Eventually, the positivity of the strain cycle possesses the objectivity, but that of the stress cycle is not objective. (4) The assumption that the interior of the yield surface is the purely elastic domain is adopted in Drucker’s interpretation. On the other hand, this assumption is not required by the postulate of the positivity of second-order plastic relaxation work rate, which holds on the quite natural premise that the purely elastic deformation is induced at the moment of reverse loading. Eventually, it can be stated that postulate of the positivity of second-order plastic relaxation work rate is more general than the Drucker’s postulate. However, even the former is based on the premise that the direction of the plastic strain rate is dependent on the normal component but independent of the tangential component of stress rate to the yield surface. It is observed in the test result that the inelastic deformation is induced even by the tangential component. The inelastic strain rate induced by the component of stress rate tangential to yield surface will be described in Sect. 9.7.
8.7
Anisotropy
The plastic strain rate described in Sect. 8.3 concerns the yield condition with the function of stress invariants and scalar-valued internal variables. Therefore, it is limited to the materials exhibiting the isotropic responses in the plastic deformation behavior. In what follows, first the isotropy in constitutive equation is defined. Then, the plastic strain rate extended to the anisotropy will be explained in this section.
8.7.1
Definition of Isotropy
An isotropic material is defined as one exhibiting identical mechanical response that is independent of the chosen direction of material element or of the coordinate
234
8
Elastoplastic Constitutive Equations
system by which the response is observed. Here, the input/output variables are the stress rate and the strain rate in the irreversible deformation. The rate-type constitutive equation is described in general as follows:
fðr; r; Hi ; eÞ ¼ O
ð8:68Þ
where Hi ði ¼ 1; 2; 3; Þ denotes collectively scalar-valued or tensor-valued internal state variables. When the following equation holds by giving coordinate transformations only for stress (rate) and strain rate tensors in the tensor-valued function f, it can be stated that Eq. (8.68) describes the constitutive equation of isotropic material.
fðQ r QT ; Q r QT ; Hi ; Q e QT Þ ¼ Qfðr; r; Hi ; eÞQT
ð8:69Þ
In the plastic constitutive equation formulated incorporating the yield and/or plastic potential function, the isotropy holds if the yield and/or plastic potential function is given by the function of stress invariants and scalar internal variables. Then, designating these scalar-valued functions by f , it must fulfill the following equation. f ðQ r QT ; Hi Þ ¼ f ðr; Hi Þ
ð8:70Þ
In contrast, the anisotropic plastic constitutive equation is described by incorporating the yield and/or plastic potential function including tensor-valued internal variable in addition to the stress invariants and scalar internal variables. Then, Eqs. (8.69) and (8.70) do not hold in anisotropic constitutive equations.
8.7.2
Elastoplastic Constitutive Equation with Kinematic Hardening
It is well-known through the experimental observation that if the monotonic loading proceeds towards a certain direction in the stress space, the hardening develops in that direction but the yield stress lowers in the opposite direction. From the microscopic viewpoint, this phenomenon is induced by the staticallyindeterminable deformation of internal structure and is called the Bauschinger effect. To reflect this effect in the elastoplastic constitutive equation, the translation or the rotation of the yield surface is adopted widely. The translation of the yield surface, called the kinematic hardening, is realized by introducing the back-stress translating towards the loading direction and replacing the stress tensor with the relative tensor given by subtracting the back-stress tensor from the stress tensor. On the other hand, soils, which is the assembly of particles with weak adhesion among them, can bear a far larger compression stress than the tensile stress. Therefore, they exhibit a strong frictional property that the deviatoric yield stress increases with the pressure, while the yield surface only slightly includes the tensile stress regime near
8.7 Anisotropy
235
the origin of the stress space. Therefore, once the yield surface translates leaving the origin, it can never come back to include the origin again because the yield surface contracts with the plastic volume expansion leading to the softening. Therefore, the kinematic hardening cannot be applied but the rotation of the yield surface around the origin of stress space, i.e. the rotational hardening, is suitable to soils as will be described in Chap. 13. Now, let the yield condition (8.16) be extended to describe the anisotropy by introducing the internal variables of second-order tensor a as follows: f ð^ rÞ ¼ FðHÞ
ð8:71Þ
^ ra r
ð8:72Þ
where
að¼ a0 Þ being the back-stress (kinematic hardening variable) proposed by Prager (1956) in order to describe the induced anisotropy, the evolution rule of which depends on stress and internal variables as will be described in the next section. Note that an anisotropic hardening variable is deviatoric in general. Here, it is ^ in the homogeneous degree-one assumed that f in Eq. (8.71) is the function of r fulfilling f ðjsj^ rÞ ¼ jsjf ð^ rÞ. Then, the yield surface (8.71) maintains the similar shape and orientation with respect to the stress point r ¼ a. The material time-derivative of Eq. (8.71) leads to the consistency condition. @f ð^ rÞ @f ð^ rÞ 0 :r : a F H ¼ 0 @^ r @^ r
ð8:73Þ
)
p p
p p F; e = e e
H ¼ fHe ðr; F; e Þ ¼ fH^n r;
p p
p p a ¼ f ke ðr; F; a; e Þ ¼ f ke r; F; a; e = e e
ð8:74Þ
where
p
noting that they are homogeneous functions of e in degree-one since they are p induced only in the plastic loading process e 6¼ O and the first-order time-differential quantities. The associated flow rule in Eq. (8.24) is extended as
p p ^ ðk [ 0; jj^ njj ¼ 1; e ¼ kÞ e ¼kn
ð8:75Þ
@f ð^ rÞ
@f ð^ rÞ
^ = n @r @r
ð8:76Þ
with
236
8
Elastoplastic Constitutive Equations
noting @f ðb r Þ=@r ¼ @f ðb r Þ=@ b r Substituting 9 @f ðb rÞ > > :b r ¼ f ðb rÞ ¼ F > = @b r @f ðb r Þ
:b r @f ðb
@f ðb
n :b > r Þ
r> > @r
n
r Þ ¼ ^ ; ^ = ¼ 1=
@r
f ðb rÞ @r F
ð8:77Þ
based on the Euler’s homogeneous function of tensor variable (see Appendix D), Eq. (8.73) leads to
^:r ^ 0 n F H¼0 F
^n : r ^n : a
The substitution of the associated flow rule in Eq. (8.75) into this equation leads to
^n : r k M p ¼ 0
ð8:78Þ
where
F0 fH^n b M ^n : r þ f k^n F
ð8:79Þ
p
^ Þ ¼ H = k; fH^n ðr; H; n
^Þ ¼ a = k f k^n ðr; F; a; n
ð8:80Þ
It follows from Eq. (8.78) that
^ :r n ; Mp
k¼
p
e ¼
^ :r n ^ n Mp
ð8:81Þ
and
e ¼ E1 : r þ
^ :r n ^¼ n Mp
^n ^ n :r E1 þ Mp
ð8:82Þ
from which it follows that
^ : E :e n ^ : E :n ^ Mp þ n ^n ^:E ^ : E :e E :n n ^ ¼ E p E :n :e r ¼ E :e p ^ : E :n ^ ^ : E :n ^ M þn M þn
K¼
ð8:83Þ ð8:84Þ
8.7 Anisotropy
237
The loading criterion is given as follows (Hashiguchi 1994, 2000)
p
)
^ : E :e [0 6 O for f ð^ rÞ ¼ FðHÞ and K [ 0 or n e ¼ p e ¼ O for others
8.7.3
ð8:85Þ
Kinematic Hardening Rules
The evolution rule of the back-stress for the plastically-incompressible metals was given by Prager (1956) as follows (see Fig. 8.9):
a ¼ ck e
p
ð8:86Þ
where ck is the material constant with the dimension of stress. Equation (8.86) is described in the uniaxial loading process by
p
aa ¼ c k e a
ð8:87Þ
where ð Þa designates the axial component. If ck is extended to be a monotonic decreasing function of the equivalent plastic strain, the nonlinear behavior is described in the initial monotonic loading process but the peculiar behavior is described in the inverse loading process as shown in Fig. 8.9b. It should be noted that this extension is not accepted, resulting in the worsening, although it has been recommended in some literatures (e.g. de Souza Neto et al. 2008). Shield and Ziegler (1958) and Ziegler (1959) proposed the following modification of Prager (1956)’s linear kinematic hardening rule in Eq. (8.86), insisting that Preger’s rule possesses the mathematical inconvenience such that the components of back-stress tensor are induced also in the directions in which the components of stress tensor are zero.
a
a
ck
1 0
(a) ck : constant
1
p
a
0
ck
p
a
(b) ck : monotonic-decreasing function of plastic strain
Fig. 8.9 Rate-linear kinematic hardening rules illustrated in one-dimensional state
238
8
Elastoplastic Constitutive Equations
0 p
^ e
a ¼ cz r
ð8:88Þ
b 0i ¼ 0 which is seen in a where cz is the material constant, noting that ai ¼ 0 for r plane stress condition. The directions of the translations of the back-stresses in Eqs. (8.86) and (8.88) are identical to each other for the Mises yield surface whose 0 0 ^¼b r jj holds, while section cut by the deviatoric plane is circle for which n r =jjb they are obviously different from each other for the Tresca yield surface for instance. It would be merely the disturbance in the history of plasticity, although it has been introduced in a lot of literatures (e.g. Chakrabarty 1987; Duszek and Perzyna 1991; Khan and Huang 1995; Lubarda 2002; Asaro and Lubarda 2006). The following nonlinear-kinematic hardening rule of metals was formulated by Hashiguchi (2000). 1 p 1 p ^ ke ka ¼ k f k^n ; f k^n ¼ ck n a a ¼ ck e bk F bk F
ð8:89Þ
pffiffiffiffiffiffiffiffi where ck and bk ð 3=2Þ are the material constants, while the first and the second terms in Eq. (8.89) are called the hardening part and the dynamic recovery part, respectively. On the other hand, the original Armstrong-Frederick nonlinear kinematic hardening rule is given as 1 p p a ¼ ck e ke ka bk
from which Eq. (8.89) is modified as the saturation value of the movement of the yield surface depend on the size of the yield surface F by the modifications bk ! bk F . The pertinence ofEq. (8.89) for the evolution rule of the kinematic hardening can be perceived by the physical interpretation that the saturation of the kinematic hardening would decrease with the isotropic softening and further will be confirmed in Subsection 15.5.1 as the saturation of the kinematic hardening decreases with the damage evolution. This pertinent property is not provided in the Armstrong-Frederick rule. Now, Eq. (8.89) is reduced to the following equation for the uniaxial loading process of the Mises metals exhibiting the plastically-incompressibility, i.e. p p el ¼ ea =2. ! rffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 3 1 p p2 p2 p e a þ 2 e l aa ¼ c k 1 aa e a aa ¼ c k e a bk F 2 bk F p p : ea [ 0; þ : ea \0
leading to
ð8:90Þ
8.7 Anisotropy
239 a
2b F 3 k
ck (1
Limit kinematic hardening surface
3 a 2 bk F
)
1
ck
0
ck (1
3 a 2 bk F
1
) ck 1
1
1
2ck
1
p
a
2ck
Limit kinematic hardening surface
2b F 3 k
Fig. 8.10 Nonlinear kinematic hardening rule illustrated for uniaxial loading process
) 8 p pffiffiffiffiffiffiffiffi 0 for ea [ 0 > > > in aa ¼ 2=3bk F > p > > > < 2ck for ea \0 p aa = ea ¼ ck in aa ¼ 0 > ) > p > > pffiffiffiffiffiffiffiffi 2ck for ea [ 0 > > > in aa ¼ 2=3bk F : p 0 for ea \0
p
ð8:91Þ
p
noting el ¼ ea =2, where ðÞl designates the lateral component. Equation (8.90) is
time-integrated under the isotropic nonhardening state F ¼ 0 as follows:
ð8:92Þ
where A is the saturation value of aa , i.e. A
pffiffiffiffiffiffiffiffi 2=3bk F
ð8:93Þ
The relation of the axial components aa of the back-stress versus the axial plastic strain epa in the uniaxial loading is illustrated in Fig. 8.10 in which the kinematic hardening gradually saturates in the monotonic loading process but the abrupt increase of the kinematic hardening rate occurs at the moment of reverse loading. pffiffiffiffiffiffiffiffi The kinematic hardening saturates at 2=3bk F. The Armstrong-Frederick kinematic hardening rule have been worsened to various complex forms by Chaboche (1991), Hassan and Kyriakides (1991), Ohno and Wang (1993), Dafalias et al. (2008), Hassan et al. (2008), Dafalias and Feigenbaum (2011), etc. Here, it should be noted that the kinetic hardening is important for plastically-incompressible metals but it is irrelevant to the plastically-incompressible
240
8
Elastoplastic Constitutive Equations
materials, e.g. soils, rocks, concrete, ceramics, etc. for which the rotational hardening is dominant as will be described in Chap. 13. Further, the directional distortional hardening has been studied hitherto, which leads to the distortion of the shape of the yield surface such that a region of high curvature develops in the direction of loading and flattening develops in the opposite direction by Kurtyka and Zyczkowski (1996), Voyiadjis and Forooze (1990), Feigenbaum and Dafsalias (2007, 2014), Shi et al. (2017), etc.
8.8
Plastic Spin
The infinitesimal hyperelastic-based plastic constitutive equation is formulated so far. In order to transform it to the hypoelastic-based plastic constitutive equation, the plastic spin in Eq. (3.44) must be incorporated. It is given specifically by the following equation in terms of the plastic-work conjugate pair ðs; dp Þ as would be suggested from Eq. (5.58) in Sect. 5.58 and shown specifically in Sect. 17.6. wp ¼ gp ant½s dp
ð8:94Þ
where gp is the material parameter. Equation (8.94) is based on the non-coaxiality between the stress and the plastic strain rate, which is generally accompanied with the anisotropy. Obviously, the plastic spin is not induced in isotropic materials
^Þ are co-axial and thus they are fullfilling a ¼ O holds, for which s and dp ð¼ kn commutable. Equation (8.94) is the modification of the proposition wp ¼ gp ant½r dp by Zbib and Aifantis (1988) in which the work-conjugacy is ignored. It will be extended to the multiplica-tive hyperelastic-based plasticity in Sect. 17.9.
8.9
Physical Interpretation of Nonlinear Kinematic Hardening Rule
The strain e is additively decomposed into the elastic strain ee as the storage part and the plastic strain ep as the pseudo dissipative part, and further the plastic strain ep is additively decomposed into the storage part epks which induces the kinematic hardening and the dissipative part epkd , i.e. ep ¼ epks þ epkd
ð8:95Þ
The stress r and the kinematic hardening variable a which describe the variation of substructures in material are given by the partial derivatives of the potential functions of the elastic strain and the storage parts of the plastic strain so that the following relations hold.
8.9 Physical Interpretation of Nonlinear Kinematic Hardening Rule
r¼
@we ðee Þ ; @ee
a¼
@wk ðepks Þ @epks
241
ð8:96Þ
Now, assuming 1 1 we ðee Þ ¼ ee : E : ee ; wk ðepks Þ ¼ ck epks : epks 2 2
ð8:97Þ
it follows from Eq. (8.96) that a ¼ ck epks ¼ ck ðep epkd Þ
r ¼ E : ee ¼ E :ðeep Þ;
ð8:98Þ
The time-differentiation of Eq. (8.98) reads:
p
p
p
r ¼ E :ðe e Þ; a ¼ ck ðe ekd Þ
ð8:99Þ
Then, the rates of the storage and the dissipative parts for the stress and the kinematic hardening are given as Storage part ⎧ 644 7448 ⎪ p • • ⎪σ = : ( ε − ), ε• { ⎪ Dissipative part ⎪ Storage part ⎨ 6447448 ⎪ ⎪α• = ck ( ε• p − 1 || ε• p||α ) bk F ⎪ 14243 ⎪ Dissipative part ⎩
ð8:100Þ
from Eqs. (8.5) and (8.89). The dissipative parts of the strain and the plastic strain satisfy the positivity of the dissipation energy, i.e. • ⎧ •p ⎪σ : ε = (σ : nˆ )λ ≥ 0 ⎨ • p ⎪ α : ε• kd = (α : α )λ / (b k F ) ≥ 0 ⎩
8.10
ð8:101Þ
Limitations of Conventional Elastoplasticity
The conventional elastoplasticity described in this chapter is premised on the assumption that the interior of yield surface is a purely elastic domain. Therefore, the relation of stress rate versus strain rate is predicted to change abruptly at the moment when the stress reaches the yield surface. Therefore, the smooth stress– strain curve observed in real materials is not predicted as shown in Fig. 8.11. This results in the defects: a determination of offset value, i.e. plastic strain at yield point
242
8
Elastoplastic Constitutive Equations
Experiment Elastic state Elastoplastic state
Prediction by conventional plasticity Unrealistic prediction: Excessively high peak stress
Roughly real prediction
0
0 Hardening
Softening
Fig. 8.11 Prediction of monotonic loading behavior by conventional plasticity
is required, which is influenced by an arbitrariness since a plastic deformation develops gradually and thus a stress–strain curve is smooth usually in real materials, and a peak stress value is predicted to be excessively high in a softening behavior observed in frictional materials, e.g. soils and rocks and concrete. Further, only an elastic deformation is predicted for the cyclic loading of stress below the yield stress. In real materials, however, plastic deformation is accumulated even for stress cycles lower than the yield stress and the strain is accumulated resulting in the failure as will be described in subsequent chapters. Thus, the conventional plasticity possesses the above-mentioned various defects in the application to the mechanical design of machines and structures in engineering practice.
Chapter 9
Unconventional Elastoplasticity Model: Subloading Surface Model
Elastoplastic constitutive equations with the yield surface enclosing the elastic domain possess many limitations in the description of elastoplastic deformation, as explained in the last chapter. They are called the conventional model in the Drucker’s (1988) classification of plasticity models. Various unconventional elastoplasticity models have been proposed, which are intended to describe the plastic strain rate induced by the rate of stress inside the yield surface. Among them, the subloading surface model is the only pertinent model fulfilling the mechanical requirements for unconventional models. These mechanical requirements are first described and then the subloading surface model is explained in detail.
9.1
Mechanical Requirements
There exist various mechanical requirements, e.g., the thermodynamic restriction and the objectivity for constitutive equations. Among them, the continuity and the smoothness conditions are violated in many elastoplasticity models, while their importance for formulation of constitutive equations has not been sufficiently recognized to date. Before formulation of the plastic strain rate, these conditions will be explained below (Hashiguchi 1993a, b, 1997, 2000).
9.1.1
Continuity Condition in the Small
It is observed in experiments that “stress rate changes continuously for a continuous change of strain rate”. This fact is called the continuity condition in the small and is expressed mathematically as follows (Hashiguchi 1993a, b, 1997, 2000).
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_9
243
244
9
Unconventional Elastoplasticity Model: Subloading Surface Model
lim rðr; Hi ; e þ d eÞ rðr; Hi ; eÞ ! O
ð9:1Þ
d e!O
where Hi ði ¼ 1; 2; 3; Þ denotes collectively scalar-valued or tensor-valued internal state variables. In addition, dðÞ stands for an infinitesimal variation. The response of the stress rate to the input of strain rate in the current state of stress and internal variables is designated by rðr; Hi ; eÞ. Uniqueness of solution is not guaranteed in constitutive equations violating the continuity condition, predicting different stresses or deformations for identical input loading. The violation of this condition is schematically shown in Fig. 9.1. Ordinary elastoplastic constitutive equations in which the plastic strain rate is derived obeying the consistency condition, fulfill the continuity condition. As described later, however, no elastoplastic constitutive equation fulfills it except for the subloading surface model when they are extended to describe the tangential inelastic strain rate. The concept of the continuity condition was first advocated by Prager (1949). However, a mathematical expression of this condition was not given. The condition was defined as the continuity of strain rate to the input of stress rate by Prager (1949) inversely to the definition given above. However, identical stress rate directing inwards the yield surface can induce different strain rates in loading and unloading states for softening materials. Also, identical stress rate along the yield surface can induce different strain rates in a perfectly-plastic material as illustrated in Fig. 9.2. Besides, it is noteworthy that a stress rate cannot be given arbitrarily since there exists a limitation in strength of materials although a strain rate can be given arbitrarily. For that reason, the Prager’s (1949) notion does not hold in the general loading state including softening and the perfectly plastic states.
jump
jump
Input : ε•
Fig. 9.1 Violation of continuity condition in the small
Response : σ•
9.1 Mechanical Requirements
245
σ Input: dσ ( > < [ 0 for R\1 R ¼ 0 for R ¼ 1 > > : ð\0 for R [ 1Þ e ¼ 0 for e ¼O R 6 O \0 for ee ¼
for ep 6¼ O
for ep ¼ O
ð9:7Þ
ð9:8Þ
Here, the rate of the normal-yield ratio evolves with the plastic strain rate, obeying Eq. (9.7). On the other hand, when only the elastic strain rate is induced, the updated value of R needs to be calculated by solving Eq. (9.6) for R. Note that in the calculation, the stress is updated obeying the elastic constitutive equation and internal variables are fixed since no plastic strain evolves during this process. Then, it follows that
R ¼ UðRÞjj ep jj ¼ UðRÞ k for ep 6¼ O
Fig. 9.5 Plastic strain rate based on the subloading surface concept
ð9:9Þ
9.2 Subloading Surface (Hashiguchi) Model
R¼
f ðrÞ for ee 6¼ O and ep ¼ O F
249
ð9:10Þ
where UðRÞ is the monotonically-decreasing function of R fulfilling the conditions (see Fig. 9.6a). 8 ! þ 1 for R ¼ 0 ðelastic stateÞ > > < [ 0 for R\1 ðsubyield stateÞ UðRÞ ¼ 0 for R ¼ 1 ðnormal-yield stateÞ > > : \0 for R [ 1 ðover normal-yield stateÞ
ð9:11Þ
The explicit form of the function UðRÞ satisfying Eq. (9.11) is given by p UðRÞ ¼ u cot R 2
ð9:12Þ
where u is the dimensionless material parameter regulating the increase of the normal-yield ratio R for a certain plastic strain increment, noting dR=jjdep jj ¼ UðRÞ ¼ u cot½ðp=2ÞR. Therefore, the stress vs. strain curve is smoother for a smaller value of the parameter u, while the conventional elastoplastic behavior is realized for u ! 1. However, note that there exist a lot of materials including usual metals in which the plastic strain rate is hardly induced blow a certain value of the normal-yield ratio. Then, let Eq. (9.11) be modified to the following equation, by which the plastic strain rate is not induced until the normal-yield ratio R reaches a certain value of the material parameter Re ð\1Þ (see Fig. 9.6b).
Fig. 9.6 Function UðRÞ in rate of normal-yield ratio R which is attract to unity in plastic loading process
250
9
Unconventional Elastoplasticity Model: Subloading Surface Model
8 ! þ 1 for 0 R Re ðelastic stateÞ > > < [ 0 for Re \R\1 ðsubyield stateÞ UðRÞ ¼ 0 for R ¼ 1 ðnormal-yield stateÞ > > : \0 for R [ 1 ðover normal-yield stateÞ
ð9:13Þ
The material parameter Re is interpreted to be the ratio of the (half) stress amplitude rfl at the fatigue (or endurance) limit, i.e. the fatigue limit stress to the yield stress ry under the zero value of average stress r, i.e. Re ¼ rfl =ry jr¼0 . Fatigue limit is observed in steels, titanium, etc. but it is not observed in other materials involving non-ferrous metals. Note here that the incorporation of the material parameter Re does not mean the incorporation of the yield surface enclosing a purely-elastic domain as known from the fact: The plastic strain rate is predicted for the cyclic loading with a small stress amplitude under a high average stress by the subloading surface model with the incorporation of Re but it cannot be predicted if the yield surface enclosing a purely-elastic domain is incorporated as seen in the cyclic kinematic hardening model, i.e. the cylindrical yield surface, the multi-surface, the two-surface models which will be described in Sect. 10.3. Equation (9.12) is modified to satisfy Eq. (9.13) as follows: p hR Re i UðRÞ ¼ u cot 2 1 Re
ð9:14Þ
where h i is the Macaulay’s bracket defined by hsi ¼ ðs þ jsjÞ=2, i.e. s\0:hsi ¼ 0 and s 0: hsi ¼ s (s: arbitrary scalar variable). Equation (9.14) conforms to the fulfillment of the smoothness condition since the function UðRÞ decreases continuously from the infinite value UðRe Þ ð! 1Þ. If u is fixed to be constant, Eq. (9.9) with Eq. (9.14) can be integrated analytically as 9 2 p R0 Re p ep ep0 > 1 > exp u þ Re > R ¼ ð1 Re Þ cos cos > > p 2 1 Re > 2 1 Re > = p R0 Re cos for R0 Re 2 1 Re 2 1 Re > > > ln ep ep0 ¼ > > p R Re p u > > cos ; 2 1 Re ð9:15Þ R where ep jj ep jjdt and ep0 is the initial value of ep , whilst one must set R0 ¼ Re for R0 \Re . However, the judgment whether of R\Re or R Re is required in Eq. (9.14), although the yield judgment is not required. The time-differentiation of Eq. (9.6) of the subloading surface leads to
9.2 Subloading Surface (Hashiguchi) Model
251
@f ðrÞ : r R F R F ¼ 0 @r
ð9:16Þ
The following relation holds by virtue of the homogeneity of the function f ðrÞ in degree-one. @f ðrÞ : r ¼ f ðrÞ ¼ RF @r
ð9:17Þ
which yields @f ðrÞ :r @r ¼ RF ; n :r ¼ @f ðrÞ @f ðrÞ @r @r
1 ¼ n :r @f ðrÞ RF @r
ð9:18Þ
where @f ðrÞ @f ðrÞ = n @r @r
ð9:19Þ
Equation (9.16) with Eq. (9.18) results in " n:
! # F R r þ r ¼0 F R
ð9:20Þ
Now, adopt the associated flow rule
ep ¼ k n
ðk ¼ jj ep jj [ 0Þ
ð9:21Þ
The evolution rule of the normal-yield ratio is described as the equation R ¼ UðRÞ k instead of Eq. (9.9) for the expression of the flow rule in Eq. (9.21). Note, however,
that the equation R ¼ UðRÞ k does not hold for the expression of the flow rule
ep ¼ k @f ðrÞ=@r because of jj@f ðrÞ=@rjj 6¼ 1 in general. Substituting Eqs. (8.28) and (9.9) with Eq. (9.21) into Eq. (9.20), one has "
! # ' F UðRÞ fHn ðr; H; nÞ k þ k r ¼ 0 ð H ¼ fHn ðr; H; nÞ kÞ n : r F R
ð9:22Þ
252
9
Unconventional Elastoplasticity Model: Subloading Surface Model
It follows from Eqs. (9.21) and (9.22) that
k¼
n: r p ; M
p
e ¼
n: r p n M
ð9:23Þ
where F' UðRÞ fHn ðr; H; nÞ þ n :r M F R p
ð9:24Þ
which is reduced to the plastic modulus of the conventional elastoplasticity in p Eq. (8.27) in which note that M p is replaced to M in Eq. (9.24), i.e. p
M ¼
F' fHn ðr; H; nÞn : r ¼ M p F
ð9:25Þ
in the normal-yield state ðR ¼ 1 ! UðRÞ ¼ 0Þ. The strain rate is described by extending Eq. (8.30) to the subloading surface model as follows: • −1 n ⊗ n • = ( + ε• = −1 : σ• + n : σ p ): σ p n
M
M
ð9:26Þ
from which the magnitude of plastic strain k in terms of strain rate, denoted by the
symbol K, is derived as follows: •
Λ=
n : : ε• n : : ε• •p n , = ε p M p + n:: n M + n : : n
ð9:27Þ
The stress rate is given by extending Eq. (8.33) as follows: σ• = : ε• −
n : : ε• :n ⊗ n : ) : ε• : n = ( − p M + n:: n M + n : : n p
ð9:28Þ
The loading criterion is given by •
ε• p ≠ O : R > R e and Λ > 0 or n: : ε• > 0 ε• p = O : others
ð9:29Þ
where the judgment whether or not the stress reaches the yield surface is not required since the plastic strain rate is induced continuously as the stress approaches the normal-yield surface.
9.2 Subloading Surface (Hashiguchi) Model
253
1
1 Conventional plasticity model (u
)
Subloading surface model
σ R 1
0
1
0
2
Subloading surface f (σˆ ) = R F (H )
3
Normal-yield surface
ˆ ) = F (H ) f (σ
Fig. 9.7 Smooth stress–strain curve predicted by the subloading surface
The stress vs. strain curve by the subloading surface model is illustrated for the simplest case of the perfectly-plastic material in Fig. 9.7.
9.3
Distinguished Advantages of Subloading Surface Model
This model possesses the following distinguished abilities. (1) Smooth transition from elastic to plastic state is described, which is observed in real material behavior. Then, we don’t need suffer from the determination of an offset value (plastic strain value in yield point). In contrast, the determination is required in all of the other elastoplastic models since they assume a surface enclosing a purely-elastic domain leading to the abrupt elastic–plastic transition, while the determination is accompanied with an arbitrariness. The influences of the material parameter u on the function UðRÞ and the stress–strain curve are depicted in Figs. 9.8 and 9.9, respectively. The larger the material parameter u, the more rapidly the normal-yield ratio R increases causing the more rapid increase of stress, i.e. approaching the behavior of the conventional plasticity. (2) Plastic strain rate can be described even for low stress level and for cyclic loading process under small stress amplitudes since a purely-elastic domain is not assumed.
254
9
Unconventional Elastoplasticity Model: Subloading Surface Model
U ( R)
Smaller value of u
0
1
R
Fig. 9.8 Influence of u on function UðRÞ
(3) The yield-judgment whether or not the stress reaches the yield surface is unnecessary since the plastic strain rate develops continuously as the stress approaches the normal-yield surface. In contrast, the yield judgment is required in all of the other elastoplastic models since they assume a surface enclosing a purely-elastic domain. (4) The stress is automatically pulled back to the normal-yield surface when it goes
over the surface in numerical calculation because of R \0 for R [ 1 from Eq. (9.7) with Eq. (9.11)4 as seen in Fig. 9.10. In contrast, the particular operation to pull back the stress is required in all of the other models because they assume a surface enclosing a purely-elastic domain. The above-mentioned stress controlling function involved in the subloading surface model to pull-back the stress automatically to the yield surface in the plastic deformation process is quite beneficial for the explicit method in the numerical
0 Fig. 9.9 Influence of u on stress–strain curve
9.3 Distinguished Advantages of Subloading Surface Model
255
Fig. 9.10 Stress is automatically controlled to be attracted to the normal-yield surface in the subloading surface model
stress integration process. Besides, it is particularly important for the softening state in which the tangent stiffness modulus changes intensely deviating from the linear isotropic hardening curve. For the normal-yield state R ¼ 1 ðU ¼ 0Þ, the plastic strain rate in Eq. (9.23) with Eq. (9.24) is reduced to Eq. (8.29) with Eq. (8.27) for the conventional plasticity, i.e. p
e ¼
n: r
F' fHn ðr; H; nÞn : r F
n
For u ! 1 leading to the sudden decrease of the function U from U ! 1 p for R\1 to U ¼ 0 for R ¼ 1 in Eq. (9.11), the plastic modulus M in Eq. (9.24) drops suddenly from the infinite value to the value M p in Eq. (9.25) so that the present model behavior is reduced to the conventional elastoplasticity model behavior by choosing a large value of the material parameter u, thereby exhibiting an sudden transition from the elastic to plastic state. On the other hand, as u becomes smaller, a gentler transition from the elastic to plastic state is described. Therefore, u plays the role to alleviate the sudden transition from the elastic to plastic state. It follows from Eqs. (8.29) and (9.23) in the plastic loading process, fulfilling
k 0 or k 0, that
256
9
Unconventional Elastoplasticity Model: Subloading Surface Model
9 > M p [ 0 ! n : r [ 0; F [ 0 : normal hardening = p M ¼ 0 ! n : r ¼ 0; F ¼ 0 : normal nonhardening > ; M p \0 ! n : r \0; F \0 : normal softening
ð9:30Þ
for the conventional model and 9 p > M [ 0 ! n : r [ 0 : subloading hardening = p M ¼ 0 ! n : r ¼ 0 : subloading nonhardening > ; p M \0 ! n : r \0 : subloading softening
ð9:31Þ
for the subloading surface model. Here, it should be noted that the signs of M p or p
M and n : r coincide with each other in both models but they do not necessarily
coincide with the sign of F in the subloading surface model. The distinguished advantages of the subloading surface model in the descriptions of irreversible mechanical phenomena can be obtained by the simple modification of existing computer program for the conventional elastoplasticity model to add only one material parameter u for the evolution rule of the normal-yield ratio without any expense.
9.4
Numerical Performance of Subloading Surface Model
The stress controlling function of the subloading surface model is described in Sect. 7.3. This fact will be shown below by the numerical calculation for the response of the uniaxial loading behavior, adopting the simplest subloading surface model for the isotropic Mises material with the evolution rule of the normal-yield ratio in Eq. (9.9) with Eq. (9.12). The response of the conventional elastoplastic constitutive model is also shown for the comparison. The relations of the axial stress ra and the normal-yield ratio R versus the axial strain ea are depicted in Fig. 9.11. The responses adopting the linear isotropic hardening F ¼ F0 þ hc eep (hc : material constant) are depicted in Fig. 9.11a and those for the nonlinear isotropic hardening in Eq. (8.42) are shown in Fig. 9.11b. The two levels of axial strain increment dea ¼ 0:0006 and 0:0055 are input in the numarical calculations. Here, any special stress controlling algorithm to pull it back to the yield surface is not introduced. The material parameters are chosen as follows:
9.4 Numerical Performance of Subloading Surface Model
257
Material constants: Youg’s modulus : E ¼ 100000 MPa; Linear isotropic : hc ¼ 7000 MPa Hardening Nonlinar isotropic : sr ¼ 0:8; cH ¼ 50; Evolution of normal-yield ratio : u ¼ 200: Initial values: Hardening function: F0 ¼ 500 MPa; Stress : r0 ¼ O MPa The nonsmooth curves bent at the yield stress are expressed by the conventional model. Moreover, the stress deviates from the exact curve of concventional elastoplasticity. The deviation becomes large with the increases in the nonlinearity of hardening and in the increase of input strain increment. On the other hand, the stress is automatically attracted to the normal-yield surface in the subloading subloading surface model even for the quite large strain increment dea ¼ 0:0055 ð0:55%Þ. The zigzag lines tracing the exact curve are calculated such that the stress rises up when the it lies below the normal-yield surface but it drops 1000 0.0055
d a = 0.0006
800
600
a
0.0055
(MPa)
400
d a = 0.0006
200
0
2.0
1.50
d a = 0.0006 d a = 0.0055
1.0 d a = 0.0055 d a = 0.0006
R
0.50
0
0.02
0.04
a
0.06
0.08
(a) Linear isotropic hardening Exact curve of conventional elastoplasticity Calculated by the conventional elastoplastic model Calculated by the subloading surface model
Fig. 9.11 Numerical accuracies of the conventional elastoplastic model and the subloading surface model: Uniaxial loading behavior of Mises material with isotropic hardening
258
9
Unconventional Elastoplasticity Model: Subloading Surface Model 1000
0.0055 d a = 0.0006
800
a (MPa)
600 0.0055 400
d a = 0.0006
200
0
2.0
1.50
0.0055
0.0006
1.0 0.0055 d a = 0.0006
R
0.50
0
0.02
0.04
a
0.06
0.08
(b) Nonlinear isotropic hardening
Fig. 9.11 (continued)
down immediately if it goes over the normal-yield surface, obeying the evolution
rule of normal-yield ratio in Eq. (9.9) with Eq. (9.12), i.e. R [ 0 for R\1 and
p
R \0 for R [ 1. The plastic modulus M lowers than that in the conventional one and further it can be negative at the over normal-yield state R [ 1 leading to U\0,
p
while n : r \0 (subloading softening defined in Eq. (9.31)) holds for M \0
because of k [ 0 as known from Eqs. (9.23)–(9.25). The amplitude of zigzag decreases gradulally in the monotonic loading process, while, needless to say, the amplitude is smaller for a smaller input increment of strain. Eventually, the subloading surface model posseses the distinguished high ability for numerical calculation as verified also quantitatively in these concrete examples, which has not been attained in any other elastoplastic constitutive equations assumng a purely-elastic domain as will be described in Chap. 10. The drastic improvement of the conventional eladtoplasticity model by the subloading surface model is realized in Fig. 9.12. Neddless to say, this model is essntial for predicting the fatigue phenomenon in which a plastic deformation is induced during the cyclic loading behavior with a small stress amplitude. Besides, the two scale model (Lemaitre 1990) for the damage phenomenon in a high cycle fatigue, which requires the complex and irrigorous formulation, can be disused by adopting the subloading surface model as will be described in Sect. Sect. 15.9.
9.4 Numerical Performance of Subloading Surface Model
259
Conventional plasticity model
Excessively high peak stress
Pull-back of stress to yield surface is necessary
Yield judgement is necessary
Elastic deformation
0 Hardening state
Elastic deformation
0
Softening state
Yield judgement is unnecessary
0
Hardening state
Improvement
Subloading surface model
Offset value
0
Stress is pulled-back automatically to yield surface
Softening state
Improvement of monotonic loading behaviors
0
Improvement of accuracy and efficiency in numerical calculation
Fig. 9.12 Improvement of conventional elastoplasticity model by subloading surface model
These improvements can be attained merely by adding only one matrial constant u
in the evolution rule of the normal-yield ratio R. Then, the cost and the time required for the industrial design and development can be reduced tremendously even if we input irresponsible large value for the material constant u, noting that the deformation behavior based on the conventional elastoplastic constitutive equation is calculated for u ! 1.
9.5
On Bounding Surface Model with Radial-Mapping: Misuse of Subloading Surface Concept
The bounding surface model with a radial mapping by Dafalias and Herrmann (1980), Dafalias (1986), Dafalias and Talebat (2016), etc. possesses the primitive mathematical formulation which is basically different from the subloading surface model, although Dafalias stated: “It appears that the first time a radial mapping formulation was proposed, it was in reference to granular materials by Hashiguchi and Ueno (1977)” which is the original sentence in Dafalias (1986, p. 980). However, it has been adopted to the description of deformation behavior of soils as SANICLAY and SANISAND models (Dafalias et al. 2006; Dafalias and Talebat 2016, etc.). First note that the coined word “bounding surface” is nothing but the usual word “yield surface”. Therefore, this coined word must be disused in order to avoid the disturbance/confusion. Accordingly, let the “bounding surface model with a radial mapping” be called merely the “radial mapping model” hereinafter.
260
9
Unconventional Elastoplasticity Model: Subloading Surface Model
Note the following fundamental difference of the radial mapping model from the subloading surface model. (1) Any loading surface passing through the current stress is not incorporated, (2) Then, the strain rare cannot be formulated by rigorously based on the consistency condition and then it is formulated irrationally by the interpolation method. Therefore, the radial mapping model possesses the following crucial drawbacks and limitations which are not involved in the subloading surface model. (a) It is not guaranteed that the stress approaches the yield surface in the plastic loading process in the general loading state. (b) The cyclic loading behavior cannot be described leading to the open hysteresis loop in the unloading–reloading process and thus resulting in the excessively large plastic strain accumulation, while the subloading surface model has been extended to describe the cyclic loading behavior accurately as will be described in Chap. 11. Nevertheless, the bounding surface model with radial-mapping is used for the description of soil deformation, although it is never used for the description of metal deformation. The similar situation is seen in the hypoplasticity (Kolymbas and Wu 1993) in Eq. (7.113). The constitutive modelling in terms of the rate-nonlinear equation of the strain rate is performed in the hypoplasticity by extending the hypoelasticity in terms of the rate-linear equation of strain rate (Treuesdell 1955), while the fundamental requirements, i.e. the decomposition of the strain rate into the elastic and the plastic parts and the yield surface are ignored in the hypoplasticity. Unfortunately, the hypoplasticity is still now studied for soils, although it was studied once but faded out soon for metals ten years after the proposition of the hypoelasticity (Treuesdell 1956). This difference between metals and soils would be caused by the facts: The prediction of deformation behavior is required a high accuracy for metals but a quite low accuracy for soils. The radial mapping model and the hypopolasticity will disappear in the near future by the sound development of soil mechanics. Eventually, the ones using the bounding surface model with radial mapping for soils, i.e. the SANICLAY and the SANISAND models should dispose them and instead they should notice the subloading surface model for the rigorous deformation analyses and the sound development of soil plasticity.
9.6
Incorporation of Kinematic Hardening
The subloading surface based on the yield surface in Eq. (8.71) with the kinematic hardening is described as f ðb r Þ ¼ RFðHÞ The material-time derivative of Eq. (9.32) leads to
ð9:32Þ
9.6 Incorporation of Kinematic Hardening
261
@f ðb r Þ @f ðb rÞ :r : a RF H R F ¼ 0 @b r @b r
ð9:33Þ
Substituting the associated flow rule
^ ep ¼ k n
ðk ¼ jj ep jj 0Þ
ð9:34Þ
Equation (9.33) is rewritten substituting Eq. (8.74) as @f ðb r Þ @f ðb rÞ ^Þ RF k fH^n ðr; H; n ^Þ U k F ¼ 0 : r : k f k^n ðr; a; F; n @b r @b r
ð9:35Þ
Noting 9 > > > > > > > > > > =
@f ð^ rÞ :^ r ¼ f ð^ rÞ ¼ RF @^ r @f ð^ rÞ @f ð^ rÞ n ^ ¼ @^ @^ r r
ð9:36Þ
> > > > rÞ > @f ð^ > > :^ r @f ð^ > ^ r Þ @f ð^ r Þ n :^ r > @^ r > = ¼ ¼ 1= ; f ð^ rÞ @^ @^ r r RF Equation (9.35) is rewritten as
^: r ^ ^' Þ n ^: r ^ n n : k f k^n ðr; a; F; n
F' U ^Þ þ k ¼ 0 k fH^n ðr; H; n F R
ð9:37Þ
from which one has
^ : r p n ^: r n ^ k¼ p ; e ¼ p n M M
ð9:38Þ
where p
^: M n
h
F'
U ^ F fH^n ðr; H; nÞ þ R
'
^Þ b r þ f k^n ðr; a; F; n
i
ð9:39Þ
The loading criterion is given by ⎧• p • • ⎪ε ≠ O for R > R e and Λ > 0 or nˆ : : ε > 0 ⎨ ⎪ε• p = O for others ⎩
ð9:40Þ
262
9.7
9
Unconventional Elastoplasticity Model: Subloading Surface Model
Incorporation of Tangential-Inelastic Strain Rate
As presented in Eqs. (8.29), (8.81) and (9.38), the inelastic strain rate in the traditional constitutive equation has the following limitations. (i) The inelastic strain rate depends solely on the stress rate component normal to the yield surface, called the normal stress rate, but is independent of the component tangential to the yield surface, called the tangential stress rate, since it is derived merely based on the consistency condition. (ii) The direction of inelastic strain rate is determined solely by the current state of stress and internal variables but it is independent of the stress rate. On the other hand, it has been verified by experiments that an inelastic strain rate induced by the deviatoric part of the tangential stress rate, called the deviatoric tangential stress rate, influences considerably on a deformation in the non-proportional loading process deviating from the proportional loading path normal to the yield surface, which is called the tangential inelastic strain rate. Here, the spherical part of the tangential stress rate does not induce an inelastic strain rate, as Rudnicki and Rice (1975) verified based on the fissure model. In addition, the tangential inelastic strain rate is induced considerably in the plastic instability phenomena with the strain localization induced by the generation of the shear band and it influences on the macroscopic deformation and strength characteristics. To remedy these insufficiencies of the traditional plastic constitutive equation, various models have been proposed to date as follows: (1) Intersection of plural yield surfaces: Various models assuming the intersection of plural yield surfaces have been proposed (Batdorf and Budiansky 1949; Koiter 1953; Bland 1957; Mandel 1965; Hill 1966; Sewell 1973, 1974). The Koiter’s (1953) model was adopted by Sewell (1973, 1974), but it is indicated that the applicability of the model is limited to the inception of uniaxial loading. Models in this category cannot describe the latent hardening pertinently and are not readily applicable to general loading processes (cf. Christoffersen and Hutchinson (1979)). (2) Corner theory: The singularity of outward-normal of the yield surface is introduced by assuming the conical corner or vertex at the stress point on the yield surface. Therefore, the direction of plastic strain rate can take a wide range surrounded by the outward-normal of the yield surface (Christoffersen and Hutchinson 1979; Ito 1979; Gotoh 1985; Goya and Ito, 1991; Petryk and Thermann, 1977). There exist the two kinds of models: One kind is based on the assumption of an imaginary infinitesimal vertex and the other subsumes a finite projecting cone. The evolution rule of the cone cannot be formulated and the reloading from the cone surface after partial unloading cannot be described pertinently in the latter models. It was described by Hecker (1976) and Ikegami (1979) that the yield surface projects towards the loading direction generally but the formation of the so-called vertex is doubtful.
9.7 Incorporation of Tangential-Inelastic Strain Rate
263
(3) Hypoplasticity: This term was first used by Dafalias (1986) in the analogy to the term hypoelasticity introduced by Truesdell (1955) described in Sect. 5.3. Models in this category are classified into the two kind of models in which the direction of plastic
strain rate depends on the direction of the stress rate r =jj r jj (Mroz, 1966; Wang et al. 1970; Dafalias and Popov 1977; Hughes and Shakib 1986; Hashiguchi 1993a) and the models in which the direction of the plastic strain rate depends on
the direction of strain rate e =jj e jj (Hill 1959; Simo 1987b; Hashiguchi 1997; Kuroda and Tvergaard 2001). The singularity in the field of direction of plastic strain rate is introduced in the algebraic ways into these models, although it is done geometrically in the models described in (1) and (2). However, the magnitude of the plastic strain rate is derived from the consistency condition. Therefore, the plastic strain rate diminishes when the stress rate is directed tangentially to the yield surface, as in the traditional constitutive equations without the vertex. The constitutive equations described in (1)–(3) possess the following problems. (i) A formulation of pertinent model which fulfills the consistency condition and is applicable to the general loading process is difficult. (ii) The stress rate vs. strain rate relation becomes nonlinear. Therefore, the inverse expression cannot be derived, which renders deformation analysis to be difficult. Differently from the above-mentioned models, the following linear relation between the stress rate vs. strain rate with the tangential-inelastic strain rate, called the J2 -deformation theory, has been formulated by Budiansky (1959) and later Rudnicki and Rice (1975) by extending Eq. (6.73) with the isotropic Mises yield condition as follows: ð9:41Þ which can be rewritten as
ð9:42Þ
where the rate-linearity is retained. On the other hand, Hencky’s deformation theory (Hencky 1924) is described as ð9:43Þ
264
9
Unconventional Elastoplasticity Model: Subloading Surface Model
The time-derivative of Eq. (9.43) leads to ð9:44Þ Comparing Eq. (9.42) with Eq. (9.44), choosing Fðreq Þ so as to fulfill F ' ðeeqp ðreq ÞÞ ¼
3 1 2 /ðreq Þ þ /'ðreq Þreq
ð9:45Þ
and it is known that the J2 deformation theory coincides with Hencky’s deformation theory (9.43). However, it possesses crucial limitations as described in below. In what follows, let the tangential inelastic strain rate induced by the stress rate tangential to the loading surface be incorporated into the subloading surface model in the following (Hashiguchi 1998, 2005; Hashiguchi and Tsutsumi 2003; Hashiguchi and Protasov 2004; Khojastepour and Hashiguchi 2004a, b; Khojastepour et al. 2006). Here, note that the tangential inelastic strain rate would not be induced by the mean stress rate as would be inferred from the example for the fact that inelastic volumetric change would not be induced in metals (see Fig. 9.13). In what follows, let the inelastic strain rate induced by the deviatoric stress rate component tangential to the subloading surface be formulated for the general material unlimited to the Mises metal. Firstly, assume that the strain rate is com posed of the tangential-inelastic strain rate et in addition to the elastic strain rate and the plastic strain rate as follows (Hashiguchi 1998, 2013):
e ¼ ee þ e p þ et
ð9:46Þ
Further, assume that the tangential-inelastic strain rate et is induced by the tangential component of the stress rate r to the subloading surface in the deviatoric stress space, which is denoted by rt (Fig. 9.14) defined by ð9:47Þ
ð9:48Þ T't ≡ ' − n' ⊗ n' , T t'ijkl ≡ 'ijkl − n 'ijn 'kl
ð9:49Þ
9.7 Incorporation of Tangential-Inelastic Strain Rate
265
Fig. 9.13 Example showing the fact that tangential-inelastic strain rate is not induced by mean stress rate since inelastic volumetric change is not induced in metals
σ 'n
σ' 0 Subloading surface
n' σ ' σ 't ij'
f (σ ) = R F ( H )
Normal-yield surface f (σ) = F ( H )
Fig. 9.14 Normal and tangential stress rates for subloading surface model in deviatoric stress plane
fulfilling
ð9:50Þ
by virtue of ð9:51Þ
266
9
Unconventional Elastoplasticity Model: Subloading Surface Model
The fourth-order tensor is the deviatoric projection tensor in Eq. (1.194). The fourth-order tensor T't is the deviatoric-tangential projection tensor which transforms an arbitrary second-order tensor to the tangential part to the subloading surface in the deviatoric stress space and the second-order tensor subjected to this projection is designated by ðÞ't, i.e. t't T't: t leading further to T't: t't ¼ t't. Then, r't is
the deviatoric-tangential projection tensor of the stress rate r, which is called the deviatoric-tangential stress rate. Note that the tangential-inelastic strain rate does not affect the yield surface because it is relevant to the tangential component but irrelevant to the normal component to the yield surface.
Now, assume that the tangential-inelastic strain rate et is related linearly to the tangential-deviatoric stress rate r't in the normal-yield state (R ¼ 1) as follows: • −1 ε t = : σ• 't
ð9:52Þ
Let Eq. (9.52) be extended for the sub-yield state as follows: • −1 ε t = T ( R) : σ• 't
ð9:53Þ
In this equation TðRÞ was given by (Hashiguchi, 1998, 2013, 2017, 2020) as follows: TðRÞ ¼ ~cR~n
ð9:54Þ
which is the monotonically-increasing function of the normal-yield ratio R, where ~c and ~ n are the material constants, always fulfilling the continuity and the smoothness conditions in Eqs. (9.1) and (9.2). However, it increases unlimitedly with R and thus it would be physically unacceptable to the high value of the normal-yield ratio, i.e. high over-yield state R [[ 1 observed in the viscoplastic deformation. The function TðRÞ without this effect would be given by the following equation.
ð9:55Þ
9.7 Incorporation of Tangential-Inelastic Strain Rate
267
where ~c and ~ n are the material constants. The following properties are furnished in Eq. (9.55) 1) The material parameter ~c exhibits the maximum value of the function TðRÞ. 2) The function TðRÞ approaches the maximum ~c earlier for a larger value of the material constant ~ n. 3) The relation of the function TðRÞ vs. the normal-yield ration R exhibits the inflection point in a smaller value of R for a larger value of the material constant ~n. In particulier, the inflection point appears at R = 1 for ~n ¼ 1 as shown in Fig. 9.15. Adding the tangential-inelastic strain rate in Eq. (9.53) to Eq. (9.26), the strain rate is given by n: σ• • ε = −1 : σ• + p n + T ( R) −1 : σ• 't M ð9:56Þ n n −1 ⊗ −1 • ( R ) = + : T't : σ p +T
(
)
M
In what follows, we assume the elastic modulus tensor fulfilling
T't : : n = O
ð9:57Þ
R T ( R)
c
1
d 2T ( R ) 0 dR 2
0 c (1 e)e
d 2T ( R) 0 dR 2
c (1 e)e
c
T ( R)
Fig. 9.15 Function TðRÞ for ~ n¼1
1
R
268
9
Unconventional Elastoplasticity Model: Subloading Surface Model
which holds in the Hooke’s law in Eq. (5.35) for example. Taking account of Eq. (9.57), it follows from Eq. (9.56) that • • T't : : ε = [1 + T ( R)] σ't
ð9:58Þ
1 • T't : : ε 1 + T ( R)
ð9:59Þ
leading to σ• 't =
Substituting Eq. (9.59) into Eq. (9.53), one obtains •
εt =
T ( R) −1 • : T't : : ε 1 + T ( R)
ð9:60Þ
The expression Eq. (9.27) for the plastic multiplier in terms of strain rate is obtained from Eq. (9.56), noting Eq. (9.47)1. Then, the stress rate is given by substituting Eq. (9.46) with Eq. (9.27) and Eq. (9.60) into the relation based on the infinitesimal linear elasticity in Eq. (7.83) as follows:
σ• = : (ε• − ε• p− ε• t ) = : ε• −
n: : ε• :n − T ( R) T' : : ε• t 1 + T ( R) M + n: : n p
: n ⊗ n: − T ( R) T' : = − p t M + n: : n 1 + T ( R)
(
):ε
ð9:61Þ
•
Equations (9.56) and (9.61) are given for the Hooke’s law in Eq. (7.84) as follows: n n T ( R) • −1 ð9:62Þ T't : σ ε• = + ⊗p +
(
M
2G
)
T ( R) n n: T't : ε• σ• = − p: ⊗ − 2G + n: :n 1 + T ( R) M
(
)
ð9:63Þ
Here, it is known that the bulk modulus K is irrelevant to the tangential-inelastic strain rate which is induced only by the deviatoric part of stress rate. The tangential-inelastic strain rate is essentially tangential to the loading surface (yield and subloading surface) and thus it is always induced as long as the stress rate tangential to the loading surface is imposed. Therefore, the loading criterion for the tangential inelastic strain rate is not required and thus the loading criterion for the constitutive equation with the tangential-inelasticity is given by the equation identical to the loading criterion for the ordinary elastoplastic constitutive equation without the tangential-inelastic strain rate, i.e. Eq. (9.40).
9.7 Incorporation of Tangential-Inelastic Strain Rate
269
Equations (9.56), (9.60) and (9.63) were altered by Fincato and Tsutsumi (2019) and Tsutsumi et al. (2019) as
(
n n −1 ε• = + ⊗p + M
)
T ( R) T ( R) • • σ• 't T't : σ , ε t = 2G [1 − T ( R)] 2G [1 − T ( R)]
ð9:64Þ
n n: σ• = − p: ⊗ − 2GT ( R)T't : ε• M + n: : n
(
)
ð9:65Þ t
This alteration cannot be allowed because the tangential-inelastic strain rate e is induced infinitely for TðRÞ ! 1, i.e. R ! c1=n for Eq. (9.54) in which the sinn gularity of the tangential inelastic strain rate field is induced. It contradicts to the original physical meaning of the tangential-inelastic strain rate induced by the stress rate tangential to the subloading surface, which is ignored in the associated flow rule. It cannot be applied to the over normal-yield state ðR [ 1Þ occurring in the overstress state which will be described in Chap. 14. 1
σ 0 2
Yield surface
3
Input : σ
1
1
ε t : suddenly induced.
ε t : gradually develops as the subloading surface expands.
0 2
Subloading surface
0 3
2
Yield surface
3
Normal-yield surface Subloading surface model: Hashiguchi (2005), fulfilling continuity and smoothness conditions.
Conventional plasticity model: Rudnicki and Rice model (1975), violating continuity and smoothness conditions.
Fig. 9.16 Incorporation of tangential inelastic strain rate illustrated for von Mises yield surface
270
9
Unconventional Elastoplasticity Model: Subloading Surface Model t
The tangential-inelastic strain rate e develops gradually as the current stress approaches the normal-yield surface, i.e. the subloading surface expands fulfilling the continuity and the smoothness condition in the subloading surface model as shown in Fig. 9.16 for the isotropic Mises material. The validity of Eq. (9.62) or (9.63) has been verified by Hashiguchi and Protasov (2004) for metals and Hashiguchi and Tsutsumi (2001, 2003, 2007) and Tsutsumi and Hashiguchi (2005) and for geomaterials. On the other hand, if the tangential-inelastic strain rate is incorporated into the plasticity model assuming the yield surface enclosing a purely-elastic domain, both of the continuity and the smoothness conditions in Eqs. (9.1) and (9.2) are violated, since the tangential-inelastic strain rate is induced suddenly at the moment when the stress reaches the yield surface as illustrated for the J2 -deformation model of Rudnicki and Rice (1975) in Fig. 9.16.
9.8
Limitation of Initial Subloading Surface Model
The conventional elastoplastic constitutive equation is incapable of predicting the plastic strain accumulation because the conventional yield surface is assumed to enclose the purely-elastic domain as shown in Fig. 9.17a. On the other hand, an excessively large strain accumulation is described during the cyclic loading process because only elastic deformation is induced in the unloading process by the initial subloading surface model described in this chapter as shown in Fig. 9.17b. The cyclic plasticity models proposed to describe the cyclic loading behavior will be described in the subsequent chapters.
Elastoplastic state Elastic state
0
(a) Conventional elastoplastic model
0
(b) Initial subloading surface model
Fig. 9.17 Stress versus strain curve predicted by conventional elastoplastic model and initial subloading surface model in cyclic uniaxial loading
Chapter 10
Classification of Plasticity Models: Critical Reviews and Assessments
Accurate description of plastic deformation induced during a cyclic loading process is required for the mechanical design of machinery subjected to vibration and buildings and soil structures subjected to earthquakes since the middle of the last century. Elastoplastic constitutive model formulated for this aim is called the cyclic plasticity model. Substantially, the key of the pertinence in cyclic plasticity model is how to describe appropriately a small plastic strain rate induced by the rate of stress inside the yield surface. Therefore, a quite delicate formulation of plastic strain rate developing gradually as the stress approaches the yield surface is required to this end. Here, needless to say, the continuity and the smoothness conditions described in Sect. 8.1 would have to be fulfilled in a cyclic plasticity model. Various cyclic plasticity models have been proposed to date, while most of them violate the continuity and/or the smoothness conditions unfortunately. Then, the beginners for the cyclic plasticity model would be perplexed as to which model is most pertinent and should be chosen for their study and analyses. In this chapter, the basic mathematical structures of plasticity models proposed hitherto will be discussed and assessed comprehensively. In this context, explicit evolution rules of particular internal variables such as the isotropic hardening, the kinematic hardening rule for metals and the rotational hardening for soils, etc. which differ depending on the kinds of materials will not be discussed in detail. On the other hand, the kinematic hardening rules which are concerned only to the metals are discussed still now extensively by J.-L. Chaboche, S. Hassan, S. Kyrakides, N. Ohno, Y. F. Dafalias, etc., where unfortunately the basic mathematical structures of plasticity models themselves influencing basically the cyclic loading behavior have not been discussed enough.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_10
271
272
10.1
10
Classification of Plasticity Models: Critical Reviews …
Cyclic Loading Behavior
The actual stress versus strain curve during the cyclic loading under a constant stress amplitude is shown by the blue colored line in Fig. 10.1, showing the accumulation of strain, i.e. the mechanical ratchetting. Here, note that the slight compressive plastic strain is induced in the unloading process to the stress-free state by the following mechanical reason: The solids are composed of material particles, e.g. the crystal particles in metals and soil particles in soils. Now, suppose the uniaxial loading process of the solid. The material particles are elongated to the tensional direction and the mutual slips are induced so that the tensional plastic strain is induced in the loading process. On the other hand, the elongated material particles return to the initial shapes which cause the slight mutual slips to the compressive direction so that the slight compressive plastic strain is induced in the unloading process to the stress free state. Further, the tensional elastic deformation and then the tensional plastic strain rate are induced in the reloading process. Consequently, the closed hysteresis loop is induced as shown in Fig. 10.1. However, the stress versus strain curve in which an merely an elastic deformation is repeated during the cyclic loading is predicted by the conventional elastoplastic model as shown by the red-colored line in Fig. 10.1. On the other hand, the excessively large strain accumulation is predicted by the initial subloading surface model described in Chap. 9 as was shown in Fig. 9.17, because only the elastic deformation is induced in the unloading process. Consequently, the cyclic loading behavior cannot be predicted realistically by the conventional elastoplastic model and the initial subloading surface model. Then, various plasticity models described in the next section have been proposed hitherto.
Fig. 10.1 Cyclic loading (mechanical ratchetting) behavior
10.2
Classification and Assessment of Plasticity Models
10.2
273
Classification and Assessment of Plasticity Models
The plasticity models, i.e. the cylindrical yield surface (Chaboche) model (Chaboche et al. 1979), the multi-surface (Mroz) model (Mroz 1967; Iwan 1967), the two-surface (Dafalias) model (Dafalias and Popov 1975; Krieg 1975) and the subloading surface (Hashiguchi) model (Hashiguchi 1978, 1980, 1989; Hashiguchi and Ueno 1977) proposed to date are classified from some aspects (Fig. 10.2). (1) Classification from conventional and unconventional models The plasticity models were classified into the conventional plasticity models and the unconventional plasticity models by Drucker (1988). The inside of the yield surface is assumed to be the purely-elastic domain inside which the stress lies always in the former. On the other hand, the latter aim at the description of the plastic deformation induced by the rate of stress inside the yield surface. The plasticity models proposed so far are classified as follows (Figure 10.3): Conventional plasticity model: Cylindrical yield surface (Chaboche) model, The Chaboche model is the conventional elastoplasticity model possessing the cylindrical yield surface which obeys the isotropic and the kinematic hardenings in the principal stress space. Therefore, needless to say, it is incapable of describing the cyclic loading behavior for the stress amplitude smaller than the size of the cylindrical surface. Moreover, it is incapable of describing the cyclic loading behavior appropriately even for the stress amplitude larger than the size of yield surface as will be described in detail in Sect. 10.3.2. Therefore, the Chaboche model is the primitive model incapable of the cyclic loading behavior, although it has been used and studied widely against the progress of time. First of all, it would be quite reckless to describe the cyclic loading behavior only by making the conventional yield surface expand (isotropic hardening) and translate (kinematic hardening). Unconventional plasticity models: Multi-surface (Mroz) model, Two-surface (Dafalias) model and the subloading surface (Hashiguchi) model. The cyclic plasticity model would have to belong to the unconventional model. In this aspect, the Chaboche model cannot be classified into the cyclic plasticity model. (2) Classification from existence or removal of elastic-domain Existence of elastic domain: Cylindrical yield surface (Chaboche) model, Multi-surface (Mroz) model and Two-surface (Dafalias) model, Removal of elastic-domain: Subloading surface (Hashiguchi) model (3) Classification from geometrical variation of elastic domain or loading surface The cyclic plasticity models are classified from the aspect of the geometrical variation of elastic domain or loading surface as follows:
Fig. 10.2 Classification of plasticity models
Chaboche et al. (1979) Ohno-Wang (1993)
Two-surface model Dafalias-Popov (1975) Yoshida-Uemori (2002)
Multi-surface model Mroz (1966, 1967)
Plural surface models with small elastic-domain
Hashiguchi-Ueno (1977) Hashiguchi (1980, 1989)
Subloading surface model
Unconventional plasticity models
10
Cylindrical yield surface model
Conventional plasticity model
Plasticity models
274 Classification of Plasticity Models: Critical Reviews …
Classification and Assessment of Plasticity Models
Fig. 10.3 Conventional and unconventional plasticity models
10.2 275
276
10
Classification of Plasticity Models: Critical Reviews …
Translation of elastic domain: Cylindrical yield surface (Chaboche) model, Multi-surface (Mroz) model and Two-surface (Dafalias) model Expansion of loading surface: Subloading surface (Hashiguchi) model The critical reviews for each of the above-mentioned plasticity models will be given in the subsequent sections.
10.3
Plasticity Models with Elastic Domain
The models other than the subloading surface model, i.e. the cylindrical yield surface model (Chaboche et al. 1979; Ohno and Wang 1993), the multi-surface model (Mroz 1967) and the two-surface model (Dafalias and Popov 1975; Yoshida and Uemori 2002a) assume the existence of the purely elastic domain inside which the stress lies always. The basic mathematical structures and the mechanical properties of these models will be examined in this section.
10.3.1
Common Drawbacks in Models with Elastic-Domain
The plasticity models assuming the elastic-domain inside which the stress lies always possess the following common crucial drawbacks. (1) The plastic strain rate cannot be described for the rate of stress inside the elastic-domain surface, so that the plastic strain accumulation, i.e. mechanical ratchetting phenomenon cannot be predicted at all. In facts, the original proposers (Chaboche, Morz, Dafalias) and their followers (Ohno and Yoshida in Japan unfortunately) have never shown a simulation of the mechanical ratchetting behavior in the tension (positive) or compression (negative) one side, i.e. the pulsating loading except for the physically unacceptable method by Yoshida’s group (e.g. Yoshida and Uemori 2002b, 2003; Yoshida and Amaishi 2020) as will be explained in Sect. 10.3.3. (2) The tangent stiffness modulus changes discontinuously resulting in the violation of the smoothness condition in Eq. (9.2) at the moment when the stress reaches the elastic-domain surface, (3) The judgement whether the stress reaches the elastic-domain surface is necessary, (4) Then, the numerical calculation of cyclic loading behavior is the time-consuming. Here, note that an excessive downsizing of the elastic-domain would results in the instability of its outward-normal, so that the direction of the plastic strain rate based on the normality-rule (associated flow rule) changes unstably as shown in Fig. 10.4. Therefore, the cyclic plasticity models other than the subloading surface model are inapplicable to the description of the cyclic loading behavior under the small stress amplitude.
10.3
Plasticity Models with Elastic Domain
277 ε p : stable
Yield surface
σ ε p : unstable σ Downsized yield surface or elastic domain
0
ij
Fig. 10.4 Downsizing of elastic-domain causes the instability of direction of plastic strain rate even for slight change of stress.
Needless to say, the cylindrical yield surface model is incapable of describing the cyclic loading (mechanical ratchetting) behavior at all, since it is the typical conventional elastoplasticity model as will be described in detail in Sect. 10.3.2. Further, the multi-surface (Mroz) model and the two-surface (Dafalias) model are also incapable of describing that behavior as will be described in Sects. 10.3.3 and 10.3.4. The plasticity models with an elastic domain, i.e. the cylindrical elastic-domain model, the multi-surface model and the two surface model described in this section have not been propagated by the original proposers themselves after the recent ten years. The proposers of these models themselves (Chaboche, Mroz, Dafalias) would have recognized the crucial defects of their own-models so that they have not commented on their models in the recent ten years, although these models are still studied by their followers (Ohno, Yoshida, etc.) unfortunately yet at present.
10.3.2
Cylindrical Yield Surface (Chaboche) Model: Ad Hoc. Primitive Conventional Model Limited to Simple Metal Behavior
The conventional plasticity model limited to the cylindrical yield surface with the plural kinematic hardening rules was proposed by Chaboche et al. (1979) and improved by Chaboche and Rousselier (1983), Chaboche (1989), Ohno and Wang (1993), etc. It may be called the cylindrical yield surface model which is called the
278
Classification of Plasticity Models: Critical Reviews …
10
multikinematic model by Chaboche (2008) himself. Strictly speaking, however, it may not be called a cyclic plasticity model, since it is incapable of describing the plastic strain rate by the rate of stress inside the yield surface. (a) Chaboche model The following yield surface with the isotropic hardening and the Armstrong-Frederick (1966) kinematic hardenings was introduced by Chaboche et al. (1979). rffiffiffi 3 0 jj^ r jj ¼ F0 þ F 2
ð10:1Þ
where the isotropic hardening function F evolves by the following equation.
F ¼ cðF s FÞe epq
ð10:2Þ
c is the material constant and F s is the material constant describing the saturation value of F. Equation (10.2) leads to F ¼ F s þ ðF 0 F s Þ exp½cðeeqp eeqp 0 Þ which coincides with Eq. (8.42) by setting Fs ¼ ð1 þ sr ÞF0 and c ¼ cH . The kinematic hardening rule was given by the superposition of the several non-linear kinematic hardening rules of Armstrong and Frederick (1966) in Eq. (8.89) as follows:
a¼
n X
ai
ð10:3Þ
i¼1
where ai
rffiffiffi 2 bi ai jjep k ðno sumÞ ¼ Ai e 3 p
ð10:4Þ
Ai and bi ði ¼ 1; 2; ; nÞ are the material constants, while n is chosen usually 4 8. Equation (10.3) is integrated for the uniaxial loading process as follows: aa ¼
n X Ai i¼1
bi
½1 expðbi epa Þ
ð10:5Þ
for epa [ 0 under the initial condition aa ¼ 0 and epa ¼ 0. The uniaxial loading behavior of Chaboche model is illustrated in Fig. 10.5, where the isotropic hardening is not incorporated by setting c ¼ 0 for simplicity.
10.3
Plasticity Models with Elastic Domain
279
Fig. 10.5 Chaboche model (after Lamaitre and Chaboche 1990)
The mechanical ratchetting is overestimated by use of the Armstrong-Frederic nonlinear kinematic hardening equation in its original form, because the kinematic hardening is suppressed but it induced excessively in the opposite direction in the reverse loading by the dynamic recovery part in the Armstrong-Frederic nonlinear kinematic hardening rule. Then, Chaboche (1991) modified Eq. (10.4) by incorporating the threshold for activation of the dynamic recovery part into the fourth kinematic equation, while Eq. (10.4) is used for the first to the third kinematic equations, as follows: a4
rffiffiffi 2 a a4 jjep jj b4 1 ¼ A4 e 3 jja4 jj p
ð10:6Þ
where a is the material constant designating the threshold value of the activation of the dynamic recovery which is zero for jja4 jj\a. (b) Ohno-Wang model Based on the opinions: (1) By introducing the threshold for the activation of the dynamic recovery the excessive mechanical ratchetting is suppressed as was proposed by Chaboche (1991), (2) The rate of the dynamic recovery depends on direction of the backstress rate in the multi-dimensional deformation. Ohno and Wang (1993) have modified the Armstrong-Frederic kinematic hardening rule by incorporating the idea of the threshold value by Chaboche (1991) as follows:
280
10
Classification of Plasticity Models: Critical Reviews …
a¼
n X
ai
ð10:7Þ
i¼1 ai
2 p p ai ai ¼ hi e H½fi e : ai r i 3
ðno sumÞ
ð10:8Þ
where hi and ri ði ¼ 1; 2; ; nÞ are the material constants. The dynamic recovery activates when the following condition is satisfied in each evolution rule. fi a2i ri2 where
ð10:9Þ
rffiffiffi 3 ai jjai jj 2
ð10:10Þ
H½ is the Heaviside step function, i.e., H½s ¼ 1 for s 0, H½s ¼ 0 for s\0. Then, the kinematic hardening proceeds when fi a2i ri2 ¼ 0 is satisfied and the plastic strain rate is induced directing outward of the surface described by Eq. (10.9) but it ceases until fi a2i ri2 ¼ 0 ðH½fi ¼ 1Þ is satisfied. The fact that Eq. (10.8) is the possible candidate for the rate of the kinematic hardening fulfilling the time-derivative of Eq. (10.9) is as ascertained by
fi ¼ 2ai ai ¼ 2ai
rffiffiffi rffiffiffi 3 3 ai : ai ðjjai jjÞ ¼ 2ai 2 2 jjai jj
rffiffiffi 3 ai 2 ai ai ¼ 2ai :hi ep H½fi ep : ai ri 2 jjai jj 3 rffiffiffi a 3 ai 2 ai ai ai i :hi k k : ¼ 2ai jjai jj ai ai 2 jjai jj 3 jjai jj 0 1 * + rffiffiffi a a 3 ai 2 a a i i i i B C qffiffi ¼ 2ai :hi @ k k : qffiffi A¼0 jjai jj 2 jjai jj 3 jjai jj 3 3 jja jj jja jj 2
i
2
i
The piecewise-linear relation of the backstress versus plastic strain is depicted for the proportional loading process by Eq. (10.7) with Eq. (10.8). Further, Eq. (10.8) was modified by replacing the Heaviside step function to the power function proposed by Henshall et al. (1987) as follows:
ai ¼ hi
mi 2 p ai ai ai e ep : ri ai r i 3
ðno sumÞ
ð10:11Þ
10.3
Plasticity Models with Elastic Domain
281
where mi ði ¼ 1; 2; ; nÞ are the material constants. On account of this modification, the backstress versus plastic strain curve in the monotonic loading process becomes smooth. Here, note that Eq. (10.11) for mi ! 1 is reduced to Eq. (10.8). On the other hand, Eq. (10.11) for mi ¼ 0 is reduced to the Armstrong-Frederick nonlinear kinematic hardening rule. Further, the following model was proposed by Abdel-Karim and Ohno (2000).
a i ¼ hi
2 p ai ai e li ai jjep jj H½fi ep : ai r i 3
ðno sumÞ
ð10:12Þ
which comprises both of the Armstrong-Frederick’s (1966) and Ohno and Wang’s (1993) equations, li being the material constants. The defect of the original Armstrong-Frederick nonlinear kinematic hardening rule has been revealed such that the overshooting of the stress vs strain curve is predicted in the reloading process after the small reverse loading (cf. Dafalias, 2008), when this model is applied to the prediction of the cyclic loading behavior by the conventional plasticity model with the yield surface enclosing the purely-elastic domain obeying the isotropic and the kinematic hardenings. The overshooting of the stress-strain curve resulting in the overestimation of the mechanical ratchetting is the unavoidable nature of this model, because the reloading curve translates in parallel to the preceding monotonic curve in its underside keeping the almost constant interval which is caused by the dynamic recovery part in the kinematic hardening rule. Nevertheless, various modifications of the Armstrong-Frederick nonlinear kinematic hardening model have been proposed by Henshall et al. (1987), Hassan et al. (2008), Dafalias et al. (2008), Dafalias and Feigenbaum (2011), Okorokov et al. (2019), etc. in addition to Chaboche et al. (1979), Chaboche (1991), Ohno and Wang (1993) and Abdel-Karim and Ohno (2000) described above. Then, the formulation has become more and more complicated. Therein, the modifications of the Armstrong-Frederick kinematic hardening model have been repeated through its superposition by Chaboche et al. (1979) in the first stage, the incorporation of the threshold term (Chaboche, 1991) or the power function (Henshall et al., 1987), the multiplication (Dafalias, 2008), etc. Unfortunately, however, they adopt only the cylindrical yield surface (Chaboche-Ohno) without taking care about the afore-mentioned defect of the basic primitive structure of this model as a cyclic plasticity model, although they are concerned only with the large stress amplitude surrounding the yield surface. In facts, the Chaboche-Ohno model has been irrespective to the general cyclic loading behavior including the partial unloading-reloading, the pulsating loading behaviors, etc. because it falls within the framework of the conventional plasticity. Nevertheless, an appropriate modification of Chaboche-Ohno model could not be found and will not found forever even for the description of the cyclic loading behavior with large stress amplitude contrary to the expectation of these researchers represented by Chaboche. The reason for the impossibility is caused by the fact that they are concerned only with the unloading behavior through the dynamic recovery part. However, the rational formulation of the cyclic loading behavior cannot be
282
10
Classification of Plasticity Models: Critical Reviews …
attained by the modification of the unloading behavior in the kinematic hardening rule but it can be attained by the modification of the reloading behavior as will be shown actually in Sect. 11.6 for the extended subloading surface model. Eventually, it should be concluded that the conventional plasticity model including the Chaboche model is basically incapable of describing the cyclic loading behavior even for a large stress amplitude, although a half century has been wasted after Chaboche et al. (1979). First of all, the superposition of equations started by Chaboche et al. (1979) itself results in containing many material parameters whose physical meanings are unclear, exposing the physical and mathematical immaturity as a constitutive equation. The Armstrong-Frederick model itself is physically and mathematically reasonable for metals by the slight modification in Eq. (8.89). In fact, the general cyclic loading behavior of metals can be described accurately by adopting it in the extended subloading surface model as will be shown in Chaps. 11 and 12. Besides, the kinematic hardening is relevant to the plastically-incompressible metals but it is irrelevant to the plastically-compressible materials, e.g. soils, rocks, concrete, ceramics, glass, etc. for which the rotational hardening is dominant as will be described in Chap. 13 for soils. The cyclic plasticity model should be studied seriously by adopting the unconventional yield surface and formulating the plastic strain rate induced by the rate of stress inside the yield surface. Besides, the improvement for avoiding the undershooting curve cannot be attained by the modification of the unloading property as the dynamic recovery part but can be attained by the improvement of the reloading behavior in the unconventional plasticity model as will be described in Sect. 11.6. Consequently, it should be recognized that 1. The cyclic plasticity formulation can never be realized by adopting the conventional plasticity model enclosing the purely-elastic domain. 2. It is impossible to describe the cyclic loading behavior only by the modification of the kinematic hardening rule. 3. Only the cyclic loading behavior under the large stress amplitude in positive and negative stress sides covering the yield surface has been shown in the Chaboche model and its related models. It can be described realistically even by the initial subloading surface model in Sect. 9.2 as shown by Hashiguchi (1981). However, the description of the cyclic loading behavior under the half stress amplitude in positive or negative one side inducing the mechanical ratchetting is of crucial imporatance for the engineering practice. 4. The rational formulation for the cyclic plastic model can be realized only by the extended subloading surface model described in detail from the subsequent chapters in which the plastic strain rate is induced by the rate of stress inside the yield surface and reloading behavior is modified so as to return promptly to the preceding stress vs. strain curve, keeping the kinematic hardening rule in Eq. (8.89) as it is. On the other and, the various nonsense modifications of the unloading behavior have been repeated through the alteration of the dynamic recovery part in the Armstrong-Frederick kinematic hardening equation.
10.3
Plasticity Models with Elastic Domain
283
Unfortunately, however, the Chaboche model was officially implemented in the commercial software Marc and Abaqus, LS-DYNA etc. so that it is used widely by metal engineers because it can be understood even by the beginners of elastoplasticity theory possessing the elementary knowledge only of the Mises yield condition with the kinematic hardening. However, it should be recognized that the mechanical designs of solids and structures subjected to the general cyclic loading ehaviour by this model would result in dangerous accidents because this model is incapable of describing the cyclic loading behavior pertinently even for the stress amplitude larger than the yield surface, while, needless to say, it cannot describe the cyclic loading behavior for the stress amplitude smaller than the yield surface at all.
10.3.3
Multi-surface (Mroz) Model: Incapable of Describing Mechanical Ratchetting
Mroz (1966, 1967, 1976) and Iwan (1967) proposed the multi-surface (Mroz) model based on the following basic assumptions. (a) The elastic-domain surface and plural encircled subyield surfaces are incorporated inside the yield surface, while the ratios of the sizes of these surfaces to the yield surface are kept constant throughout a deformation. (b) The elastic-domain surface and the subyield surfaces are pushed out by the current stress point so that plural surfaces contact at a point. (c) Plastic modulus is prescribed by the size of the subyield surface on which the current stress lies, while it is smaller for a more outer subyield surface. The uniaxial deformation ehaviour of the multi-surface model is illustrated in Fig. 10.6 for the simple material without the isotropic and the kinematic hardenings. However, this model possesses the following defects in addition to the crucial defects caused by the incorporation of the elastic-domain described in Sect. 10.3.1. (1) Plural subyield surfaces contact at the current stress point and there exist plural plastic moduli at that contact point so that the singular point in the field of plastic modulus is induced therein. Numerical calculation becomes unstable for the cyclic loading behavior in the vicinity of contact point. (2) It is physically impertinent that the elastic-domain surface contacts with the conventional yield surface in the fully-plastic state. (3) Stress transfers to a larger subyield surface by moving the half of the difference of the radii of subyield surfaces in the initial loading process as shown in Fig. 10.6b. On the other hand, it transfers to a larger subyield surface by moving the difference of the diameters of subyield surfaces in the unloading-reverse loading process as shown in Fig. 10.6d. Therefore, the Masing rule (Masing 1926) meaning that the curvature of stress–strain curve in the unloading-reverse loading decreases to a half of the curvature of initial loading curve is described exactly and simply as shown in Fig. 10.6d. By virtue
284
10
Classification of Plasticity Models: Critical Reviews …
Fig. 10.6 One-dimensional loading ehaviour predicted by multi-surface model
of this mechanical feature, this model has been used widely. However, the variation of curvature observed in real material behavior is not so large as described by the Masing rule. (4) The stress–strain curve converges to the fixed loop for the nonhardening yield surface as shown in Fig. 10.7a. The hysteresis loop of the stress-strain curve becomes thinner with increase of the cycles for the hardening yield surface as shown in Fig. 10.7b. Therefore, the accumulation of plastic strain during a pulsating stress loading, called the mechanical ratcheting, cannot be described at all by this model. Therefore, the deformations of machinery and structures subjected to cyclic loading are predicted to be unrealistically small by the multi-surface model, resulting in a risky mechanical design of solids and structures.
10.3.4
Two Surface (Dafalias) Model: Incapable of Describing Plastic Strain Rate in Unloading Process
Dafalias and Popov (1975, 1976) and Krieg (1975) proposed the two surface model based on the following assumptions:
10.3
Plasticity Models with Elastic Domain
285
Fig. 10.7 Prediction of cyclic loading behavior by the multi surface model under a constant stress amplitude: Plastic strain accumulation cannot be described at all
(a) The elastic domain is incorporated inside the conventional yield surface which is renamed as the “bounding surface” by Dafalias and Popov (1975) and “limit surface” by Krieg (1975). The bounding surface is merely the usual yield surface. Therefore, this coined word must be disused in order to avoid the disturbance/confusion. (b) The ratio of the size of the elastic-domain surface to that of the yield surface is kept constant throughout the deformation. (c) The elastic-domain is pushed out by the current stress point and translates toward the conjugate point on the yield surface, while the outward-normal at conjugate point on the yield surface is identical to that at the current stress point on the elastic-domain surface. (d) The plastic modulus depends on the distance from the current stress on the elastic-domain surface to the conjugate stress on the yield surface. Here, it is required that the elastic-domain surface must translate so as not to intersect with the yield surface because the direction of plastic strain rate becomes indeterminate if they intersects. The rigorous translation rule was derived by Hashiguchi (1981, 1988). This model has been adopted widely for the prediction of deformation behavior of metals (cf. e.g. Dafalias and Popov 1976; McDowell 1985, 1989; Ohno and Kachi 1986; Ellyin 1989; Hassan and Kyriakides 1992; Yoshida and Uemori 2002a, 2002b, 2003; Yoshida and Amaishi 2003).
286
10
Classification of Plasticity Models: Critical Reviews …
Fig. 10.8 Uniaxial loading behavior predicted by two surface model
The uniaxial loading behavior of the two surface model is illustrated in Fig. 10.8 for the simple material without the isotropic and the kinematic hardenings. However, this model possesses the following defects in addition to the crucial defects caused by the incorporation of the elastic-domain described in Sect. 10.3.1. (1) The plastic modulus depends on the distance from the current stress to the conjugate stress irrespective of the loading process, i.e. the initial, the reverse and the reloading processes and thus the curvature of stress–strain curves are identical irrespective of these processes. Then, the Masing effect cannot be described at all contrary to the multi-surface model. (2) Only the elastic deformation without a plastic strain rate is induced in unloading process, although the plastic deformation in the opposite direction to the preceding deformation is induced in real material behavior as seen in Fig. 10.1. Consequently, the open hysteresis loop is depicted in the unloading–reloading process. (3) The excessively large mechanical ratcheting phenomenon is depicted as shown in Fig. 10.9. However, the mechanical ratcheting cannot be described at all if the inner yield surface (purely elastic domain) is assumed to be larger than the half of the outer-yield (bonding) surface. Therefore, the finitely different results
10.3
Plasticity Models with Elastic Domain
287
Bounding surface Constant stress amplitude
Elastic state
p
0
(a) Small inner yield surface: Excessively large strain accumulation
Bounding surface Constant stress amplitude
Elastic state
0
p
(b) Large inner yield surface: No strain accumulation
Fig. 10.9 Prediction of cyclic loading behavior by the two-surface model under a constant stress amplitude: excessively large strain accumulation
of the mechanical ratcheting are described by the selection of the size of the inner yield surface. Besides, the spring-back phenomenon can never be predicted by this model, since the description of the plastic deformation induced in the unloading process is of crucial importance for the prediction of this phenomenon. Nevertheless, Yoshida et al. (Yoshida and Uemori, 2002b, 2003; Yoshida and Amaishi, 2020) insist that they can predict the spring-back phenomenon by the two surface model by making the Young’s modulus decrease asymptotically to the saturation value with the plastic equivalent strain or by using the chord modulus defined by the inclination of the straight line connecting the beginning and the unloaded points of the unloading stress-strain curve, which decreases with the plastic deformation. In fact, however, the exact Young’s modulus represented by the inclination of the stress vs. strain curve at the moment just after the reverse loading in the uniaxial loading is kept constant as far as a large deformation causing the damage is not induced as known from the elementary knowledge of the continuum damage mechanics. Besides, once the damage begins, the damage develops increasingly to the failure in the monotonic loading process. In fact, however, the plastic deformation is induced during the unloading process, so that the spring-back phenomenon must be analyzed by an elastoplastic constitutive equation which is capable of describing the plastic
288
10
Classification of Plasticity Models: Critical Reviews …
deformation induced in the unloading loading process. Besides, the mechanical ratcheting phenomenon, i.e. the strain accumulation during the cyclic loading cannot be described by such easy-going method altering the elastic deformation behavior which woud result in the risky mechanical design. The spring-back phenomenon can be analyzed rigorously by the extended subloading surface model which will be explained in the next section and Sect. 12.6. Unfortunately, however, the two surface model worsened by the Yoshida’s group has been installed as the standard program in the commercial software Ls-Dyna, PAM-STSAMP and JSOL-JSTAMP. Further, the similar irrational methods have been reported by Sun and Wagoner (2011), Wagoner (2013), Chen et al. (2016), Okorokov et al. (2019), etc. One should make effort seriously to describe the plastic deformation induced in the unloading process in order to predict rationally the spring-back phenomenon. Vanishing Elastic Domain Model The modification of the two surface model was proposed by Dafalias and Popov (1977) in which the inner yield surface is contracted to a point, naming it “vanishing elastic domain”. However, this modification results in the serious defect that the direction of the plastic strain rate depends on the direction of the stress rate and thus it becomes unstable deviating from the associated flow rule. In other words, it is the quite peculiar model ignoring the physical background of the associated flow rule described in Sect. 8.6. Bounding Surface Model with Radial-Mapping The bounding surface model with radial-mapping by Dafalias and Harmann (1980) is basically different from the above-mentioned two-surface model and it is basically different also from the subloading surface model contrary to the Dafalias’ statement (Dafalias 1986) as was explained exhaustively in Sect. 9.5. The term “bounding surface” should be excluded from the formal technical terms, since it is merely a yield surface itself so that all the plasticity models using the yield surface are included in the bounding surface model if this term is accepted. As studied deeper, the inherent defects become clearer in the Chaboche, the multi and the two surface models.
10.4
Extended Subloading Surface (Hashiguchi) Model: Capable of Describing General Loading Behavior
The subloading surface model formulated in Chap. 9, called the initial subloading surface model hereinafter, is incapable of describing cyclic loading behavior appropriately, predicting an open hysteresis loop in an unloading-reloading process and thus overestimating a mechanical ratcheting phenomenon. The insufficiency is
10.4
Extended Subloading Surface (Hashiguchi) …
289
caused by the fact that the similarity-center of the normal-yield and the subloading surfaces is fixed at the back-stress and thus a purely-elastic deformation is described in the unloading process, resulting in the open hysteresis loop. Then, the insufficiency was remedied by making the similarity-center translate with the plastic deformation (Hashiguchi 1985b, 1986, 1989) as will be described in the following, while it is called the extended sublading surface model. The uniaxial loading behavior is depicted in Fig. 10.10 for the simple material behavior without a variation of the normal-yield surface. The similarity-center goes up following the stress by the plastic strain rate in the initial loading process as seen in Fig. 10.10a and b. The subloading surface shrinks and thus only elastic strain rate is 1
1
σ
Normal-yield surface
1
1
c α σ
σ
c
σ = c =α c
p
0
0
3
1
σ
c
σ c
α
Subloading surface: expands
3
c
(c = O)
σ
σ
0
0
α
p
σ c
0
Contracts p ( ε = O)
α p
0
3 2 (c) Beginning instant of unloading process
Expands
α Contracts p ( ε = O)
σ σ
σ c
α
α α
2 3 (e) Reloading process until reaching similarity-center 1
1
σ
c
c
(c = O)
p
0
1
1
α c α σ σ
0
2 (b) Initial loading process
3 1
1
α
p
0
Expands 0
2 (d) Reverse loading process
(a) Beginning of initial loading
1
σ
p
0
2
c α
2 (f) Reloading process
0
3
1
1
σ
σ c
α
0
Expands
c α p
0 2
3
Expansion of (f): Closed hysteresis loop is depicted in unloading-reloading process.
Fig. 10.10 Prediction of uniaxial loading behavior by extended subloading surface model: a initial state, b initial loading process, c unloading process until similarity-center, d unloading-inverse loading process after passing similarity-center, e reloading process until reaching similarity-center and f reloading process.
290
10
Classification of Plasticity Models: Critical Reviews …
Fig. 10.11 Modification of subloading surface model to describe cyclic loading behavior.
induced until the stress goes down to the similarity-center in the unloading process as seen in Fig. 10.10c. After that the subloading surface begins to expand and thus the plastic strain rate in the compression is induced in the unloading-inverse loading process whilst the similarity-center goes down following the stress by the compressive plastic strain rate as seen in Fig. 10.10d. Further, only the elastic strain rate is induced until the stress goes up to the similarity-center in the reloading process from the complete unloading as seen in Fig. 10.10e. After that the subloading surface begins to expand and thus the plastic strain rate is induced whilst the similarity-center goes up following the stress by the plastic strain rate as seen in Fig. 10.10f. The expanded figure of Fig. 10.10f is shown in the lowest part of Fig. 10.10. Consequently, the closed hysteresis loop is depicted realistically as shown in this figure. Besides, the spring-back phenomenon can be described rigorously only by the extended subloading model, while it can never be analyzed by such irrational model as the Yoshida et al’s model based on the two surface model (Yoshida and Uemori, 2002b, 2003; Yoshida and Amaishi, 2020) as was explained in Sect. 10.3.4. The extended subloading surface model would describe the cyclic loading behavior realistically as illustratively shown in Fig. 10.11, resolving all drawbacks in the cyclic plasticity models based on the kinematic hardening concept, while the continuity and the smoothness conditions in Eqs. (9.1) and (9.2) are satisfied only in this model. Then, it has been applied to the descriptions of rate-independent (elastoplastic) and rate-dependent (elasto-viscoplastic) deformation behavior of not only metals but also geomaterials and further the friction phenomena between solids and the crystal plasticity as will be described in detail in the subsequent chapters.
10.5
Overall Assessment of Plasticity Models
The overall assessment of cyclic plasticity models is summarized in Table 10.1. All the serious drawbacks involved in the other plasticity models with elastic domain, i.e. the cylindrical yield surface model Chaboche-Ohno), the multi-surface model
Table 10.1 Classification and over-all evaluation of plasticity models
10.5 Overall Assessment of Plasticity Models 291
292
10
Classification of Plasticity Models: Critical Reviews …
(Mroz) and the two surface model (Dafalias-Yoshida) are resolved in the extended subloading surface model. Eventually, it can be concluded definitely that the rigorous plasticity model is the extended subloading surface model which possesses the capability of describing the general rate-independent and—dependent monotonic and the cyclic loading behavior for any small stress amplitude and the finite deformation behavior by the multiplicative elastic-plastic decomposition. The detailed formulations for the extended subloading surface model will be described in the subsequent chapters.
Chapter 11
Extended Subloading Surface Model
As was assessed in detail in Chap. 10, it can be concluded that only the extended subloading surface model is capable of describing the general loading behavior of materials appropriately. The explicit formulation of the extended model is described in detail in this chapter. The central point of the extension in the formulation from the initial subloading surface model in Chap. 9 is the translation of the similarity-center of the subloading and the normal-yield surfaces, so that the cyclic loading behavior with the closed hysteresis loop is described realistically. The extended model will be applied to the description of the elastoplastic deformations of various materials, e.g. metals, soils and the various phenomena, e.g. damage, phase transformation and friction and some of their validities will be verified by comparisons with test data of metals and soils in the subsequent chapters.
11.1
Normal-Yield and Subloading Surfaces
The normal-yield surface with the isotropic and the kinematic hardening is described as f ð^ rÞ ¼ FðHÞ
ð11:1Þ
as shown in Eq. (8.71) already. The extended subloading surface associated with the normal-yield surface in Eq. (11.1) is given as follows (see Fig. 11.1). f ðrÞ ¼ RFðHÞ
ð11:2Þ
where Rð0 R 1Þ is the normal-yield ratio and
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_11
293
294
11
Extended Subloading Surface Model
n
εp
σy
σ
σ
ˆ ) = F (H ) f (σ Subloading surface
nˆ c
σ
σ
Rσ
Normal-yield surface
f (σ) = RF (H )
c
cˆ
α α
Elastic-core surface
f (cˆ) = R cF (H )
ij
0
3
n
σy
α
α
ˆ ' || = F ( H ) 3 / 2||σ
σ
cˆ
( / R) σ
Normal-yield surface
εp
σ Rσ
f (cˆ ) = F ( H )
σˆ y σ y α σ / R σˆ σ α σˆ y
(a) General material
Limit elastic-core surface
Subloading surface
σ σ
ˆ' || RF ( H ) 3 / 2 || σ' Rc
c
Elastic-core surface
0
1
ˆ' || = R cF ( H ) 3 / 2 || c
2 (b) Mises material in deviatoric stress plane
Fig. 11.1 Normal-yield, subloading and elastic-core surfaces
Limit elastic-core surface ˆ' || = F ( H ) 3 / 2 || c
11.1
Normal-Yield and Subloading Surfaces
295
rra
ð11:3Þ
a stands for the conjugate (similar) point in the subloading surface to the point a representing the back-stress tensor for kinematic hardening in the normal-yield surface. Here, note that the purely elastic behavior is induced when the stress coincides with the similarity-center of the normal-yield and the subloading surfaces and thus the normal-yield ratio becomes zero ðR ¼ 0Þ. Then, let the similarity-center be called the elastic-core and let it be denoted by the bold letter c. Here, the following relation holds by virtue of the similarity (see Fig. 11.1). a c ¼ Rða cÞ
ð11:4Þ
a ¼ c R^c
ð11:5Þ
which yields
_
r ¼ r þ R^c
ð11:6Þ
where ^c c a
) ð11:7Þ
_
rrc
Besides, ry in Fig. 11.1 is the conjugate stress on the normal-yield surface to the current stress r on the subloading surface, fulfilling ra ¼ Rðry aÞ. The time-derivative of a is described from Eq. (11.5) as
a ¼ R a þ ð1 RÞ c R ^c
ð11:8Þ
All the relations in Eqs. (11.4)–(11.6) hold by virtue of the similarity of the subloading surface to the normal-yield surface. The subloading surface in Eq. (11.2) is rewritten by Eq. (11.6) as follows: _
f ðr þ R^cÞ ¼ RFðHÞ
ð11:9Þ
Adopt the associated flow rule for the subloading surface: e p ¼ k nðk ¼ e p [ 0Þ
ð11:10Þ
296
11
Extended Subloading Surface Model
where n
@f ðrÞ @r
@f ðrÞ @r
ðknk ¼ 1Þ
ð11:11Þ
The rate of the isotropic hardening variable is described by using the function fHn ^ to n as given from Eqs. (8.23) and (8.28) with Eq. (11.10) and the replacement of n follows:
p H ðr; H; e Þ ¼ fHn ðr; H; nÞ k
ð11:12Þ
The rate of the kinematic hardening variable is described by using the function ^ to f kn given from Eqs. (8.89) and (8.100) with Eq. (11.10) and the replacement of n n as follows: Storage part
zfflfflfflfflfflfflfflfflfflfflfflffl}|fflfflfflfflfflfflfflfflfflfflfflffl{ 1 p a ¼ ck ðe p e a Þ ¼ k f kn ; bk F |fflfflfflfflfflffl ffl{zfflfflfflfflfflfflffl}
f kn ¼ ck
1 a n bk F
ð11:13Þ
Diddipative part
11.2
Evolution Rule of Elastic-Core
The rigorous evolution rule of the elastic-core is formulated in this section, which plays the central role in the description of the cyclic loading behavior leading to the closed hysteresis loop as shown in Fig. 10.1. Now, the following facts should be noticed. (1) In the physical view-point, the elastic-core at which a purely-elastic deformation is induced must not lie on the normal-yield surface at which a fully-plastic deformation is induced. In addition, it should be noted that a smooth elastic– plastic transition leading to the continuous variation of the tangent stiffness modulus tensor is not described violating the continuity condition in the large, i.e. the smoothness condition in Eq. (9.5), if the elastic-core lies on the normal-yield surface. On the other hand, the other plasticity models, i.e. the cylindrical superposed kinematic hardening model (Chaboche et al. 1979; Ohno and Wang, 1993), the multi surface model (Mroz, 1967; Iwan, 1967) and the two surface model (Dafalias and Popov, 1975; Krieg, 1975; Yoshida and Uremori, 2003) violate the smoothness condition, since the purely-elastic domain contacts to the yield surface directly. (2) In the mathematical view-point, the subloading surface is not determined uniquely if the stress coincides with the elastic-core lying on the normal-yield surface, noting the fact that R is indeterminate as known from the relation
11.2
Evolution Rule of Elastic-Core
297
0 ¼ R0 which is induced by substituting r ¼ c ¼ ry into rc ¼ Rðry cÞ based on the similarity of the subloading surface to the normal-yield surface. Consequently, the elastic-core is not allowed to approach to the normal-yield surface unlimitedly. Now, let the following elastic-core surface be introduced, which always passes through the elastic-core c and maintains a similarity to the normal-yield surface with respect to the kinematic-hardening variable a. ð11:14Þ where ℜ c designates the ratio of the size of the elastic-core surface to the normal-yield surface (see Fig. 11.1) so that let it be called the elastic-core yield ratio. Then, let it be postulated that the elastic-core can never reach the normal-yield surface designating the fully-plastic stress state so that the elastic-core does not go over the following limit elastic-core surface. f ð^cÞ ¼ vFðHÞ
ð11:15Þ
where vð\1Þ is the material constant designating the maximum value of ℜ c and the following inequality must be satisfied. f (c ˆ ) ≤ χ F ( H ), i.e. ℜ c ≤ χ
ð11:16Þ
The material-time derivative of Eq. (11.16) at the limit state that c lies on the limit elastic-core surface in Eq. (11.15) yields: ð11:17Þ Here, noting ð11:18Þ on account of the Euler’s homogeneous function f ð^cÞ in degree-one for the variable ^c, the substitution of Eq. (11.18) to Eq. (11.17) leads to ð11:19Þ The inequality (11.16) is called the enclosing condition of elastic-core and Eq. (11.19) is its rate form.
298
11
Extended Subloading Surface Model
Now, assume the equation (Hashiguchi 2018b) r F c a ^c ¼ ce e p ðrv cÞ ¼ ce e p v ^c F R
ð11:20Þ
where ce is the dimensionless material constant and rv is the conjugate stress on the limit elastic-core surfaces to the current stress r on the subloading surface, i.e. rv
v r þ a; R
rv c
v r ^c R
ð11:21Þ
which is based on the relation rv a ¼ vðry aÞ ¼ v
ra r ¼v R R
ð11:22Þ
where ry is the conjugate stress on the normal-yield surface (Fig. 11.2). Equation (11.20) means that the elastic-core translates to approach the conjugate
stress rv in the nonhardening state: F ¼ 0 and a ¼ O. It follows from Eq. (11.19) with Eq. (11.20) that
σy
(nˆc : (σ c ) 0 )
σ
σ
Rσ
nˆ c
σ c
σ
Normal-yield surface
ˆ ) = F (H ) f (σ Subloading surface f (σ) = RF (H )
cˆ
α
Limit elastic-core surface
α 0
f (cˆ ) = F ( H )
ij
σˆ y
σy α σ /R σˆ
σ α σˆ y ( / R)σ Fig. 11.2 Translation of elastic-core when it lies on the limit surface ℜ c = χ fulfilling the enclosing condition in Eq. (11.19) because of the convexity of the surface
11.2
Evolution Rule of Elastic-Core
299
ð11:23Þ
as far as the normal-yield surface satisfies the convexity condition (Appendix H), noting that @f ð^cÞ=@c which is the outward-normal of the elastic-core surface at the current elastic-core c and rv c makes an obtuse angle when c lies on the limit ^c is the norelastic-core surface, while rv lies on the limit elastic-core surface. n malized outward-normal of the elastic-core surface (see Fig. 11.2), i.e. @f ð^cÞ ^c n @c
, @f ð^cÞ ^c k¼ 1Þ @c ðkn
ð11:24Þ
Therefore, Eq. (11.20) satisfies the enclosing condition of the elastic-core so that the elastic-core can never go out from the limit elastic-core surface. Then, the evolution rule of the elastic-core is given from Eq. (11.20) by v F c ¼ ce ep r^c þ a þ ^c F R
ð11:25Þ
The second and the third terms in the right-hand side designate the influences by the kinematic and the isotropic hardening. These terms cannot be excluded in general as known from the fact that the quantity in the left-hand term in Eq. (11.19) leads to the following equation if these terms are ignored. " ! # v @f ð^cÞ @f ð^cÞ F F p p r^c a ^c ¼ : ce e : ce e ðrv cÞ a ^c F F @^c R @^c @f ð^cÞ @f ð^cÞ F ¼ ce :ðrv cÞep : a f ð^cÞ F @^c @^c
which is not necessarily negative for the softening process F \0 and thus does not fulfill the enclosing condition of the elastic-core (Eq. 11.19) in general as observed in soils exhibiting the softening (see Chap. 13). The translation of the elastic-core would have to be influenced by the variation of the isotropic hardening/softening and the kinematic hardening of the normal-yield surface in general, since its movement is limited to the interior of the normal-yield surface which expands and translates with the plastic deformation. This property is different from the evolution rule of the kinematic hardening which causes the movement of the normal-yield surface. Equation (11.25) with Eqs. (11.12) and (11.13) leads to
300
11
Extended Subloading Surface Model
v F' f 1 p Hn p c ¼ ce ep r^c þ ck ep e a þ e ^c ¼ f cn k ð11:26Þ F R bk F where f cn ce
v 1 F' fHn ^c a þ r^c þ ck n F R bk F
ð11:27Þ
Now, considering the uniaxial loading behavior of the simple material with the pffiffiffiffiffiffiffiffi p p p0 p plastic incompressibility e ¼ e leading jjje k ¼ 3=2jea j without the isotropic
hardening F ¼ 0 and the kinematic hardening a ¼ O and designating the compop p nents of c and e in the axial direction by ca and ea , respectively, it follows noting p
p
rva ¼ vF ðupper: ea [ 0; lower: ea \0Þ that
ca ¼ c e
pffiffiffiffiffiffiffiffi p 3=2ð ea ÞðvF ca Þ
ðupper : eap [ 0; lower : eap \0Þ
leading to
ca ¼
pffiffiffiffiffiffiffiffi 3=2ce ðvF ca Þ eap
ðupper : eap [ 0; lower : eap \0Þ
ð11:28Þ
from which one has 8 pffiffiffiffiffiffiffiffi > ¼0 > < 3=2ce vF for ca ( ca pffiffiffiffiffiffiffiffi 3=2ce ðvF ca Þ ¼ pffiffiffi vF and eap \0 p ¼ > ea > : 6ce vF for ca ¼ vF and eap [ 0
ð11:29Þ
The time-integration of Eq. (11.28) is given as follows: pffiffiffiffiffiffiffiffi vF ca ¼ exp½ 3=2ce ðepa epa0 Þ vF ca0 i.e. ca ¼ vF ðvF ca0 Þ exp½
pffiffiffiffiffiffiffiffi 3=2ce ðepa epa0 Þ
ðupper: eap [ 0; lower : eap \0Þ
ð11:30Þ
where ca0 and epa0 are the initial values of ca and epa , respectively. The relation of ca versus epa is shown for the incompressible material without hardening in Fig. 11.3 in which the nonlinear evolution rule is depicted actually. The plastic strain ep be additively decomposed into the storage part epcs and the dissipative part epcd for the elastic-core c, i.e.
11.2
Evolution Rule of Elastic-Core
301
Fig. 11.3 Translation of elastic-core in uniaxial loading for nonhardening Mises material
ep ¼ epcs þ epcd ;
p p ep ¼ ecs þ ecd
ð11:31Þ
Then, the elastic-core c is formulated using the elastic strain energy function wc ðepcs Þ as follows: c¼
@wc ðepcs Þ @epcs
ð11:32Þ
Here, let wc ðepcs Þ be given by the quadratic equation 1 wc ðepcs Þ ¼ cs epcs : epcs 2
ð11:33Þ
Then, it follows that
p p c ¼ cs epcs ¼ cs ðep epcd Þ; c ¼ cs ecs ¼ cs ðep ecd Þ;
ð11:34Þ
where cs is the material constant with the dimension of stress. The storage and the dissipative parts in the rate of the elastic-core in Eq. (11.26) are expressed formally noting Eq. (11.34) as follows:
ð11:35Þ
302
11
Extended Subloading Surface Model
Here, note that c e || ε p || ( / R ) σ / c s in the elastic-core itself, ck ε p / c s in the p
kinematic hardening and ( F' f Hn / F )|| ε || cˆ / c s in the isotropic hardening contribute p
to the evolution of the elastic-core. On the other hand, c e || ε ||cˆ / c s in the elastic-core itself and ck [1/ (bk F )] || ε p || α / c s in the kinematic hardening suppress the evolution of p the elastic-core. In other words, c e || ε ||cˆ / c s and ck [1/ (bk F )] || ε p || α / c s lead to the dissipative part in the rate of the elastic-core. Equations (11.13) and (11.35) will be used in the formulation of the dissipative part of the plastic strain rate for the rates of the kinematic hardening variable and the elastic-core in the multiplicative hyperelastic-based plasticity in Sect. 17.9.
11.3
Plastic Strain Rate
The material-time derivative of Eq. (11.2) leads to the consistency condition of the subloading surface as follows: @f ðrÞ @f ðrÞ :r :aRF RF ¼ 0 @r @r
ð11:36Þ
@f ðrÞ :r ¼ f ðrÞ ¼ RF @r
ð11:37Þ
Here, one has
based on the homogeneous function f ðrÞ of r in degree-one by the Euler’s theorem. Then, it follows that @f ðrÞ :r @r ¼ RF ; n :r ¼ @f ðrÞ @f ðrÞ @r @r
1 n :r @f ðrÞ ¼ RF @r
ð11:38Þ
The substitution of Eq. (11.38) into Eq. (11.36) leads to "
n : r n :
# ! F R þ rþ a ¼ 0 F R
ð11:39Þ
The substitution of Eq. (11.8) into Eq. (11.39) leads to " # F R n : r n : r þ R a þ ð1 RÞ c þ ðr R^cÞ ¼ 0 F R
ð11:40Þ
11.3
Plastic Strain Rate
303
Noting the relation _
r R^c ¼ r a ðc aÞ ¼ r
ð11:41Þ
it follows from Eq. (11.40) that F R_ n: r n : r þ R a þ ð1 RÞ c þ r ¼ 0 F R
ð11:42Þ
The substitutions of Eq. (11.10) with Eqs. (9.9), (11.12), (11.13) and (11.26) into Eq. (11.42) leads to F' U_ n : r n : k fHn r þ R k f kn þ ð1 RÞ k f cn þ k r ¼ 0 F R
ð11:43Þ p
from which the plastic multiplier k and the plastic strain rate e are given as follows:
k¼
n : r p n : r p ; e ¼ p n M M
ð11:44Þ
where
F' U fHn r þ Rf kn þ ð1 RÞf cn þ _ M ¼ n: r F R
p
ð11:45Þ
The plastic modulus in Eq. (11.45) is reduced to the one for the conventional model in the normal-yield state by setting R ¼ 1 leading to U ¼ 0.
11.4
Stain Rate Versus Stress Rate Relations
The strain rate is given by substituting Eqs. (8.5) and (11.44)2 into Eq. (8.1) as follows:
e ¼ E1 : r þ
n :r p n¼ M
n :n :r E1 þ p M
ð11:46Þ
from which the magnitude of plastic strain rate described in terms of the strain rate,
denoted by K instead of k, in the flow rule of Eq. (11.47) is given as follows:
304
11
K¼
Extended Subloading Surface Model
n:E: e n:E: e p ;e ¼ p n p M þ n:E:n M þ n:E:n
ð11:47Þ
The stress rate is given from Eq. (8.5) with Eq. (11.47) as follows: ð11:48Þ The loading criterion is given as follows (Hashiguchi 2000, 2013b):
ep ¼ 6 O for R [ Re and K [ 0 or n : E : e [ 0 p e ¼ O for others
ð11:49Þ
p
premising M þ n : E : n [ 0 based on the same physical background to that for Eq. (8.55) described in Sect. 8.5, where the judgment whether or not the stress reaches the yield surface is not required since the plastic strain rate is induced continuously as the stress approaches the normal-yield surface. Equation (11.49) is applicable not only to the hardening state but also to the perfectly-plastic and softening state (Hashiguchi 2000, 2013a).
11.5
Calculation of Normal-Yield Ratio in Unloading Process
The normal-yield ratio R is calculated by the time-integration of Eq. (9.9) in the plastic loading process ep 6¼ O. On the other hand, it can be calculated efficiently for large incremental steps by using the analytical solution in Eq. (9.15) under u ¼ const: during an incremental step. The former is used for the forward-Euler method premising on the infinitesimal incremental steps and the latter is used in the return-mapping method for large incremental steps. On the other hand, R is calculated by solving Eq. (11.9), i.e. the following equation in the elastic loading process ee 6¼ O; ep ¼ O. _
f ðr þ R^cÞ ¼ RFðHÞ
ð11:50Þ
by substituting the updated values of r; a; c; F. Equation (11.50) is expressed by the quadratic equation and thus the normal-yield ratio can be expressed analytically for the Mises material as will be shown in Chap. 12. For soil plasticity, however, Eq. (11.50) is expressed by higher-order nonlinear equation and thus the normal-yield ratio must be calculated numerically for soils as will be described in Chap. 13.
11.5
11.6
Calculation of Normal-Yield Ratio in Unloading Process
305
Improvement of Inverse and Reloading Responses
The unique relation ep ep0 ¼ f ðR R0 Þ holds as shown in Eq. (9.15) under the initial condition ep ¼ ep0 : R ¼ R0 in the monotonic loading process if U in Eq. (9.9) is the function of only the normal-yield ratio R. Therefore, ep induced during a certain change of R in the monotonic loading process is identical irrespective of the difference of loading processes, e.g. the initial loading, the reloading and the inverse loading and of the proportional and non-proportional loadings. This property causes the description that the returning of the reloading stress–strain curve to the previous loading curve is unrealistically gentle as shown in Fig. 11.4a. Therefore, it engenders the impertinent prediction of cyclic loading behavior, i.e. the prediction of the unrealistically large plastic strain accumulation during the
Fig. 11.4 The defect of past subloading surface model: Unrealistically gentle returning to preceding loading curve
306
11
Extended Subloading Surface Model
cyclic loading process as shown in the upper part of Fig. 11.4b. This insufficiency in the past formulation of the subloading surface model was criticized intensely by Dafalias, stating “the predictions reported in Hashiguchi (1980) for the uniaxial loading of metals were quite unrealistic, basically due to the strong undershooting phenomenon” (Dafalias 1986, p. 980). However, the insufficiency in the description of deformation behavior by the past formulation of the subloading surface model would not originate from the intrinsic nature of this model contrary to the criticism by Dafalias (1986), although the identical defect is involved in the Chaboche model as the unavoidable (intrinclic) nature as described in Sect. 10.3.2. In what follows, the past formulation of subloading surface model will be modified so as to remedy the insufficiency in the description of reloading behavior. First, note the following facts: (1) The difference between the curvatures in the loading and the inverse loading curves becomes larger as the plastic deformation proceeds continuously. This fact is known as the Masing rule (Masing 1926). (2) The similarity-center corresponding to the most elastic stress state approaches the normal-yield surface as the plastic deformation proceeds continuously, and the approaching degree of the similarity-center to the normal-yield surface is expressed by the elastic-core yield ratio ℜ c ( = f (cˆ ) / F ) in Eq. (11.14) as described in Sect. 11.2. (3) The transition from the elastic to plastic state is more abrupt, i.e. the curvature of stress–strain curve is greater for a larger value of the material parameter u in the function UðRÞ in Eq. (9.14), as described in Sect. 9.2. The faster and slower rising up of the stress vs. strain curve can be described by setting a larger value and a smaller value, respectively, to the material parameter u. (4) The larger difference between the values of u in the reloading and the inverse loading states can be described by the larger value of ℜ c. ^c (5) The direction n of plastic strain rate is close to that of the outward-normal n (Eq. 11.24) of the elastic-core surface in the reloading process but it is far from ^c in the unloading process. The degree how near the or even opposite to that of n reloading process can be expressed by the following scalar product of these unit tensors: ^c : n Cn n
ð1 Cn 1Þ
ð11:51Þ
^c involved in the variable C n are shown in where the unit normal tensors n and n Fig. 11.5. Eventually, let the material parameter u in the function U(R) for the evolution rule of the normal-yield ratio R in Eq. (9.14) be extended to the following equation for the extended subloading surface model in which the elastic-core, i.e. the similarity-center c moves with the plastic deformation, introducing the variables ℜ c and Cn .
11.6
Improvement of Inverse and Reloading Responses
n
nˆ c
307
Normal-yield surface
ˆ ) F (H ) f (σ σ
c α α
σ σ
Subloading surface f (σ ) RF ( H ) Elastic-core surface f (cˆ ) / F ( H ) c
ij
^c : n ð1 Cn 1Þ for improvement of reloading behavior Fig. 11.5 Scalar variable Cn n
⎧u exp(uc χ ) (largest) for ℜ c = χ and Cn = 1 ⎪ u → u exp(ucℜ c Cn ) = ⎨ (medium) for ℜ c = 0 or Cn = 0 u ⎪u exp(−u χ ) (smallest) for ℜ = χ and C = −1 n c c ⎩
ð11:52Þ
leading to the replacement 〈 R − Re 〉 ) → U ( R, ℜc , Cn ) = u exp(ucℜc Cn ) cot(π2 〈 R1 −−RRee〉 ) U ( R) = u cot(π 2 1 − Re
ð11:53Þ
where uc is the material constant. Then, the plastic positive proportionality factor k is re-placed to the following equation by Eq. (11.53), noting Eq. (9.9) with Eq. (9.14).
ð11:54Þ By the above-mentioned modification, the phenomenon that the reloading curve after a partial unloading returns rapidly to the preceding loading curve and the curvature of inverse loading curve decreases can be described realistically as shown in Fig. 11.6a. Besides, the excessive plastic strain accumulation for the cyclic loading process in the neighborhood of yield surface is suppressed as shown in the lower part of Fig. 11.6b. The above-mentioned improvement of the reloading behavior is inevitably required for the description of the accurate cyclic loading behavior. On the other hand, the improvement of the cyclic loading behavior can never be attained by the modification of the dynamic recovery part of the Armstrong-Frederick kinematic hardening rule as has been repeated after Chaboche et al. (1979), since it is relevant to the reverse loading behavior but irrelevant to the reloading behavior.
308
11
uc 0
Extended Subloading Surface Model
R =1
uc = 0
uc 0
c
uc = 0
c
uc 0
0 uc = 0 uc 0
uc = 0
c
p
R =1
(a) Reloading after partial-unloading and reverse loading
R =1
Elastic-core
0
p
(b) Cyclic loading under small stress amplitude near normal-yield state Fig. 11.6 Stress-plastic strain curve predicted by modified evolution rule of normal-yield ratio: Rapid recovery to preceding monotonic loading curve
11.7
Loading Criterion for Large Loading Increment
The exact loading criterion will be given in this section, which is required for the numerical calculation with large loading increments involving not only the monotonic but also the inverse loading.
11.7
Loading Criterion for Large Loading Increment
11.7.1
309
Exact Judgment of Loading
The following loading criterion for the return-mapping method was incorporated by Hashiguchi (2013, 2017a) and has been used for the FEM analyses by Hashiguchi (2013), Anjiki et al. (2016), Yamakawa et al. (2010a, b), Iguchi (2019), Fincato and Tsutsumi (2017, 2018, 2019a, b, 2020a, b) and Tsutsumi et al. (2019). (
trial trial trial Depn þ 1 ¼ O; rFinal n þ 1 ¼ rn þ 1 for f ðrn þ 1 an Þ Rn FðHn Þ 0 or f ðrn þ 1 an Þ Re FðHn Þ 0 trial Depn þ 1 6¼ O; rFinal n þ 1 6¼ rn þ 1 for other
ð11:55Þ which persists that the plastic strain rate is induced only when the subloading surface just after the elastic trial step is larger than subloading surface at the beginning of the elastic trial step, i.e. the previous step. This standard loading criterion gives the correct judgment for conventional plasticity models. However, when this loading criterion is applied to the subloading surface model, it may give incorrect judgment due to the following facts. (1) The subloading surface in the elastic-trial step cannot be described by f ðrtrial nþ1 trial trial trial an Þ Rn FðHn Þ ¼ 0 but can be described by f ðrn þ 1 an þ 1 Þ Rn þ 1 FðHn Þ ¼ trial 0 with atrial cn exactly, noting that only the internal variables Hn , n þ 1 ¼ cn Rn þ 1^ an and cn are fixed in the elastic trial step. (2) Any elastic trial steps inducing the shrink of the subloading surface are judged as the elastic loading process by the past (incorrect) loading criterion. In fact, however, the plastic strain rate must be induced when once the subloading surface contracts but then it expands over the elastic domain in the elastic trial step even if the subloading surface after the elastic trial step is smaller than the subloading surface in the beginning of the elastic trial step. Needless to say, such incorrect loading judgment would result in the accumulation of plastic strain rate during the cyclic loading for a constant or a decreasing amplitude cannot be described at all. Then, the exact loading criterion will be formulated in the following. The following facts should be noticed for the formulation of the loading criterion required at the beginning of the plastic corrector step, referring to Fig. 11.7, where the elastic trial step for the initial subloading surface model ðc ¼ aÞ and for the extended subloading surface model ðc 6¼ aÞ are shown in Fig. 11.7a and b, respectively (The subscript n and n þ 1 is added to the variables at the end of the step n and the step n þ 1, respectively). “The plastic strain increment Depn þ 1 is not induced if the elastic trial stress rtrial nþ1 stays inside the elastic response region, i.e. f ðrtrial n þ 1 Þ Re FðHn Þ\0 in the step n or trial if the stress increment Drtrial n þ 1 makes an obtuse angle with the outward-normal nn þ 1 of the subloading surface in the elastic trial step n þ 1. Otherwise, it is induced.”
310
11
Extended Subloading Surface Model
Fig. 11.7 Correct loading criterion when elastic trial stress increment is directed inward of subloading surface at step n in return-mapping method for subloading surface model
11.7
Loading Criterion for Large Loading Increment
311
Then, the correct loading criterion in the return-mapping method for the subloading surface model is given as follows: Loading criterion in returnmapping for subloading surface model trial trial trial trial trial Depn þ 1 ¼ O and rFinal n þ 1 ¼ rn þ 1 for Rn þ 1 Rn or Rn þ 1 Re or nn þ 1 :Drn þ 1 0 p Final trial Den þ 1 6¼ O and rn þ 1 6¼ rn þ 1 for others ð11:56Þ where trial etrial rtrial n þ 1 ¼ rn þ Drn þ 1 ¼ rn þ E :Den þ 1
ð11:57Þ
@f ðrtrial n þ 1 Þ @f ðrtrial Þ nþ1 ntrial nþ1 @rtrial @rtrial nþ1 nþ1
ð11:58Þ
trial trial trial trial rtrial cn Þ n þ 1 ¼ rn þ 1 a n þ 1 ¼ rn þ 1 ðcn Rn þ 1^
ð11:59Þ
where
noting Eq. (11.5). Rtrial n þ 1 is calculated from the following equation of the subloading surface in the elastic trial step. trial f ðrtrial n þ 1 Þ ¼ Rn þ 1 FðHn Þ
ð11:60Þ
Equation (11.60) is the nonlinear equation of the normal-yield ratio Rtrial n þ 1 and thus must be calculated by the numerical method, e.g. the Newton–Raphson Rtrial nþ1 method in general. However, the function f ðrtrial n þ 1 Þ is given for the Mises material by f ðrtrial n þ 1Þ ¼
pffiffiffiffiffiffiffiffi trial pffiffiffiffiffiffiffiffi trial' trial ' 3=2jjrn þ'1 a n þ 1 k ¼ 3=2rtrial cn Þ' n þ 1 ðcn Rn þ 1^
ð11:61Þ
Rtrial n þ 1 is calculated analytically by solving the quadratic equation derived by substituting Eq. (11.61) into Eq. (11.60) as follows:
Rtrial nþ1 ¼
' ' ðrtrial c'n þ n þ 1 cn Þ: ^
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2ffi ' ' ' ' ½ðrtrial c'n 2 þ ð23 ½FðHn Þ2 ^c'n Þrtrial n þ 1 cn Þ: ^ n þ 1 cn ' 2 2 2 ^ 3 ½FðHn Þ cn
ð11:62Þ The detailed derivation of Eq. (11.62) can be referred to Sect. 12.3. Further, ntrial nþ1 is given for the Mises material as follows:
312
11
Extended Subloading Surface Model
trial ' ' ' rtrial cn Þ n þ 1 ðcn Rn þ 1^ ntrial nþ1 trial trial ' ' r ðc' R ^c' Þ nþ1
11.7.2
n
ð11:63Þ
nþ1 n
Initial Value of Normal-Yield Ratio in Plastic Corrector Step
As known from the correct loading criterion described in the latest section, the initial value of normal-yield ratio must be determined at the beginning of the plastic-corrector step. It has been formulated and derived by Hashiguchi (2018, 2020) as explained below. Noting the loading criterion described in the foregoing, the plastic corrector step must be started from (1) the subloading surface coinciding with the elastic domain surface, when the elastic-trial stress increment intersects with the elastic domain surface, for which the normal-yield ratio is given by Re . (2) the subloading surface in the transitional state at which the subloading surface once contracts and then begins to expand in the elastic-trial step, while the elastic-trial stress increment Drtrial n þ 1 contacts tangentially to the subloading surface, for which the normal-yield ratio is designated as R0n þ 1 . Then, designating the stress at which the stress increment vector Drtrial n þ 1 becomes tangential to the subloading surface which changes from the contraction to the expansion by r0n þ 1 , the following relations must be satisfied at the contact point, i.e. the loading-start stress r0n þ 1 . f ðr0n þ 1 Þ ¼ R0n þ 1 FðHn Þ n0n þ 1 : Drtrial nþ1 ¼ 0
ð11:64Þ
The first equation is required for the subloading surface to pass through the contact point r0n þ 1 and the second equation is required for the elastic-trial increment Drtrial nþ1 to contact tangentially to the subloading surface at the contraction–expansion transition. r0n þ 1 sDrtrial n þ 1 þ rn ð0 s 1Þ
ð11:65Þ
a0n þ 1 ¼ cn R0n þ 1^cn
ð11:66Þ _
0 r0n þ 1 r0n þ 1 a0n þ 1 ¼ sDrtrial cn n þ 1 þ rn þ Rn þ 1^
ð11:67Þ
11.7
Loading Criterion for Large Loading Increment
n0n þ 1
@f ðr0n þ 1 Þ @r0n þ 1
313
, @f ðr0n þ 1 Þ @r0 nþ1
ð11:68Þ
sð0 s 1Þ is the unknown scalar parameter which must be determined so as to satisfy Eq. (11.64) at the stress r0n þ 1 . The two unknown variables s and R0n þ 1 are calculated by solving Eq. (11.64) and then all the variables r0n þ 1 , a0n þ 1 , r0n þ 1 in the subloading surface passing through the contact point and the stress increment Drtrial n þ 1 at the elastic trial step are calculated. Equation (11.64) is regarded as the simultaneous quadratic equation of the unknown scalar variables R0n þ 1 and s so that the numerical calculation is required for their solutions in general. In what follows, the analytical equation of R0n þ 1 in terms of the known variables will be derived for the Mises material. Equation (11.64) is explicitly described for the Mises material with f ðr0n' þ 1 Þ ¼ pffiffiffiffiffiffiffiffi 0 3=2rn' þ 1 leading to n0n þ 1 ¼ r0n' þ 1 =r0n' þ 1 as follows: pffiffiffiffiffiffiffiffi 0 3=2rn' þ 1 ¼ R0n þ 1 FðHn Þ 0' rn þ 1 :Drtrial nþ1 ¼ 0
ð11:69Þ
i.e. 8 pffiffiffiffiffiffiffiffi _' 0 ' ' þr < 3=2 ^ þ R c sDrtrial ¼ R0n þ 1 FðHn Þ n n n þ 1 nþ1 :
_
0 ' ' ðsDrtrial c'n Þ:Drtrial nþ1 ¼ 0 n þ 1 þ rn þ Rn þ 1^
ð11:70Þ
The upper equation in Eq. (11.70) is expressed as _
_
0 0 ' ' ' ' ð3=2ÞðsDrtrial c'n Þ:ðsDrtrial c'n Þ n þ 1 þ rn þ Rn þ 1^ n þ 1 þ rn þ Rn þ 1^
¼ ðR0n þ 1 FðHn ÞÞ2 leading to _
trial' trial' 0 ' ' c'n Þ s2 Drtrial n þ 1 : Drn þ 1 þ 2sDrn þ 1 : ðrn þ Rn þ 1^ _
_
þ ðrn' þ R0n þ 1^c'n Þ : ðrn' þ R0n þ 1^c'n Þ ð2=3ÞðR0n þ 1 FðHn ÞÞ2 ¼ 0 from which we have _
0 ' ' c'n Þ þ s ¼ ½Drtrial n þ 1 :ðrn þ Rn þ 1^ _
_' pffi trial' 0 ' ' f½Drtrial c'n Þ2 Drtrial n þ 1 :ðrn þ Rn þ 1^ n þ 1 :Drn þ 1
_
trial' ' ½ðrn' þ R0n þ 1^c'n Þ:ðrn' þ R0n þ 1^c'n Þ ð2=3ÞðR0n þ 1 FðHn ÞÞ2 g=ðDrtrial n þ 1 :Drn þ 1 Þ
ð11:71Þ
314
11
Extended Subloading Surface Model
The substitution of Eq. (11.71) into the second equation in Eq. (11.70)2 leads to _'
0 ' Drtrial c'n Þ þ 1 :ðrn þ Rn þ 1^ qnffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi ' _ _' trial' _' 0 0 ' ' þ ½Drtrial c'n Þ2 Drtrial c'n Þ:ðrn þ R0n þ 1^c'n Þ ð2=3ÞðR0n þ 1 FðHn ÞÞ2 n þ 1 :ðrn þ Rn þ 1^ n þ 1 :Drn þ 1 ½ðrn þ Rn þ 1^ ' _
0 ' þ Drtrial c'n Þ ¼ 0 n þ 1 :ðrn þ Rn þ 1^
i.e. _
0 ' ' c'n Þ2 ½Drtrial n þ 1 :ðrn þ Rn þ 1^ _
_
trial' 0 ' ' Drtrial c'n Þ:ðrn' þ R0n þ 1^c'n Þ ð2=3ÞðR0n þ 1 FðHn ÞÞ2 ¼ 0 n þ 1 :Drn þ 1 ½ðrn þ Rn þ 1^
ð11:72Þ which is the quadratic equation of R0n þ 1 . Equation (11.72) is rewritten as trial' 02 ' ' ' c' Þ2 ðDrtrial' :Drtrial' Þð^ ' cn Þ þ ð2=3ÞðFðHn ÞÞ2 ðDrtrial ½ðDrtrial n n þ 1 :^ nþ1 n þ 1 cn :^ n þ 1 :Drn þ 1 ÞRn þ 1 _
_
trial' ' trial' ' ' c' ÞR0 ' ' þ 2½ðDrtrial cn Þ ðDrtrial n nþ1 n þ 1 :rn ÞðDrn þ 1 :^ n þ 1 :Drn þ 1 Þðrn :^ _
_
_
trial' trial' ' 2 ' ' ' þ ðDrtrial n þ 1 :rn Þ ðDrn þ 1 :Drn þ 1 Þðrn :rn Þ ¼ 0
The solution of R0n þ 1 in Eq. (11.72) is given by R0n þ 1
¼
B
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi B2 AC A
ð11:73Þ
where 8 > A S2ca Sss ½^c'n :^c'n ð2=3ÞðFðHn ÞÞ2 > < _ B Ssc Sca Sss rn' :^c'n > > _ _ : C S2sc Sss rn' : rn'
ð11:74Þ
with trial' trial' ' ' Sss Drtrial cn ; n þ 1 :Drn þ 1 ; Sca Drn þ 1 :^
_
' ' Ssc Drtrial n þ 1 :rn
ð11:75Þ
One must choose the solution satisfying 0 R0n þ 1 1 in Eq. (11.73). Here, we must set R0n þ 1 ¼ Re if R0n þ 1 Re in the calculated result. It is enough to calculate the initial normal-yield ratio R0n þ 1 , while it is not necessary to calculate the scalar number s and the stress r0n þ 1 for the return-mapping calculation. The correct loading criterion for the return-mapping method in the initial subloading surface model is given by setting c ¼ a in the above-mentioned formulations, referring to Fig. 11.7a.
11.8
11.8
Plastic Spin
315
Plastic Spin
The plastic spin in Eq. (3.44) for the hypoelastic-based plasticity is given by extending Eq. (8.94) to the extended subloading surface model as follows:
wp ¼ gp ant½r dp ¼ gp k xp ;
11.9
xp ¼ ant½r n
ð11:76Þ
Incorporation of Tangential-Inelastic Strain Rate
The tangential-inelastic strain rate described in Sect. 9.7 is extended for the extended subloading surface model as shown in this section. '
The tangential-deviatoric stress rt (Fig. 11.8) is defined by ð11:77Þ
ð11:78Þ ð11:79Þ fulfilling ð11:80Þ
by virtue of
'
The fourth-order tensor Tt is the tangential-deviatoric projection tensor which transforms an arbitrary second-order tensor to the tangential part to the subloading surface in the deviatoric stress space and the second-order tensor subjected to this ' ' projection is designated as ð Þ't , i.e. t't Tt : t leading further to Tt : t't ¼ t't .
316
11
Extended Subloading Surface Model
n'
σ '
σ 'n
σ' σ'
σ'
σ 't
ε t =
T (R ) 2G σ't
c'
α' α'
Subloading surface f (σ) = RF ( H )
'ij
0
Normal-yield surface f (σˆ ) = F ( H )
Fig. 11.8 Tangential-deviatoric stress rate and tangential-inelastic strain rate in the deviatoric stress plane
All the relations in Eqs. (9.53), (9.62) and (9.63) hold under the replacements of p ð Þ't ! ð Þ't ; M p ! M . Consequently, it follows that t
n n TðRÞ ' :r E1 þ þ T p 2G t M
ð11:82Þ
r¼
ð11:81Þ
e¼
'
e ¼ TðRÞE1 : rt ¼ TðRÞE1 : T't : r
E:n n:E TðRÞ ' 2G E p Tt : e 1 þ TðRÞ M þn : E : n
ð11:83Þ
The loading criterion in Eq. (11.49) which was given for the equation without the tangential-inelastic strain rate holds as it is, as was described in Sect. 9.7. The importance for the introduction of the tangential-inelastic strain rate will be verified by comparison with test data on a metal in Sect. 12.5. The subloading surface model has been applied to metals (Hashiguchi 1980; 1989; Hashiguchi and Yoshimaru 1995; Hashiguchi and Tsutsumi 2001; Hashiguchi and Protasov 2004; Khojastehpor et al. 2006; Tsutsumi et al. 2001, 2003, 2005, 2007; Hashiguchi et al. 2012; Fincato and Tsutsumi 2017, 2018, 2019, 2020a, b; 2021; Anjiki et al. 2020; Anjiki and Hashiguchi 2021, Liu et al. 2022) and soils (Hashiguchi and Ueno 1977; Hashiguchi 1978; Topolnicki 1990; Kohgo et al. 1993; Asaoka et al. 1997; Hashiguchi and Chen 1998; Chowdhury et al. 1999;
11.9
Incorporation of Tangential-Inelastic Strain Rate
317
Hashiguchi et al. 2002; Khojastehpor and Hashiguchi 2004a, b; Khojastehpor et al. 2006; Nakai and Hinokio 2004; Hashiguchi and Tsutsumi 2001, 2003, 2005, 2007; Hashiguchi and Mase 2007, 2011; Wongsaroj et al. 2007; Yuanming et al. 2009, 2016; Maranha et al. 2016; Hashiguchi et al. 2021; Yamada et al. 2022; Lu et al. 2022), rocks (Xiong et al. 2018), asphalt (Darabai et al. 2012), etc. Consequently, its capability has been verified widely. The explicit constitutive equations for metals and soils will be described in the subsequent chapters.
Chapter 12
Constitutive Equations of Metals
The plasticity theory has been highly developed through the prediction of deformation of metals up to date. The reason for this would lie in the fact that, among various materials exhibiting plastic deformation, metals are used most widely as engineering materials and exhibit the simplest plastic deformation behavior without a pressure dependence, a plastic compressibility, a dependence on the third invariant of deviatoric stress and a softening. Nevertheless, metals exhibit various particular aspects, e.g., the kinematic hardening and the stagnation of isotropic hardening in a cyclic loading. Explicit constitutive equations of metals will be delineated in this chapter, which are based on the extended subloading surface model described in the preceding chapters. The constitutive equation detailed in this chapter has been implemented to the commercial FEM software “Marc” marketed from MSC Software Ltd. as the standard uploaded (ready-made) program by the name “Hashiguchi model” after the 2017 version, so that Marc users can apply these models to their deformation analyses. Further, the performance that the material parameters can be determined automatically by inputting the stress–strain curve is installed after the 2019 version, so that even the users unfamiliar to the elastoplasticity theory can execute an accurate analysis by the subloading surface model. Incidentally, the computer program is shown in Appendix L(a).
12.1
Yield Surface, Isotropic, Kinematic Hardening and Elastic-Core
The yield function for the Mises yield condition is extended to incorporate kine^0 in Eq. (8.41) as follows: matic hardening by replacing r0 to r rffiffiffi 3 0 ^k f ð^ rÞ ¼ kr 2 © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_12
ð12:1Þ
319
320
12
Constitutive Equations of Metals
for which the subloading function f ðrÞ for Eq. (12.1) is given by ð12:2Þ
Further, the elastic-core function in Eq. (11.14) for Eq. (12.1) is given by rffiffiffi rffiffiffi 3 0 @f ð^cÞ 3 ^c0 ¼ f ð^cÞ ¼ k^c k ; 2 @^c 2 k^c0 k
ð12:3Þ
It follows from Eqs. (11.14) and (11.24) for Eq. (12.3) that ð12:4Þ
The isotropic hardening function is given by Eq. (8.42), i.e. FðHÞ ¼ F0 f1 þ Sr ½1 expðcH H Þg:
ð12:5Þ
F 0 ¼ F0 Sr cH expðcH H Þ
ð12:6Þ ð12:7Þ
The rates of the nonlinear kinematic hardening and the elastic-core are given by Eqs. (11.13) and (11.26) as ð12:8Þ
ð12:9Þ
The plastic modulus is given by substituting Eqs. (12.2), (12.3), (12.7), (12.8) and (12.9) into Eq. (11.45) as follows:
12.1
Yield Surface, Isotropic, Kinematic Hardening and Elastic-Core
321
(rffiffiffi 2 F0 r0 1 a r þ ck R 3F kr0 k bk F " ) pffiffiffiffiffiffiffiffi 0 # v 2=3F r0 1 U_ ^c þ r þ ð1 RÞ ce r ^c þ ck a þ F R R kr0 k bk F
r0 M : kr0 k p
ð12:10Þ
12.2
Cyclic Stagnation of Isotropic Hardening
As a plastic deformation induced by the mutual slips of solid particles proceeds, the mutual slips becomes difficult due to the closer packaging of solid particles, the accumulation of obstacles, etc. resulting in the isotropic hardening. However, if the loading direction is reversed, the solid particles are released from the close packaging, the obstacles, etc. so that the mutual slips recover so that the isotropic hardening stagnates temporarily. This phenomenon considerably affects the cyclic loading behavior in which the reverse loading is repeated. To describe this phenomenon, the concept of the cyclic stagnation of isotropic hardening, i.e. the existence of nonhardening region was proposed by Chaboche et al. (1979; see also Chaboche, 1989) and studied also by Ohno’s group (Ohno 1982; Ohno and Kachi 1986; Ohno et al. 2021). The concept insists that isotropic hardening does not proceed when the plastic strain given by the time-integration of plastic strain rate lies inside a certain region, called the nonhardening region, in the plastic strain space. The non-hardening region expands and translates when the plastic strain lies on the boundary of the region and the plastic strain rate is induced directing outwards the region. It is similar to the notion of the yield surface based on the assumption that the plastic strain rate is induced only when the stress lies on that surface, while the plastic strain and the rate of isotropic hardening variable for the nonhardening region correspond to the stress and the plastic strain rate, respectively, for the yield surface. The extended formulation for the isotropic hardening stagnation by the notion of the subloading surface concept will be shown in this section (Hashiguchi 2009; 2013a). Assuming that the isotropic hardening stagnates when the plastic strain ep lies inside a certain region, let the following surface, called the normal-isotropic hardening surface, be introduced. ~ gð~ep Þ ¼ K
ð12:11Þ
322
12
n~ p
p
ρ ε~ p
Constitutive Equations of Metals
Normal-isotropic hardening surface g (ε~ p ) = K~ Sub-isotropic hardening surface ~ ~ g (ε~ p ) = RK
ρ p
0
ij
Fig. 12.1 Normal- and sub-isotropic hardening surfaces
where ~ep ep q
ð12:12Þ
~ and q designate the size and the center, respectively, of the normal-isotropic K hardening surface, the evolution rules of which will be formulated later. Furthermore, we introduce the surface, called the subloading-isotropic hardening surface, which always passes through the current plastic strain ep and which has a similar shape and a same orientation to the normal-isotropic hardening surface (see Fig. 12.1). Here, the plastic strain ep is regarded as an internal variable. The same situation can be reminded in the Prager’s lineal kinematic hardening rule in Eq. (8.86). The subloading-isotropic stagnation surface is expressed by the following equation. ~K ~ gð~ep Þ ¼ R
ð12:13Þ
~ 0R ~ 1 is the ratio of the size of subloading-isotropic hardening surwhere R face to that of the normal-isotropic hardening surface. It plays the role as the measure for the approaching degree of the plastic strain to the normal-isotropic ~ is referred to as the normal-isotropic hardening ratio. It hardening surface. Then, R ~ in terms of the known values of ep , q ~ ¼ gð~ep Þ=K is calculable from the equation R ~ and K.
12.2
Cyclic Stagnation of Isotropic Hardening
323
The consistency condition of the subloading isotropic hardening surface is given by @gð~e p Þ p @gð~e p Þ ~ ~ ~ ~ K: :e : q ¼ RK þ R p p @~e @~e
ð12:14Þ
Let the following postulates be adopted for the formulations of the evolution ~ and q. rules of K ~ and q evolve when the plastic strain rate e p is induced directing outwards the (1) K subloading-isotropic hardening surface, fulfilling @gð~ep Þ p : e [ 0: @~ep
ð12:15Þ
~ and q increase as the plastic strain approaches the (2) The rates of K normal-isotropic hardening surface, i.e. as the normal-isotropic hardening ratio ~ increases. Therefore, they are monotonic-increasing function of the K ~ normal-isotropic hardening ratio R. p (3) The plastic strain e is assumed to exist inside the normal-isotropic hardening surface. Therefore, it must hold that
~¼1 ~ ¼ 0 for R R
ð12:16Þ
(4) The consistency condition in Eq. (12.14) is reduced to the following relation which must be fulfilled when the plastic strain just lies on the normal-isotropic hardening surface. @gð~ep Þ p @gð~ep Þ ~ : e : q ¼ K @~ep @~ep
for
~¼1 R
ð12:17Þ
Then, we assume the following equations so as to fulfill all these postulates. ð12:18Þ ð12:19Þ where 0 C 1 and fð 1Þ are the material constants and ð12:20Þ
324
12
Constitutive Equations of Metals
~
R
1 ~ K
g (ε~ p ) ~
εp
: εp
n~ : ε p 0 0
1
ε p O R~
Fig. 12.2 Evolution of normal-isotropic hardening ratio: Plastic strain is attracted to the normal-isotropic hardening surface
~ and q into Substituting Eqs. (12.18) and (12.19) for the evolution rules of K Eq. (12.14), the rate of the normal-isotropic hardening ratio is given by
ð12:21Þ
~ fulfilling which is the monotonically-decreasing function of R 8 > 1 @gð~ep Þ p > > ~¼0 h ¼ : e ið[ 0Þ for R > > ~ > p K > @~ e > > > < p ~ \ 1 h@gð~e Þ : e p ið[ 0Þ for R\1 ~ R > ~ p > K > @~ e > > > ~¼1 > ¼ 0 for R > > > : ~[1 \0 for R
ð12:22Þ
12.2
Cyclic Stagnation of Isotropic Hardening
325
~ increases as shown in Fig. 12.2. Therefore, the normal-isotropic hardening ratio R when the plastic strain moves to the outward of the sub-isotropic hardening surface but it decreases such that the normal-isotropic hardening surface encloses the plastic strain when the plastic strain tends to go out from the normal-isotropic hardening
~ [ 1 as shown in ~ \0 for R surface encloses by virtue of the inequality R Eq. (12.22). In other words, the normal-isotropic hardening surface is automatically controlled such that the plastic strain does not go out from that surface. Furthermore, needless to say, the judgment of whether the plastic strain reaches the normal-isotropic hardening surface is not necessary in the present formulation consistently based on the subloading concept in all aspects. In contrast, the judgment whether the plastic strain or the back-stress reaches the isotropic hardening (stagnation) surface is required in the other models (Chaboche et al., 1979; Chaboche, 1991; Ohno, 1982; Yoshida and Uemori, 2002b). It is assumed that the isotropic hardening variable H evolves under the following conditions.
(1) The isotropic hardening is induced when the plastic strain rate ep is induced directing outwards the sub-isotropic hardening surface, i.e.
H
n
~ : ep [ 0 [ 0 for n ¼ 0 for others:
ð12:23Þ
(2) The isotropic hardening rate increases as the plastic strain approaches the normal-isotropic hardening surface, i.e. as the normal-isotropic hardening ratio ~ increases. Then, H is the monotonically-increasing function of R. ~ Here, in R order that the isotropic hardening develops continuously, its rate must be zero, ~ ¼ 0, i.e. when the plastic strain lies just on the center of the i.e. H ¼ 0 for R normal-isotropic hardening surface because the rate is zero during the process in which the plastic strain moves towards the inside of the sub-isotropic stagnation surface. (3) The isotropic hardening rule of Eq. (12.7) in the monotonic loading process holds ~ ¼ 1Þ and when the plastic strain lies on the normal-isotropic hardening surface ðR the plastic strain rate is induced in the outward-direction of that surface. Eventually, let the following evolution rule of isotropic hardening be assumed by extending Eq. (12.7). ð12:24Þ
326
12
Constitutive Equations of Metals
where t ð = 1Þ is the material constant and ~fHn R ~ t h~ ifHn n:n
ð12:25Þ
The three material constants C, 1 and t in the isotropic hardening stagnation can be fixed as C = 0.5, 1 = 3 and t = 3 actually. Employing the extended isotropic hardening rule in Eqs. (12.24) instead of Eqs. (11.12) into Eq. (11.45), the plastic modulus is modified as follows: (
ð12:26Þ
in general and (
ð12:27Þ
for metals with Eq. (12.7). The function gð~ep Þ is given in the simplest form as follows: gð~ep Þ ¼ jj~ep jj;
ð12:28Þ
~ep ~ep @gð~ep Þ ~ ; n ¼ ¼ jj~ep jj jj~ep jj @~ep
ð12:29Þ
The Yoshida’s group (Yoshida and Uemori, 2002a, b; Yoshida and Amaishi, 2020) proposed the isotropic hardening stagnation by incorporating the Armstrong-Frederick (1966)’s nonlinear kinematic hardening variable (back-stress) instead of the plastic strain. However, it is to be physical worsening and the mathematical complication, since the physical mechanism of the isotropic hardening is different from that of the kinematic hardening.
12.3
Calculation of Normal-Yield Ratio in Unloading Process
The normal-yield ratio R must be calculated from the equation of the subloading surface in the unloading process (ep ¼ O). It can be calculated directly by R ¼ f ð^ rÞ=F in the initial subloading surface model. However, it has to be calculated by solving the equation of the subloading surface in the extended subloading surface model as described below. Substituting Eq. (11.6) into Eq. (12.2), the extended subloading surface is described as follows:
12.3
Calculation of Normal-Yield Ratio Unloading Process
327
rffiffiffi 3 _0 jjr þ R^c0 jj ¼ RFðHÞ 2
ð12:30Þ
2 _ trðr0 þ R^c0 Þ2 ¼ R2 F 2 3
ð12:31Þ
i.e.
The normal-yield ratio R is derived from the quadratic Eq. (12.31) as follows: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 _ _ 2 2 0 0 0 0 0 F jj^c jj jjr0 jj2 r : ^c þ ðr : ^c Þ þ 3 R¼ 2 2 F jj^c0 jj2 3 _
12.4
ð12:32Þ
Implicit Stress-Integration
The implicit stress-integration method for the extended subloading surface model without the isotropic hardening stagnation and the tangential-inelastic strain rate was given by Anjiki et al. (2016, 2019), while the generalized loading criterion (Hashiguchi, 2018a) described in Sect. 11.7 was incorporated in Anjiki et al. (2019) (see Hashiguchi (2020) in detail). Further, it was given by Iguchi et al. (2019) and Ttutsumi et al. (2019). However, the tangential-inelastic strain rate is ignored and the inefficient calculation by the double loop is used in the former. The isotropic hardening stagnation is ignored and the inefficient return-mapping by the cutting-plane projection is adopted in the latter. The formulation of the complete implicit calculation method by the closed point projection for the extended subloading surface model incorporating all the functions including the isotropic hardening stagnation and the tangential-inelastic strain rate is required. The simultaneous equation for the evolution rules of the variables to this aim is given in the following (Hashiguchi, 2021a).
328
12
Constitutive Equations of Metals
ð12:33Þ
where eetnal n þ 1 is the elastic strain in the elastic trial step and Fn þ 1 ¼ F0 ½1 þ Sr f1 expðcH Hn þ 1 Þg n þ 1 ¼ n
0n þ 1 r jj r0n þ 1 jj
_
0n þ 1 ¼ r0n þ 1 þ Rn þ 1^c0n þ 1 ; r
n þ 1 D epn þ 1 ¼ epn þ n kn þ 1
ð12:34Þ ð12:35Þ ð12:36Þ
ð12:37Þ
12.4
Implicit Stress-Integration
329
Tn þ 1 ¼ ~cf½1 expðRn þ 1 Þge
ð12:38Þ
~ep ~n þ 1 ¼ np þ 1 ; ~epn þ 1 ¼ epn þ 1 qn þ 1 n ~e
ð12:39Þ
nþ1
ð12:40Þ ð12:41Þ ^c n þ 1 : n n þ 1 Cn n þ 1 n
ð12:42Þ
u n +1 = u exp(ucℜ cn +1 Cn n +1)
ð12:43Þ
~ n þ 1, R ~ n þ 1 and The eight unknown variables rn þ 1 , H n þ 1 , an þ 1 , cn þ 1 , qn þ 1 , K Dkn þ 1 are involved in the simultaneous Eq. (12.33) composed of the eight equations, where en þ 1 and en are the known input variables.
12.5
Material Parameters and Comparisons with Test Data
Material parameters will be collectively shown and the capability of the subloading surface model to describe various loading behavior will be verified by the comparisons with test data in this section.
12.5.1
Material Parameters
Material parameters are shown collectively below for three versions of the subloading surface model. (i) The initial subloading surface model, which is the improvement of the conventional elastoplasticity model only with the isotropic and the kinematic hardenings, contains the following 7 material constants and 2 initial values. Material constants: Elastic moduli: E;m ( isotropic : sr ; cH Hardening kinematic : ck ; bk Evolution of normal - yieldratio : u
330
12
Constitutive Equations of Metals
Initial values:
Normal-yield surface
size : F0 center: a0
The computer program is shown in Appendix L(a) (i). (ii) The extended subloading surface model, in which the tangential-inelastic strain rate and the isotropic hardening stagnation are ignored, contains the following 12 material constants and 3 initial values. Material constants: Elastic moduli: E;m ( isotropic : isotropic : sr ; cH Hardening kinematic : ck ; bk Evolution of normal - yield ratio : u; uc ; Re ð\1Þ Translationon of elastic-core : ce ; v ( \1) Initial values: ( Normal - yield surface
size : F 0 center : a0
Elastic-core: c0 The computer program is shown in Appendix L(a) (ii). (iii) The general subloading surface model, which possesses all the behaviors involving the tangential-inelastic strain rate and the isotropic stagnation, contains the following 16 material constants and 5 initial values at most, while the full version of computer program is shown in Appendix L(a) (iii). Material constants: Elastic moduli: E;m ( isotropic : sr ; cH Hardening kinematic : ck ; bk Evolution of normal - yield ratio : u; uc ; Re ð\1Þ Translationon of elastic - core : ce ; v ( \1) Tangential inelasticity : ~c; ~ nð 1Þ Stagnation of isotropic hardening : C (0 C 1); 1 ([ 1); t
12.5
Material Parameters and Comparisons with Test Data
Initial values: Normal-yield surface
(
331
size : F 0 center : a0
Elastic-core: c0 Normal - isotropic hardening surface
(
~0 size : K center : q0
The determination of these material parameters is explained below in brief. (1) Young’s modulus E and Poisson’s ratio m are determined from the slope and the ratio of lateral to axial strains in the initial part of stress–strain curve. (2) sr , cH and F0 for the isotropic hardening and ck , bk and a0 for the kinematic hardening are determined from stress–strain curves in the initial and the inverse loadings. (3) u, uc and Re ð\1Þ for the evolution of the normal-yield ratio are determined from the stress–strain curve in the subyield state, i.e. the elastic–plastic transitional state. (4) ce , v and c0 for the elastic-core are determined from the stress–strain curves in cyclic loading. (5) ~c and ~ n for the tangential-inelastic strain rate are determined by the difference of the strain in the non-proportional loading from that in the proportional loading. ~ 0 and q0 for the isotropic-hardening stagnation are determined from (6) C, 1, t, K the stress–strain curves in cyclic loading under a constant strain amplitude. All of these material parameters except for ~c and n~ for the tangential-inelastic strain rate can be determined only by the stress–strain curves in the uniaxial loading for initial isotropic materials. One can put a0 ¼ c0 ¼ q0 ¼ O for the initial isotropy, ~ 0 ¼ ~ep0 ~ 0 by K which is assumed in all the subsequent simulations. We calculate K ~ 0 ¼ 1 by inputting a small value of ep0 and q0 ¼ O in order that the leading to R isotropic hardening rule in Eq. (12.7) without the isotropic stagnation holds in the initial loading process for all the present simulations. Needless to say, the tangential inelasticity is irrelevant to the proportional loading. The efficient and accurate determination of the material parameters in the subloading surface model for metals can be referred to Liu et al. (2022) by adopting the artificial neural network.
12.5.2
Comparisons with Test Data
The capability of the present model for describing the deformation behavior of metals is verified through comparisons with several basic test data in this section, referring to Hashiguchi and Yamakawa (2012) and Hashiguchi and Ueno (2017). Capability of
332
12
Constitutive Equations of Metals
unconventional plasticity model aimed at describing plastic strain rate induced by a rate of stress inside yield surface must be evaluated by a degree in which cyclic loading behavior can be described appropriately. Then, various cyclic loading test data in uniaxial loading are first simulated and thereafter a circular strain path test datum is simulated to verify capability for describing non-proportional loading behavior. The cyclic loading behavior under the stress amplitude to both positive and negative sides can be predicted to some extent by any models, including even the conventional plasticity model. On the other hand, the prediction of the cyclic loading behavior under the stress amplitude in positive or negative one side, i.e. the pulsating loading inducing the so-called mechanical ratcheting effect requires a high ability for the description of plastic strain rate induced by the rate of stress inside the yield surface. Furthermore, it is noteworthy that we often encounter the pulsating loading phenomena in the boundary-value problems in engineering practice, e.g. railways and gears. The comparison with the test data for the 1070 steel under the cyclic loading of axial stress between 0 and þ 830 MPa after Jiang and Zhang (2008) is depicted in Fig. 12.3, where the material parameters are selected as follows: Material constants: Elastic moduli : E ¼ 160; 000 MPa; m ¼ 0:3; ( isotropic hardening : sr ¼ 0:58; cH ¼ 170; Hardening kinematic hardening : ck ¼ 5; 000 MPa; bk ¼ 0:5; Evolution of normal - yield ratio: u ¼ 200; uc ¼ 6; Re ¼ 0:5 Translationon of elastic - core: ce ¼ 1; 000; v ¼ 0:7; Stagnation of isotropic hardening: C ¼ 0:5; 1 ¼ 5; t ¼ 0:1; Initial values: Isotropic hardening function : F0 ¼ 507 MPa: The relation of the axial stress and the axial components of back-stress and similarity-center versus the axial strain and the relation of the axial strain versus the number of cycles are depicted in Fig. 12.3a, where the axial components are designated by ðÞa . The accumulation of axial strain is simulated closely by the present model. The calculation is controlled automatically such that the stress and the back-stress are attracted to the normal-yield and the normal-isotropic hardening surfaces, respectively, as known from the variations of the normal-yield ratio R and ~ depicted in Fig. 12.3b. Accumulation of the normal-isotropic hardening ratio R axial strain is overestimated as depicted in Fig. 12.3c if the reloading behavior is not improved by setting uc ¼ 0 ignoring the Masing effect. Despite of the improvement for reloading behavior, however, hysteresis loops are simulated as
12.5
Material Parameters and Comparisons with Test Data
333
Test result (Jiang and Zhang, 2008)
Model simulation
900
3.0
800
2.5
700
a
(MPa)
a
600
(%)
2.0
sa
500
1.5
400
1.0
300 200
a
0.5
100
0.0
0 0.0
0.5
1.0
1.5
2.0
2.5
3.0 a
3.5
0
(%)
100
200
(a) 1.0
R
1.0
R
0.8
0.8
0.6
0.6
0.4
0.4
0.2
0.2
0.0 0.0
300 400 500 600 Number of cycles
0.0
0.5
1.0
1.5
2.0
2.5
3.0 a
0.0
3.5 (%)
0.5
1.0
1.5
2.0
2.5
3.0 a
(b)
3.5 (%)
3.5 900
3.0
800
2.5
700 a
600
sa
(MPa) 500
a
(%)
2.0
1.5
400
1.0
300
a
200
0.5
100
0 0.0
0.0 0.5
1.0
1.5
2.0
2.5
3.0 a
(%)
3.5
0
(c)
100
200
300 400 500 Number of cycles
600
Fig. 12.3 Uniaxial cyclic loading behavior under the pulsating loading between 0 and 830 MPa of 1070 steel (Test data after Jiang and Zhang (2008)): a Test result and simulation result and simulation without stagnation of isotropic hardening, b Variations of normal-yield ratio and normal-isotropic hardening ratio and c Test result and simulation without improvement of reloading behavior
narrower than those in the test result in order to fit the strain accumulation in the test result. A further improvement is desirable for this insufficiency. Next, examine the uniaxial cyclic loading behavior under the constant stress amplitude to both positive and negative sides with different magnitudes. Comparison with the test data for the 304L steel under the cyclic loading of axial stress between þ 250 and 150 MPa after Hassan et al. (2008) is depicted in Fig. 12.4 where the material parameters are selected as shown below.
334
12
Constitutive Equations of Metals
Test result (Hassan et al., 2008) a
Model simulation
(MPa)
300
10 20
2.5
200
2.0 1.5
sa
100
1.0
a
0
0.5
1.0
1.5
2.0
2.5 a
0.5
(%)
0.0
-100
0.0
40
20
60
100
120
140
Number of cycles
(a)
-200
R
(%)
a
40 6080100cycles
1.0
1.0
R
0.8
0.8
0.6
0.6
0.4
0.4
0.2
0.2
0.0
0.0
0.5
1.0
1.5
2.0 a
0.5
2.5
(%)
1.0
1.5
2.0
2.5 a
(b)
10 20 40 60 80 100cycles
a
(%)
(%)
㻞㻚㻡㻌 㻞㻚㻜㻌 㻝㻚㻡㻌
a (MPa)
㻝㻚㻜㻌 a
㻜㻚㻡㻌
(%)
㻜㻚㻜㻌 㻜
㻞㻜
㻠㻜
(c)
㻢㻜
㻤㻜
㻝㻜㻜
Number of cycles
Fig. 12.4 Uniaxial cyclic loading behavior under the constant stress amplitude between –150 and 250 MPa of 304L steel (Test data after Hassan et al. (2008)): a Test result and simulation, b Variations of normal-yield ratio and normal-isotropic hardening ratio and c Simulation by modified Chaboche model (cf. Hassan et al. (2008))
Material constants: Elastic moduli : E ¼ 200; 000 MPa; m ¼ 0:3; ( isotropic : sr ¼ 0:3; cH ¼ 30; Hardening kinematic : ck ¼ 130 MPa; bk ¼ 0:9; Evolution of normal - yield ratio : u ¼ 2; uc ¼ 10; Re ¼ 0:5 Translationon of elastic - core : ce ¼ 143; v ¼ 0:7; Stagnation of isotropic hardening : C ¼ 0:5; 1 ¼ 15; t ¼ 1;
12.5
Material Parameters and Comparisons with Test Data
335
Initial values: Isotropic hardening function: F0 ¼ 232 MPa: The relation of the axial stress and the axial components of back-stress and similarity-center vs. the axial strain and the relation of the axial strain vs. the number of cycles are depicted in Fig. 12.4a. Both the accumulation of strain and the hysteresis loops are simulated closely by the present model. The calculation is controlled automatically such that the stress and the back-stress are attracted to the normal-yield and the normal-isotropic hardening surfaces, respectively, as known from the variations of the normal-yield ratio R and the normal-isotropic hardening ~ depicted in Fig. 12.4b. The relations of the axial stress versus the axial strain ratio R and the relation of the axial strain vs. the number of cycles simulated using the modified Chaboche model (Chaboche 1991) are also depicted in Fig. 12.4c in which the strain is simulated as larger than the test result and the hysteresis loops are simulated as narrower than the test data. The prediction of this steel deformation behavior will be improved by incorporating the rate-dependence. Further, we examine the uniaxial cyclic loading behavior for constant symmetric stress amplitude to both positive and negative sides. Comparison with test data of the 304 steel under the cyclic loading of axial stress between +182 and −182 MPa under the constant hoop stress 80 MPa after Xia and Ellyin (1994) is depicted in Fig. 12.5 where the material parameters are selected as follows: Material constants: Elastic moduli : E ¼ 190; 000 MPa; m ¼ 0:3; ( isotropic : sr ¼ 1:3; cH ¼ 100; Hardening kinematic : ck ¼ 25 MPa; bk ¼ 0:3; Evolution of normal - yield ratio : u ¼ 200; uc ¼ 6; Re ¼ 0:5; Translationon of elastic - core : ce ¼ 1; 429; v ¼ 0:7; Stagnation of isotropic hardening : C ¼ 0:5; 1 ¼ 8; t ¼ 5; Initial values: Isotropic hardening function: F0 ¼ 212 MPa: The relation of the axial stress and the axial components of back-stress and similarity-center versus the axial strain and the circumferential strain el with the number of cycles are shown in Fig. 12.5a, while the back-stress is induced quite slightly so that it is invisible in this figure. The simulations for the accumulation of
336
12
Constitutive Equations of Metals Model simulation
Test result ( Xia and Ellyin, 1994 ) N=20, 15 1 to 10
a
(MPa) a
(MPa)
N=1 to 10
N=15, 20
l
a
(%)
(a)
R
a
N=20, 15
N=1 to 10
(%)
(MPa)
a
(MPa)
200 150 100 50 0 50 100 150 200 0.0
200 a
0.0
(c)
a
a
200
100 0.2
(b)
0
0.4 (%)
R
100
0.6
(%)
N=1 to 10
0.2
(%)
N=15, 20
0.4
0.6 l
(%)
Fig. 12.5 Uniaxial cyclic loading behavior under the constant stress amplitude between 182 and +182 MPa of 304L steel (Test data after Xia and Ellyin (1994)): a Test result and simulation, b Variations of normal-yield ratio and normal-isotropic hardening ratio and c Simulation by Xia and Ellyin (1994)
axial strain and the hysteresis loops agree well with the test result, except for the prediction of hysteresis loops as narrower than the test result in the initial stage. Here, the axial strain and the lateral strain are accumulated to the compression side and the extension side, respectively, by the application of the hoop stress 80 MPa. The calculation is automatically controlled such that the stress and the back-stress are attracted to the normal-yield and the normal-isotropic hardening surfaces, respectively, as known from the variations of the normal-yield ratio R and the ~ depicted in Fig. 12.5b. The relations of the axial normal-isotropic hardening ratio R stress vs. the axial and lateral strains and the relation with the number of cycles simulated by Xia and Ellyin (1994; cf. also Ellyin 1997) are also depicted in Fig. 12.5c where both the axial and the circumferential strains are overestimated.
12.5
Material Parameters and Comparisons with Test Data
337 Model simulation
Test result (Xia and Ellyin, 1994) a
(MPa)
(MPa)
a
400
400
sa
300
sa
300
200
200 a
a
-3.0 -2.0
-1.0
0 0.0
1.0
2.0
3.0 a
-200
(%)
-3.0 -2.0
0 -1.0 0.0 -200
-400
(a)
1.0
1.0
0.6
0.6
0.4
0.4
0.2
0.2
0.0 -1 .5 -1 .0 -0 .5 0
a
a
(%)
(b) R
0.8
0.8
-2
3.0
-300
-400 R
-2 .5
2.0
-300
-3 .5 -3
1.0
0 .5
1
1 .5
2
2 .5
3
3 .5
-3 .5 -3 -2 .5
-2
0.0 -1 .5 -1 .0 -0 .5 0
(%)
a
0 .5
1
1 .5
2
2 .5
3
3 .5
(%)
(c) a
(MPa)
a
a
(d)
(MPa)
(%)
a
(%)
(e)
Fig. 12.6 Uniaxial cyclic loading behavior under the constant strain amplitude the 5 levels increasing strain amplitudes of 316 steel (Test data after Chaboche et al. (1979)): a Test result and simulation by present model, b Test result and simulation without stagnation of isotropic hardening, c Variations of normal-yield ratio and normal-isotropic hardening, d Simulation by Chaboche (2008) and e Simulation by Ellyin and Xia (1989)
Furthermore, examine the uniaxial cyclic loading behavior under the constant symmetric strain amplitudes to both positive and negative sides. Comparison with the test data of the 316 steel under the cyclic loading with the increasing axial strain amplitudes 1:0; 1.5, 2.0, 2.5, 3.0% after Chaboche et al. (1979) is depicted in Fig. 12.6 where the material parameters are selected as follows:
338
12
Constitutive Equations of Metals
Material constants: Elastic moduli : E ¼ 170; 000 MPa; m ¼ 0:3; ( isotropic : sr ¼ 0:85; cH ¼ 5; Hardening kinematic : ck ¼ 2; 000 MPa; bk ¼ 0:5; Evolution of normal - yield ratio: u ¼ 100; uc ¼ 3; Re ¼ 0:5; Translationon of elastic - core : ce ¼ 283; v ¼ 0:7; Stagnation of isotropic hardening : C ¼ 0:5; 1 ¼ 5; t ¼ 1; Initial values: Isotropic hardening function : F0 ¼ 320 MPa: The relation of the axial stress and the axial components of back-stress and similarity-center versus the axial strain are shown in Fig. 12.6a. The hysteresis loops and the stagnation of isotropic hardening are simulated closely by the present model. On the other hand, the calculated result without the cyclic stagnation of isotropic hardening overestimates the hardening behavior as shown in Fig. 12.6b. The calculation is controlled automatically such that the stress and the back-stress are attracted to the normal-yield and the normal-isotropic hardening surfaces, respectively, as known from the variations of the normal-yield ratio R and the ~ depicted in Fig. 12.6c. The relations of the axial normal-isotropic hardening ratio R stress and the axial strain simulated by Chaboche (1991) and Ellyin and Xia (1989) are depicted in Fig. 12.6d, e, respectively. The strain in the initial stage is simulated as larger than the test result by the former and the curves predicted by the latter is not smooth but piece-wise linear. Finally, we examine the non-proportional loading behavior by the comparison with the test data of the austenitic 17–12 Mo SPH carbon stainless steel after Delobelle et al. (1995). The specimen is subjected to the approximately circular strain path in the strain plane ðez ; ezh Þ by the inputs of the axial strain ez ¼ 0:004 sinðh p=2Þ and the axial-circumferential shear strain ezh ¼ 0:0036 sin h in the sinusoidal waves under the constant circumferential normal stress rh ¼ 50 MPa during 40 cycles after the uniaxial straining to ez ¼ 0:004 as depicted in Fig. 12.7. The simulation are performed by using the material parameters selected as follows:
12.5
Material Parameters and Comparisons with Test Data z
339
z
0.004
0
r 0
0.0036
z
Fig. 12.7 Circular strain path loading given by the axial strain and the axial-circumferential engineering shear strain
Material constants: Elastic moduli : E ¼ 170; 000 MPa; m ¼ 0:3ðG ¼ 65; 385 MPaÞ; ( isotropic : sr ¼ 1:7; cH ¼ 40; Hardening kinematic : ck ¼ 200 MPa; bk ¼ 0:9; Evolution of normal-yield ratio : u ¼ 800; uc ¼ 3; Re ¼ 0:5; Translationon of elastic - core : ce ¼ 283; v ¼ 0:7; ~ ¼ 3; Tan gential inelasticity : ~c ¼ 0:6; n Stagnation of isotropic hardening : C ¼ 0:5; 1 ¼ 5; t ¼ 1; Initial values: Isotropic hardening function: F0 ¼ 240 MPa: The strain path ðez ; eh Þ (eh : circumferential normal strain) and the stress path pffiffiffi rz ; 3rzh (rah : axial-circumferential shear stress) are shown for the test result and the model simulation in Fig. 12.8a(i) and (ii), respectively. The simulation of the stress path and the accumulation of lateral strain are in good agreement with the test result. The stress and the plastic strain are attracted to the normal-yield and the normal-isotropic hardening surfaces, respectively, as known from the variations of ~ depicted in the normal-yield ratio R and the normal-isotropic hardening ratio R Fig. 12.8a(ii). The circular strain path in this test produces the spiral stress path approaching the circular stress path along the Mises yield surface which is qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffi 2ffi expressed by the circle r2z þ 3rzh ¼ F in the two-dimensional stress plane pffiffiffi rz ; 3rzh . The tangential strain is added in the axial strain ez and the axial-circumferential strain ezh but it is not added to eh . Therefore, the variation of eh for the variation of ez is predicted to be larger if the tangential-inelastic strain added to ez is ignored as shown in Fig. 12.8b.
340
12
Constitutive Equations of Metals
(%)
(%) 1.8
1.8 1.6 1.4 1.2 1.0 0.8 0.4 0.2 0.0 0.5
1.6 1.4 1.2 1.0 0.8 0.6 0.4 0.2
0.4
0.3
0.2
0.1 0
0.1 z (%)
0.2
0.3
0.4
0.5
0.0 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 -0.2
0 .1
0 .2
0 .3
0 .4
0 .5
(%)
z
3 z (MPa) 600 400 200
0 -600
-400
0
-200
200
600
400
z (MPa)
-200
-400 -600
(i) R 1.0 0.8 0.6 0.4 0.2 0.0 -0.5 -0.4 -0.3 -0.2 -0.1 0.0
0 .1
0 .2
0 .3
0 .4 z (%)
0 .1
0 .2
0 .3
~
0.5
R 1.0 0.8 0.6 0.4 0.2 0.0 -0.5 -0.4 -0.3 -0.2 -0.1 0.0
(a)
(ii)
z
0 .4
0.5
(%)
Fig. 12.8 a Circular strain path loading given by the axial strain and the axial-circumferential engineering shear strain during 40 cycles after the uniaxial loading of austenitic 17–12 Mo SPH carbon stainless steel (Test data after Delobelle et al. (1995)): (i) Test result, (ii) Model simulation. b (i) Test result, (ii) Model simulation without tangential-inelasticity
12.5
Material Parameters and Comparisons with Test Data
341 (%) 2.4
(%)
2.2 2.0 1.8
1.8 1.6 1.4 1.2 1.0 0.8 0.4 0.2 0.0 0.5
1.6 1.4 1.2 1.0 0.8 0.6 0.4 0.2
0.4
0.3
0.2
0.1 0
0.1
0.2
0.3
0.4
0.5
-0.5
-0.4
-0.3
-0.2
0.0 -0.1 -0.20.0
z (%)
0 .1
0 .2
0 .3
0 .4
200
400
600
0 .1
0 .2
0 .3
0 .4
z
(%)
z
(%)
0 .5
(%)
z
3 z (MPa) 600 400 200
-600
-400
0 -200
0
z (MPa)
-200
-400 -600
(i)
R
-0 .5 -0 .4 -0 .3 -0 .2 -0 .1 0 .0
~
R
0.5
-0 .5 -0 .4 -0 .3 -0 .2 -0 .1 0 .0
(b) Fig. 12.8 (continued)
(ii)
0 .1
0 .2
0 .3
0 .4
0.5
342
12
12.6
Constitutive Equations of Metals
Analyses of Engineering Phenomena
The metal forming analyses are of importance in the industrial production. The analyses of the two typical phenomena, i.e. the springbuck analysis and the residual stress analysis by use of the subloading surface model will be described in this section. (1) Springback analysis The high tensile (strength) steel sheets and aluminum sheets exhibiting far larger springback than ordinary mild steel sheets are widely used in automobile industries. First, it should be noticed that not only elastic but also slight plastic deformations are induced in the spring-back process induced during the drawing out of the punch from the dies. Therefore, an appropriate unconventional plastic constitutive equation must be adopted for the prediction of this phenomenon, which is capable of describing the plastic deformation in the unloading process accurately. Needless to say, it cannot be described by the constitutive models which adopt the yield surface enclosing a purely-elastic domain, i.e. the conventional plasticity model represented by the cylindrical yield surface (Chaboche-Ohno) model, and the two-surface (Dafalias-Yoshida) models. In particular, it can never be described by the two-surface model at all contrary to the repeated propagations by Yoshida et al. (Yoshida and Uemori 2002a, b, 2003; Yoshida and Amaishi 2020) and the standard installation into the commercial software LS-DYNA, PAM-STAMP and JSOL-JSTAMP, since a plastic strain rate in the unloading process cannot be described as was described in Sect. 10.3.3. The similar irrational analyses by the two-surface model are performed by Wagoner et al. (2013), etc. In their analysis, the plastic deformation during the unloading process is ignored. Instead, the Young’s modulus E is replaced by the chord-Young’s modulus E which is calculated by connecting the beginning point of the unloading and the unloaded point to the stress-free state and the chord-Young’s modulus is assumed to decrease only by the equivalent plastic strain induced during the preceding monotonic loading process as shown in Fig. 12.9.
where E, E and Emin are the cord modulus, its initial value (true Young’s modulus) and the final converged value, respectively, and n is the material constant. However, their formulation is irrational against the facts observed in real material behavior described as follow: (1) The innegligible plastic deformation is induced during the unloading process. (2) The real Young’s modulus does not change in the range of the usual sheet forming process. Microscopically, only the elastic deformation induced by the elastic deformations of mate-rial particles themselves is induced in the
12.6
Analyses of Engineering Phenomena
343
Real unloading-reloading curve with plastic deformation Real (true) elastic curve induced in beginning state of reverse loading, which is constant in deformation for range of springback Chord modulus incorporated in Yoshida’s group
E 1
E 1
Emin
Cyclic loading under constant stress amplitude Accumulation of plastic strain cannot be described at all by Yoshida-Uemori model leading to dangerous design !!!
1
0
Fig. 12.9 Chord Young’s modulus used in ad hoc. Yoshida and Uemori (2003) model which is incapable of describing strain accumulation under constant stress amplitude, where E is the true Young’s modulus which is constant, E is the chord Young’s modulus and E min is its minimum value
initiation of the reverse loading process. In fact, the inclination of the stress vs. strain curve at the moment of the reverse loading process does not change as observed in the test data for the uniaxial loading. Besides, in fact, during a far larger deformation process, the decrease of the Young’s modulus is induced by the damage of materials by the generation and growth of minute cracks and thus the Young’s modulus decreases acceleratingly once it decreases in the monotonic loading process. The Yoshida’s group ignores this primitive knowledge in the continuum damage mechanics. Therefore, the ad hoc. method adopted in their analysis must not be used in the springback analysis and, needless to say, in the general loading process, since the plastic deformation during the pulsating loading process is ignored resulting in the dangerous engineering design. The pertinent calculation result of the springback is shown below, which was analyzed by Dr. Motoharu Tateishi (MSC Software, Ltd., Japan) by implementing the subloading surface model to the commercial software Marc (MSC Software, Ltd.). The calculation results of the shapes of the sheet after the springback are shown in Fig. 12.10, choosing the die diameter 5 mm and using the following values of material parameters. Conventional model (u
)
Subloading surface model
Fig. 12.10 Springback analysis
344
12
Constitutive Equations of Metals
Material constants: Elastic moduli : E ¼ 205; 000 MPa; m ¼ 0:3; ( isotropic : h1 ¼ 0:5; h2 ¼ 15; Hardening kinematic : ck ¼ 3000 MPa; bk ¼ 0:5; Evolution of normal - yield ratio : u ¼ 200; uc ¼ 3:5; Re ¼ 0:5; n ¼ 1; Translationon of elastic - core : ce ¼ 400; v ¼ 0:7; Stagnation of isotropic hardening : C ¼ 0:5; 1 ¼ 20; t ¼ 1; Initial values: Isotropic hardening function : F0 ¼ 400 MPa: The simulation by the conventional elastoplastic model predicting the springback only by the elastic deformation is also shown in Fig. 12.10, which is calculated by setting u ! 1 in the subloading surface model and almost identical to the simulations by the Chaboche model and the two surface model without resorting to the irrational method as done in the Yoshida-Uemori model (Yoshida and Uemori 2002a, b, 2003; Yoshida and Amaishi 2020), since the plastic deformation cannot be described in the unloading process as was explained in Sect. 10.3. The enough springback is predicted, which is caused by the plastic deformation in the stress-releasing process by virtue of the advantage of the subloading surface model describing the plastic strain rate due to the rate of stress inside the yield surface. In contrast, the springback is predicted slightly by the conventional elastoplastic model which is realized by using the large value for the material constant in the evolution of the normal-yield ratio u ¼ 100; 000. Then, the importance is recognized for the introduction of the rigorous elastoplastic model, i.e. the subloading surface model capable of describing the plastic strain rate in the stress-reducing process appropriately. Hereinafter, it is desirable that the prediction of springback behavior will be executed by the pertinent analysis exploiting the subloading surface model, aiming at the epochal improvement of the prediction of the springback behavior in industries. (2) Residual stress analysis (Higuchi and Okamura 2016) The prediction of the residual stress is of importance in the metal forming process. The estimations of residual stress change due to cyclic loading by the conventional elastoplasticity model and the subloading surface model are shown below, which was examined by Higuchi and Okamura (2016). A four-point cyclic bending test was conducted to examine the change of the residual stress in the cyclic loading process. A specimen with a width of 13 mm, a thickness of 13 mm and a length of 100 mm was cut from a seamless steel pipe P110. The specimen was loaded under a four-point bend configuration with the intervals of 20 mm between the inner rollers and 80 mm between the outer rollers as shown in Fig. 12.11. At first, the specimen was
12.6
Analyses of Engineering Phenomena
345
Fig. 12.11 Configuration of the four-point bending test
㻱㼤㼜㼑㼞㼕㼙㼑㼚㼠 㻿㼡㼎㼘㼛㼍㼐㼕㼚㼓㻌㼟㼡㼞㼒㼍㼏㼑㻌㼙㼛㼐㼑㼘 㻯㼔㼍㼎㼛㼏㼔㼑㻌㼙㼛㼐㼑㼘
Fig. 12.12 Comparison of stress–strain curves between experiment and simulations in uniaxial loading
plastically deformed by static bending load corresponding to the maximum bending stress of 900 MPa. After the unloading of the static bending load, the compressive residual stress was generated in the side of outer rollers and the tensile residual stress in the side of inner rollers. The distribution of the residual stress was measured by the X-ray stress measurement method. Then, the specimen was turned upside down, so that the side of outer rollers was in the tensile residual stress state. Sinusoidal waveform load between 5 and 100% of 900 MPa was applied 20 times to the specimen. The maximum bending stress was 500 MPa. The distributions of the residual stress after cyclic loading were measured by the X-ray stress measurement method. The simulations by the Chaboche model described in Subsect. 10.3.2 and the subloading surface model described in Chap. 12 are executed, while the commercial software Abaqus is used in the simulation by the Chaboche model. First, material constants of these two models were determined so as to fit to the
346
12
Distance from upper surface
(a) Chaboche model
Constitutive Equations of Metals
Distance from upper surface
(b) Subloading surface model
Fig. 12.13 Comparison of distributions of residual stresses before and after cyclic loading between experiment and simulations
test data of uniaxial cyclic loading of a round bar specimen of P110 as shown in Fig. 12.12. The test data can be simulated accurately by the subloading surface model. On the other hand, the simulation by the Chaboche model is not in agreement with the test data. Especially, it was difficult to simulate appropriately the smooth elastic–plastic transition in the reverse loading processes by the Chaboche model. Measured distributions of the residual stresses after the initial loading and the 20 times cyclic loading are simulated by the two models as shown in Fig. 12.13. The horizontal axis denotes the distance from upper to lower surface at center of the specimen. The initial residual stress distributions are simulated well by both of these models. As for the residual stress distribution after cyclic loading, however, the decreases of the residual stresses near the upper and the lower surfaces of the specimen is accurately simulated by the subloading surface model, whereas they are not simulated by the Chaboche model. This would be caused by the fact that the subloading surface model is capable of describing the cyclic loading behavior more accurately than the other constitutive models.
12.7
Orthotropic Anisotropy
The kinematic hardening incorporated in the foregoing is regarded to be the induced anisotropy. On the other hand, various inherent anisotropies are induced in the manufacturing process of metals. The typical inherent anisotropy is the orthotropic anisotropy formulated by Hill (1948).
12.7
Orthotropic Anisotropy
347
Now, consider the general yield function in the quadratic form shown as follows: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi f rij ¼ 12 Cijkl rij rkl
ð12:44Þ
where Cijkl is the fourth-order anisotropic tensor having eighty-one components fulfilling the symmetries Cijkl ¼ Cijlk ¼ Cjikl ¼ Cjilk ¼ Cklji ¼ Cklji ¼ Clkij ¼ Clkij
ð12:45Þ
by the minor symmetries Cijkl ¼ Cijlk ¼ Cjikl based on the symmetry of the stress tensor rij ¼ rji and the major symmetries Cijkl ¼ Cklij based on Cijkl rij rkl ¼ Cklij rkl rij ¼ Cklij rij rkl . Then, the independent components are reduced to twenty-one leading to Cijkl rij rkl ¼ C1111 r211 þ 2C1122 r11 r22 þ 2C1133 r11 r33 þ 2C1112 r11 r12 þ 2C1123 r11 r23 þ 2C1131 r11 r31 þ C2222 r222 þ 2C2233 r22 r33 þ 2C2212 r22 r12 þ 2C2223 r22 r23 þ 2C2231 r22 r31 þ C3333 r233 þ 2C3312 r33 r12 þ 2C3323 r33 r23 þ 2C3331 r33 r31 þ C1212 r212 þ 2C1223 r12 r23 þ 2C1231 r12 r31 þ C2323 r223 þ 2C2331 r23 r31 þ C3131 r231
ð12:46Þ which is the general form of yield function in the quadratic form. Here, assuming the plastic incompressibility, it holds that 2 @ 2f =@rpq dpq ¼ @Cijkl rij rkl =@rpq dpq ¼ Cijkl dpi dqj rkl dpq þ Cijkl rij dpk dql dpq ¼ Cppkl rkl þ Cijpp rij ¼ Cppkl rkl þ Cppij rij ¼ 2Cppkl rkl ¼ 0 This relation must hold for any rij and thus one obtains
which leads to
Cppkl ¼ Cijqq ¼ 0
ð12:47Þ
9 C1111 þ C2211 þ C3311 ¼ 0 = C1122 þ C2222 þ C3322 ¼ 0 ; C1133 þ C2233 þ C3333 ¼ 0
ð12:48Þ
9 9 C1112 þ C2212 þ C3312 ¼ 0 = C3312 ¼ ðC1112 þ C2212 Þ = C1123 þ C2223 þ C3323 ¼ 0 ! C1123 ¼ ðC2223 þ C3323 Þ ; ; C1131 þ C2231 þ C3331 ¼ 0 C2231 ¼ ðC1131 þ C3331 Þ
ð12:49Þ
348
12
Constitutive Equations of Metals
The substitution of Eq. (12.49) into Eq. (12.46) gives the expression Cijkl rij rkl ¼ C1111 r211 þ 2C1122 r11 r22 þ 2C1133 r11 r33 þ 2C1112 r11 r12 2ðC2223 þ C3323 Þr11 r23 þ 2C1131 r11 r31 þ C2222 r222 þ 2C2233 r22 r33 þ 2C2212 r22 r12 þ 2C2223 r22 r23 2ðC1131 þ C3331 Þr22 r31 þ C3333 r233 2ðC1112 þ C2212 Þr33 r12 þ 2C3323 r33 r23 þ 2C3331 r33 r31 þ C1212 r212
þ 2C1223 r12 r23 þ 2C1231 r12 r31 þ C2323 r223 þ 2C2331 r23 r31 þ C3131 r231
ð12:50Þ Further, noting Eq. (12.48), the terms in the form Ciijj rii rjj (no sum) are written as C1111 r211 þ C2222 r222 þ C3333 r233 þ 2C1122 r11 r22 þ 2C2233 r22 r33 þ 2C1133 r11 r33 ¼ C1111 r211 þ C2222 r222 þ C3333 r233 C1122 ðr11 r22 Þ2 þ C1122 r211 þ C1122 r222 C2233 ðr22 r33 Þ2 þ C2233 r222 þ C2233 r233 C1133 ðr33 r11 Þ2 þ C1133 r233 þ C2233 r211 ¼ ðC1111 þ C1122 þ C2233 Þr211 þ ðC1122 þ C2222 þ C2233 Þr222 þ ðC1133 þ C2233 þ C3333 Þr233 C1122 ðr11 r22 Þ2 C2233 ðr22 r33 Þ2 C1133 ðr33 r11 Þ2 ¼ C1122 ðr11 r22 Þ2 C2233 ðr22 r33 Þ2 C1133 ðr33 r11 Þ2
ð12:51Þ Then, by setting a1 C1122 ; a2 C2233 a3 C1133 a4 2C1112 ; a5 2C2212 ; a6 C2223 a7 2C3323 a8 2C3331 ; a9 2C1131 a10 2C1223 ; a11 2C2331 a12 2C1231 a13 C1212 ; a14 C2323 a15 C3131
9 > > > > = > > > > ;
ð12:52Þ
and substituting Eqs. (12.50) and (12.51) with Eq. (12.52) into Eq. (12.50) reads: Cijkl rij rkl ¼ a1 ðr11 r22 Þ2 þ a2 ðr22 r33 Þ2 þ a3 ðr33 r11 Þ2 þ ½a4 ðr33 r11 Þ þ a5 ðr33 r22 Þr12 þ ½a6 ðr11 r22 Þ þ a7 ðr22 r33 Þr23 þ ½a8 ðr22 r33 Þ þ a9 ðr22 r11 Þr31 þ a10 r12 r23 þ a11 r23 r31 þ a12 r31 r12
12.7
Orthotropic Anisotropy
349
þ a13 r212 þ a14 r223 þ a15 r231
ð12:53Þ
Equation (12.53) is the general yield function for the plastically-incompressible materials in the quadratic form. Now, assume orthotropic anisotropy. Then, if we describe the yield surface by _ the coordinate axes selected to the principal axes fe j g of orthotropic anisotropy, the yield function is independent of the sign of shear stress components in this coordinate system. Therefore, it must hold that a4 ¼ a5 ¼ a6 ¼ a7 ¼ a8 ¼ a9 ¼ a10 ¼ a11 ¼ a12 ¼ 0 Here, replacing the symbols ai as F ¼ a1 ; G ¼ a2 ; H ¼ a3 ; L ¼ a13 =2; M ¼ a14 =2; H ¼ a15 =2 used by Hill (1948), Eq. (12.53) leads to the Hill’s yield condition with orthotropic anisotropy:
1 2
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi F ðr11 r22 Þ2 þ Gðr22 r33 Þ2 þ H ðr33 r11 Þ2 þ 6 Lr212 þ Mr223 þ Nr231 ¼ FðHÞ
ð12:54Þ Here, note that for the isotropic material all the material parameters are unity, i.e. F ¼ G ¼ H ¼ L ¼ M ¼ N ¼ 1 holds and thus Eq. (12.54) is reduced to pffiffiffiffiffiffiffiffi f rij ¼ 3=2kr0 k which is the equivalent stress. While Eq. (12.54) is the ^
expression on the principal axis fe i g of orthotropic anisotropy, it is rewritten by the following equation stipulating this fact. 1 2
rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 2 2 ffi ^ ^ ^ ^ ^ ^ ^ ^ 2 2 F r11 r22 þ G r22 r33 þ H r33 r11 þ r Lr12 þ M r23 þ Nr31 ¼ FðHÞ
ð12:55Þ or 2 2 2 2 ^ ^ ^ ^ ^ ^ r r ^12 þ Mr ^ 223 þ Nr ^231 ¼ 1 r þG þH þ 6 Lr F 11 r22 22 r33 33 r11
ð12:56Þ
350
12
where
H
H
Constitutive Equations of Metals
G
9 > > 2> =
F
;G ;F ½2FðHÞ2 ½2FðHÞ2 ½2FðHÞ > L L L > > L ;M ;N 2 2 2; ½2FðHÞ ½2FðHÞ ½2FðHÞ ^
^
^
^
^
^
^
^
^
ð12:57Þ
^
^
^
Denoting r11 ; r22 ; r33 ; r12 ; r23 ; r31 in the yield state by ry1 ; ry2 ; ry3 ; ry12 ; ry23 ; ry31 , respectively, when only each stress applies, one has 9 ^ ^ y2 ^ y2 = ðH þ FÞry2 ¼ 1; ðG þ FÞr ¼ 1; ðG þ HÞr ¼ 1 1 2 3 ð12:58Þ ^ ^ ^ 6Ls y2 ¼ 1; 6Ms y2 ¼ 1; 6Ns y2 ¼ 1 ; 12
23
31
from which the material parameters are given by ! 1 1 1 1 F¼ þ ^y2 ^y2 ^ y2 2 r r2 r3 ! 1 ¼1 1 þ 1 1 G ^ y2 ^ ^ 2 r ry2 ry2 2 3 1 ! 1 1 1 1 ¼ þ ^y2 ^y2 H ^ y2 2 r r1 r2 3 1 1 ¼ ¼ ¼ 1 L ;M ;N ^y2 ^y2 ^ 6s 12 6s23 6s y2 23
9 > > > > > > > > > > > > > > =
ð12:59Þ
> > > > > > > > > > > > > > ;
Equation (12.55) is reduced under the uniaxial loading in the sheet metal forming as follows: 1 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi^ ^2 ¼1 ðF þ HÞr11 ¼ FðHÞ or ðF þ HÞr11 2
ð12:60Þ
Further, under the plane stress condition observed in the sheet metal forming it ^ ^ ^ holds that r23 ¼ r31 ¼ r33 ¼ 0 and thus Eqs. (12.55) and (12.56) are reduced to 1 2
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ^2 ^ ^ ^2 ^2 2Fr11 r22 þ ðF þ GÞr22 þ 6Lr12 ðF þ HÞr11 ¼ FðHÞ
ð12:61Þ
i.e. ^
^
^
^
^
2 2Fr r þ ðF þ GÞr 2 þ 6Lr 2 ¼ 1 ðF þ HÞr11 11 22 22 12
ð12:62Þ
12.7
Orthotropic Anisotropy
351
Equation (12.62) is described in the state that the principal stress directions coincide with the orthogonal anisotropy axes as follows: ^
^ ^
^
ðF þ HÞr12 2Fr1 r2 þ ðF þ GÞr22 ¼ 1
ð12:63Þ
The plastic strain rate is given for Eq. (12.63) as follows:
ð12:64Þ
noting the plastic incompressibility. The orthogonal anisotropy is induced seriously in the rolling process for the ^ ^ ^ sheet metal forming. Choosing the axes e1 ; e 2 ; e3 to the rolling, the traverse, and the thickness directions, respectively, the following Lankford R-value is adopted widely in order to evaluate the intensity of the orthotropy.
Reh ¼ eph =ep3
ð12:65Þ
where eph ð\0Þ is the lateral strain rate measured from the uniaxial tension test of the test specimen cut out at the angle h measured counterclockwise from the rolling direction. The plastic strain rate ep3 ð\0Þ in the thickness direction is calculated from the axial and the lateral strain rates by the assumption of plastic incompressibility (see Fig. 12.14). Reh ¼ 1 means the isotropy. A small Reh -value means that the produced sheet metal is easily thinned resulting in an easy failure. It follows from Eq. (12.64) that 9 ep2 F > ^ ^ e > R0 ¼ p ¼ for r2 ¼ r3 ¼ 0 > > > H = e3 ð12:66Þ p > > > e F ^ ^ > Re90 ¼ 1p ¼ for r1 ¼ r3 ¼ 0 > ; e3 G Substituting Eq. (12.66) into Eq. (12.63), one has ^
^
^ ^
ðH þ Re0 HÞr12 þ ðG þ Re0 HÞr22 2Re0 Hr1 r2 ¼ 1
352
12
Constitutive Equations of Metals
d 3p
dp
e3 e2
Rolling direction
e1
Fig. 12.14 Uniaxial tension test for Lankford R-value in metal formed by rolling process
leading to ^
r12 þ
G þ Re0 H ^ 2 2Re0 ^ ^ 1 r r1 r2 ¼ 2 e e 1 þ R0 ð1 þ R0 ÞH ð1 þ Re0 ÞH
ð12:67Þ
Substituting Eq. (12.59) into Eq. (12.66), it follows that 9 > > ¼ ð1 > > r3 r2 r1 = 1 1 1 > > > ð1 þ Re90 Þ ð1 Re90 Þ ¼ ð1 þ Re90 Þ > ^ y2 ^ y2 ^ y2 ; r3 r2 r1 1 ð1 þ Re0 Þ ^ y2
1 ð1 þ Re0 Þ ^ y2
1 Re0 Þ ^ y2
By solving this equation, one has 1 ^
r2y2 1 ^
r3y2
¼
9 Re0 ð1 þ Re90 Þ 1 > > > ^ y2 > = ð1 þ Re0 ÞRe90 r 1
Re0 þ Re90 1 > > > ¼ > ^ y2 ; ð1 þ Re0 ÞRe90 r 1
Substituting Eq. (12.68) into Eq. (12.59), it follows that
ð12:68Þ
12.7
Orthotropic Anisotropy
353
9 Re0 1 > > G¼ > ^ y2 > = ð1 þ Re0 ÞRe90 r 1
1 1 > > > H¼ > ^ y2 ; 1 þ Re0 r
ð12:69Þ
1
from which we have 9 G þ Re0 H Re0 ð1 þ Re90 Þ > > > ¼ = ð1 þ Re0 ÞH ð1 þ Re0 ÞRe90 > 1 ^ > > ¼ r1y2 ; ð1 þ Re0 ÞH
ð12:70Þ
Substituting Eq. (12.70) into Eq. (12.67), it follows that
^
r12 þ
Re0 ð1 þ Re90 Þ ^ 2 2Re0 ^ ^ ^ r r1 r2 ¼ r1y2 ð1 þ Re0 ÞRe90 2 1 þ Re0
ð12:71Þ
Equation (12.62) is rewritten as
1 1 1 2 2 ^2 ^ ^ ðG þ HÞ þ ðG þ H þ 4FÞ ðG HÞ r11 þ ðG þ HÞ ðG þ H þ 4FÞ r11 r22 4 4 2 4 4
1 1 1 ^2 ^2 þ ðG þ HÞ þ ðG þ H þ 4FÞ þ ðG HÞ r22 þ 6Lr12 ¼1 4 4 2 which is arranged as follows: 1 1 ^ ^ ^ ^ ðG þ HÞðr11 þ r22 Þ2 þ ðG þ H þ 4FÞðr11 r22 Þ2 4 4 1 ^2 ^2 ^2 ðG HÞðr11 r22 Þ þ 6Lr12 ¼ 1 2
ð12:72Þ
Denoting the angle measured in the counterclockwise direction from the principal axes of anisotropy to the principal stress as a and substituting the relations. ^
^
^
^
r11 þ r22 ¼ r1 þ r2 ; r11 r22 ¼ ðr1 r2 Þ cos 2a; ^ 2r12 ¼ ðr1 r2 Þ sin 2a
ð12:73Þ
354
12
Constitutive Equations of Metals
into Eq. (12.72), one has 1 1 ðG þ HÞðr1 þ r2 Þ2 þ ðG þ H þ 4FÞðr1 r2 Þ2 cos2 2a 4 4 1 3 ðG HÞðr21 r22 Þ cos 2a þ Lðr1 r2 Þ2 sin2 2a ¼ 1 2 2 which is rewritten as ðr1 þ r2 Þ2 2aðr21 r22 Þ cos 2a þ bðr1 r2 Þ2 cos2 2a þ 6
L 4 ðr1 r2 Þ2 ¼ GþH GþH
ð12:74Þ where a
GH G þ H þ 4F 6L ;b GþH GþH
ð12:75Þ
Here, denoting the yielding strength in the equi-two axis tension as r and that of the pure shear as s, it follows from Eq. (12.62) that r ðH þ GÞ1=2 ; s ð6LÞ1=2
ð12:76Þ
The substitution of Eq. (12.76) into Eq. (12.74) leads to ðr1 þ r2 Þ2 þ
r2 ðr1 r2 Þ2 2aðr21 r22 Þ cos 2a þ bðr1 r2 Þ2 cos2 2a ¼ ð2rÞ2 s ð12:77Þ
Equation (12.77) is extended to the following equation for the in-plane isotropy with the material constant m ð 1Þ. jr1 þ r2 jm þ
rm jr1 r2 jm ¼ ð2rÞm s
ð12:78Þ
Hill (1990) proposed the following extended orthotropic yield condition from Eqs. (12.77) and (12.78). jr1 þ r2 jm þ
r m
jr1 r2 jm h i þ jr21 þ r22 jðm=2Þ1 2aðr21 r22 Þ þ bðr1 r2 Þ2 cos 2a cos 2a ¼ ð2rÞm s
ð12:79Þ
12.7
Orthotropic Anisotropy
355
Equation (12.79) includes the five material constants, i.e. the yield stress r; s and the dimensionless number a; b; m. It is reduced to Eq. (12.77) for m ¼ 2 and to Eq. (12.78) for a ¼ b ¼ 0 (or a ¼ p=4). By use of Eqs. (12.73), (12.79) is rewritten in the anisotropic axes as follows:
m=2 r m
^ ^ ^ ^ ^2 m jðr11 r22 Þ2 þ 4r12
r11 þ r22 j þ
s
ðm=2Þ1 h i
^ ^ ^2 ^2 ^2 ^ ^ þ ðr11 þ r22 Þ2 þ 4r12 2aðr11 r22 Þ þ bðr11 r22 Þ2 ¼ ð2rÞm
ð12:80Þ Generally, the yield surface is described in the principal axes of anisotropy as follows: ^
f ðrij Þ ¼ FðHÞ
ð12:81Þ
where ^
^
^^
^
^
r ¼ R T rR; rij ¼ Rri Rsj rrs ^
^
^
Rij ðtÞ ei ej ðtÞ ¼ cos ei ; ej ðtÞ
ð12:82Þ
ð12:83Þ
^
ei ðtÞ are the base vectors taken to the directions of the principal axes of the orthotropic anisotropy. Needless to say, Eq. (12.81) is not a general tensor ^ expression but is merely the expression by the components. The variation of ei is ^ calculated using the following equation with the initial value of e i0 . Z ^ ^ ^ eidt ð12:84Þ e i ¼ ei0 þ ^
where ei is given by ^ ^ ^ e i¼ xa ei xa ¼ ei ei
^
ð12:85Þ
^
Here, the stress rate eij is given by ^
^
^
^
^
^
rij ¼ rij ¼ Rri Rsj rrs ¼ Rri Rsj ðrrs xarp rps þ rrp xaps Þ ^
ð12:86Þ
noting Eqs. (1.105) and (3.43) with Q ¼ R T . Various anisotropic yield surfaces in the plane stress state are proposed by the Barlat’s group (e.g. Barlat et al. 2007), Yoshida et al. (2015), etc.
356
12.7.1
12
Constitutive Equations of Metals
Representation of Isotropic Mises Yield Condition
The isotropic yield function described by Eq. (8.41) can be expressed in the following various forms. rffiffiffi rffiffiffi 3 0 3pffiffiffiffiffiffiffiffiffiffiffi r0rs r0rs f ðrÞ ¼ r ¼ kr k ¼ 2 2 rffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 02 3 02 02 02 02 r02 ¼ 11 þ r22 þ r33 þ 2 r12 þ r23 þ r31 2 rffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 ðr11 r22 Þ2 þ ðr22 r33 Þ2 þ ðr33 r11 Þ2 þ 6 r212 þ r223 þ r231 ¼ 2 rffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3 02 02 r02 ¼ 1 þ r2 þ r3 2 rffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 ¼ ðr1 r2 Þ2 þ ðr2 r3 Þ2 þ ðr3 r1 Þ2 ¼ F 2 ð12:87Þ eq
The combined test of the tensile stress rð¼ r11 Þ and the distortional stress sð¼ r12 Þ for a thin wall cylinder specimen is widely adopted for metal. In this case Eq. (12.87) is rewritten as pffiffiffi r2 þ ð 3sÞ2 ¼ F 2
ð12:88Þ
pffiffiffi Then, the Mises yield condition is shown by a circle of radius F in the ðr; 3sÞ plane. The visualization of the stress state can be realized in the space of three and less dimension. The stress state can be represented completely in the principal stress space when principal stress directions are fixed to materials and only the principal stress values change. In general, however, one must use the six-dimensional space or memorize the variation of the principal stress direction if the directions change. However, in the cases for which the number of independent variable components is less than three, such as the tension-distortion test described above and the plane stress and strain tests, the state of stress can be represented in the three and less dimensional stress space. The Ilyushin’s isotropic stress space (Ilyushin 1963) is convenient to depict the Mises yield surface, which depends only on the deviatoric stress, as explained below.
12.7
Orthotropic Anisotropy
357
The deviatoric stress tensor includes the five independent variables ðtrr0 ¼ 0Þ and thus the Mises yield surface in Eq. (12.87) is described by the independent components as follows: f ðrÞ ¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0 02 0 0 02 02 3r02 11 þ 3r22 þ 3r11 r22 þ 3 r12 þ r23 þ r31
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi s 2 0 3 0 2 1 0 0 02 02 þ 3 r11 þ r22 þ 3 r12 þ r23 þ r31 ¼ F r ¼ 2 11 2 and thus it can be rewritten as S21 þ S22 þ S23 þ S24 þ S25 ¼ F 2
ð12:89Þ
in the five-dimensional space with the axes pffiffiffi 1 pffiffiffi pffiffiffi pffiffiffi 3 S1 ¼ r011 ; S2 ¼ 3 r011 þ r022 ; S3 ¼ 3r012 ; S4 ¼ 3r023 ; S5 ¼ 3r031 2 2 ð12:90Þ Equation (12.89) exhibits the five-dimensional spherical super surface. Further, consider the expression of the Mises yield surface for the plane stress and strain conditions in the following.
12.7.1.1
Plane Stress State
The plane stress state fulfilling r3j ¼ 0 can be described in the three-dimensional space ðr11 ; r22 ; r12 Þ and thus the Mises yield condition (12.87) is described by the following equation. qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r211 r11 r22 þ r222 þ 3r212 ¼ F
ð12:91Þ
On the other hand, Eq. (12.91) can be described in the two-dimensional principal stress plane as follows: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r21 r1 r2 þ r22 ¼ F
ð12:92Þ
which is the section of the Mises yield condition cut by the plane r3 ¼ 0 and exhibits Mises’s ellipse in the principal stress plane ðr1 ; r2 Þ as shown in Fig. 12.15. It follows from the third equation of Eq. (1.322) that
358
Ⅱ
(1
3
( =120 ) F
(
12
Constitutive Equations of Metals
F,
2 F ) ( = 90 ) 3 ( F , F ) ( = 60 )
1 F , 1 ) ( =150 ) F 3 3
( 23 F , 1 F ) ( = 30 ) 3
( =180 ) F F ( = 0)
0 ( = 210 ) (
2 F, 3
1 ) F 3
(1
3
( = 240 ) ( F ,
F)
(
1 F, 3
F,
Ⅰ
1 ) ( = 330 ) F 3
F ( = 300 ) 2 ) ( = 270 ) F 3
Fig. 12.15 Mises yield surface in the plane stress condition (Thin curve describes Hill’s orthotropic Mises yield condition)
rffiffiffi 2 0 2 2 2 rm ¼ kr kF cos h þ p ¼ F cos h þ p 3 3 3 3
ð12:93Þ
pffiffiffiffiffiffiffiffi because of rm þ r03 ¼ 0 leading to rm ¼ r03 with kr0 k ¼ 2=3F. Substituting pffiffiffiffiffiffiffiffi Eq. (12.93) with kr0 k ¼ 2=3F into Eq. (1.322), one obtains 2 r1 ¼ F cos h þ 3 2 r2 ¼ F cos h þ 3
2 p þ 3 2 p þ 3
9 2 2 p > > F cos h ¼ pffiffiffi F sin h þ = 3 3 3 2 2 2 > ; F cos h p ¼ pffiffiffi F sin h > 3 3 3
ð12:94Þ
from which the coordinates of main points on the Mises’s ellipse are calculated as shown in Fig. 12.15. The thin curve shows the Hill’s orthotropic Mises yield surface in Eq. (12.61), which is rotated the principal axes of ellipse with the changes of its long and short radii from the isotropic Mises yield surface. Next, consider the Ilyushin’s isotropic stress space in which the variables in Eq. (12.90) are used. Here, in the present case fulfilling r3j ¼ 0 leading to S4 ¼ S5 ¼ 0 the Mises yield surface is represented by the sphere in the ðS1 ; S2 ; S3 Þ space, while it holds that
12.7
Orthotropic Anisotropy
359
S2
F
φ
−F
0
F
S1
−F Fig. 12.16 Mises yield surface in plane stress state without shear stress
9 > S1 ¼ ð3=2Þr011 ¼ ð3=2Þ½r11 ðr11 þ r22 Þ=3 ¼ r11 r22 =2 > = pffiffiffi 0 0 S2 ¼ 3 ð1=2Þr11 þ r22 > pffiffiffi pffiffiffi > ¼ 3fð1=2Þ½r11 ðr11 þ r22 Þ=3 þ ½r22 ðr11 þ r22 Þ=3g ¼ ð 3=2Þr22 ; ð12:95Þ Furthermore, in the case fulfilling r12 ¼ 0, the Mises yield surface is represented by the circle in the ðS1 ; S2 ; S3 Þ plane (Fig. 12.16). Here, setting S1 ¼ F cos /; S2 ¼ F sin /
ð12:96Þ
and substituting them into Eq. (12.95), it holds that 2 p r11 ¼ pffiffiffi F sin / þ ; 3 3
12.7.2
2 r22 ¼ pffiffiffi F sin / 3
ð12:97Þ
Plane Strain State
If the elastic strain rate can be ignored compared with the plastic strain rate in the
plane strain state, the following relation holds by substituting ep33 ¼ k r033 ¼ 0 into r0rr ¼ 0.
360
12
Constitutive Equations of Metals
1 r33 ¼ ðr11 þ r22 Þ 2
ð12:98Þ
Then, the Mises yield surface is described from Eq. (12.87)3 by rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffi r11 r22 2 3 þ r212 ¼ F 2
ð12:99Þ
which is represented by the Mohr’s circle in the plane of the normal and the shear stresses.
12.8
Subloading Surface Model with Orthotropic Anisotropy
Explicit equations required for applying the subloading surface model to the deformation analysis will be given in this section.
12.8.1
Subloading Surface with Orthotropic Anisotropy
The subloading surface for the yield condition in Eq. (12.54) is given by rffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 Fðr11 r22 Þ2 þ Gðr22 r33 Þ2 þ Hðr33 r11 Þ2 þ 2ðLr212 þ Mr223 þ Nr231 Þ ¼ RFðeeqp Þ 2
ð12:100Þ where _
rij ¼ rij þ R^cij
ð12:101Þ
(_ rij rij cij ^cij cij aij
ð12:102Þ
12.8
Subloading Surface Model with Orthotropic Anisotropy
12.8.2
361
Plastic Strain Rate
It follows from Eq. (12.100) that 8 @f 1 > > 22 Þ þ Hð 33 Þ ½Fð r11 r ¼ r11 r > > @ r 2f ðrÞ > 11 > > > @f > 1 > > 11 Þ þ Fð 33 Þ ½Gð r22 r ¼ r22 r > > > @ r 2f ðrÞ 22 > > > > 1 > @f > > 11 Þ þ Gð 22 Þ ½Hð r33 r ¼ r33 r < @ r33 2f ðrÞ > @f 1 > > L r12 ¼ > > @ r12 2f ðrÞ > > > > > @f 1 > > M r23 ¼ > > > @ r11 2f ðrÞ > > > > @f 1 > > : N r31 ¼ @ r31 2f ðrÞ
ð12:103Þ
|| ∂∂σf || =
ð12:104Þ
and 1 Ξ 2 f (σ )
where Ξ ≡
{[ F (σ 11 − σ 22 ) + H (σ 11 − σ 33 )]2 + [ F (σ 22 − σ 33 ) + G (σ 22 − σ 11 )]2 +[ H (σ 33 − σ 11 ) + G (σ 33 − σ 22 )]2 + 4[( Lσ 12 ) 2 + ( M σ 23 ) 2 + ( Nσ 31 ) 2 ]}
ð12:105Þ The components of the normalized outward-normal of the subloading surface are given by ⎧n11 = [ F (σ 11 − σ 22 ) + H (σ 11 − σ 33 )] / Ξ ⎪ ⎪n22 = [G (σ 22 − σ 11 ) + F (σ 33 − σ 11 )] / Ξ ⎪ ⎪n33 = [ H (σ 33 − σ 11 ) + G (σ 33 − σ 22 )] / Ξ ⎨ ⎪n12 = 2 Lσ 12 / Ξ ⎪n = 2M σ Ξ 23 / ⎪ 23 ⎪n = 2 Nσ 31 / Ξ ⎩ 31
ð12:106Þ
362
12
Constitutive Equations of Metals
Therefore, the plastic strain rate is given by the associated flow rule of the subloading surface as follows: ⎧ p • (σ − σ ) ⎪d11 = λ [ F 11 22 + H (σ 11 − σ 33 )] / Ξ ⎪ • ⎪d 22p = λ [G (σ 22 − σ 11 ) + F (σ 33 − σ 11 )] / Ξ ⎪ ⎪ p • (σ − σ ) ⎪d33 = λ [ H 33 11 + G (σ 33 − σ 22 )] / Ξ ⎨ ⎪d p = 2 λ• Lσ / Ξ 12 ⎪ 12 • ⎪ p ⎪d 23 = 2 λ M σ 23 / Ξ ⎪ • ⎪d p = 2 λ Nσ 31 / Ξ ⎩ 31
12.8.3
ð12:107Þ
Normal-Yield Ratio
The normal-yield ratio is calculated by the evolution rule in the plastic loading process. However, it must be calculated by the state variables, i.e. the stress, the isotropic hardening function, the kinematic hardening variable (back-stress), the elastic-core (similarity-ratio). Let the calculation method of the normal-yield ratio in the unloading process be given in the following. Equation (12.100) for the subloading surface is represented in the quadratic equation of R as follows: fFð^c11 ^c22 Þ2 þ Gð^c22 ^c33 Þ2 þ Hð^c33 ^c11 Þ2 þ 2ðL^c212 þ M^c223 þ N^c223 Þ 2½Fðeeqp Þ2 gR2 _
_
_
_
_
_
þ 2½Fðr11 r22 Þð^c11 ^c22 Þ þ Gðr22 r33 Þð^c22 ^c33 Þ þ Hðr33 r11 Þð^c33 ^c11 Þ _
_
_
þ 2ðLr12^c12 þ Mr23^c23 þ Nr31^c31 ÞR _
_
_
_
_
_
_2
_2
_2
þ Fðr11 r22 Þ2 þ Gðr22 r33 Þ2 þ Hðr33 r11 Þ2 þ 2ðLr12 þ Mr23 þ Nr31 Þ ¼ 0
ð12:108Þ R can be calculated by the following equation which is derived from Eq. (12.108). pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi b b2 ac R¼ ð12:109Þ a
12.8
Subloading Surface Model with Orthotropic Anisotropy
363
where 8 a Fð^c11 ^c22 Þ2 þ Gð^c22 ^c33 Þ2 þ Hð^c33 ^c11 Þ2 þ 2ðL^c212 þ M^c223 þ N^c223 Þ 2½Fðeeqp Þ2 > > > > _ _ _ _ _ _ < b Fðr c11 ^c22 Þ þ Gðr22 r33 Þð^c22 ^c33 Þ þ Hðr33 r11 Þð^c33 ^c11 Þ 11 r22 Þð^ > þ 2ðL^c12 þ M^c23 þ N^c31 Þ > > > : _ _ _ _ _ _ _ _ _ c Fðr11 r22 Þ2 þ Gðr22 r33 Þ2 þ Hðr33 r11 Þ2 þ 2ðLr212 þ Mr223 þ Nr231 Þ 2½Fðeeqp Þ2
ð12:110Þ
12.8.4
Elastic-Core Yield Ratio
The elastic-core (similarity) yield ratio designating the approaching degree of the elastic-core to the normal-yield surface is given by
ð12:111Þ
12.8.5
Cyclic Stagnation of Isotropic Hardening
The cyclic stagnation of the isotropic hardening described in Sect. 12.2 will be extended to the orthotropic anisotropy in the following. The normal-isotropic hardening surface in Eq. (12.28) is extended to the orthotropy as follows: rffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 gð~e Þ ¼ ep2 ep2 Fð~ep11 ~ep22 Þ2 þ Gð~ep22 ~ep33 Þ2 þ Hð~ep33 ~ep11 Þ2 þ 2ðL~ep2 12 þ M~ 23 þ N~ 31 Þ 2 ð12:112Þ p
364
12
Constitutive Equations of Metals
for which one has
ð12:113Þ
ð12:114Þ
leading to p p p p ⎧n%11 = [ F (ε% 11 − ε% 22 ) + H (ε% 11 − ε% 33 )] / Ζ ⎪ p p p p ⎪n%22 = [G ( ε% 22 − ε% 33 ) + F ( ε% 22 − ε% 11 )] / Ζ p p p p ⎪ ⎪n%33 = [ H (ε% 33 − ε%11 ) + G ( ε% 33 − ε% 22 )] / Ζ ⎨ p ⎪n%12 = 2 L ε% 12 / Ζ ⎪n% = 2M ε% p 23 / Ζ ⎪ 23 ⎪⎩n% 31 = 2 N ε% p31 / Ζ
ð12:115Þ
where Ζ ≡
{ [ F (ε%11p − ε%22p ) + H (ε%11p − ε% 33p )]2 + [G ( ε% 22p − ε% 33p ) + F (ε% 22p − ε% 11p )]2 + [ H ( ε% 33 − ε% 11 ) + G ( ε%33 − ε% 22 )]2 + 4[(L ε%12 ) 2 + (M ε% 23 ) 2 + ( N ε% 31) 2]} p
p
p
p
p
p
p
ð12:116Þ
Chapter 13
Constitutive Equations of Soils
The general importance of incorporating the subloading surface model, e.g. (1) the capability of the smooth elastic-plastic transition, (2) the unnecessity of the yield judgement in the loading criterion, (3) the automatic controlling function to pull-back the stress to the yield surface in the numerical calculation with large strain increment are explained in the former sections. The deformation in the elastic-plastic transition is quite small as the 0.2% plastic strain represented by the offset-value in metals which is rather brittle material. On the other hand, the plastic strain in the elastic-plastic transition is tremendously large as several percent plastic strain in soils. (4) the large plastic deformation in the elastic-plastic transition can be described by the suloading surface model, while it cannot be described by the conventional plasticity model with the yield surface enclosing the elastic domain represented by the Cam-clay model at all. Further, soils exhibit the remarkable softening, while only the hardening is induced in metals. The smooth peak-up and down stress vs. strain is described by the subloading surface model, although the abruptly rising-up and -down stress-strain curve is described by the conventional plasticity model. Therefore, the adoption of the subloading surface model is crucially important for the formulation of the plastic constitutive equation of soils. The elastoplastic constitutive equation of soils is formulated based on the subloading surface model in this chapter.
13.1
Isotropic Consolidation Characteristics
The isotropic consolidation characteristics is the one of the fundamentals in the constitutive equations of soils, which provides the base of isotropic hardening of soils.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_13
365
366
13 Constitutive Equations of Soils
(a) Linear relation between double-logarithm of volume and pressure Denoting the initial and the current volume as V and v, respectively, and the volume at the unloaded configuration to the stress-free state as V, the logarithmic volumetric strain ev in Eq. (4.141) is additively decomposed into the elastic logarithmic volumetric strain eev and the plastic logarithmic volumetric strain epv as follows: ev ¼ eev þ epv ðln J ¼ ln J e þ ln J p Þ
ð13:1Þ
where ev ln J ¼ ln ¼ ln
v vV v V v ¼ ln ¼ ln þ ln ; eev ln J e ¼ ln ; epv ln J p V V VV V V
V V ð13:2Þ
where J is the Jacobian representing the volume change as defined in Eq. (2.21) and, J e and J p are the elastic and the plastic parts, i.e. J det F ¼ det
@x v v V ¼ ; J e det Fe ¼ ; J p det Fp ¼ @X V V V
ð13:3Þ
which lead to the multiplicative decomposition ð13:4Þ
J ¼ JeJp
Fe and Fp are the elastic and the plastic part of the deformation gradient tensor F in the multiplicative decomposition F ¼ Fe Fp as will be defined in Chap. 17. It follows from Eqs. (13.1) and (13.2) that
e
p
ev ¼ ev þ ev
ð13:5Þ
where
ð13:6Þ As known from Eq. (13.6), the rate of the logarithmic volumetric strain is the trace of the strain rate measure d used in the hypoelasticity described in Eq. (7.109). Now, the ln v ln p linear relation (p ðtrrÞ=3: pressure) for the isotropic consolidation characteristics of soils was proposed by Hashiguchi (1947, 1955, 1958a, 1977, 2008), Hashiguchi and Ueno (1977) and later its conformity to test data was shown by Butterfield (1979). It is shown by the following equations and shown schematically in the upper part of Fig. 13.1.
13.1
Isotropic Consolidation Characteristics
367
Normal consolidation line: ðElastoplastic stateÞ Swelling and recompression line: ðElastic stateÞ
vy py 9 > ¼ ~k ln > = Vy py0 > v p > ln ¼ ~ j ln > ; p0 > V
ln
ð13:7Þ
where ðV; p0 Þ, ðVy ; py0 Þ and ðvy ; py Þ are the volumes and the pressures in the initial state, the initial yield state and the current yield state, respectively. py ð [ 0Þ is the so-called preconsolidated pressure. In addition, V is the volume in the unloaded state from the yielded state ðvy ; py Þ to the initial pressure p0 , while the unloaded v
ln v
V
Vy
Normal-consolidation line
Idealization
1
V
V
1
∼
Swelling line
∼
v vy
v vy 0
Swelling line
V Vy
p0
py0
p
py
(a) Isotropic consolidation of soils
p
1
p0
p
py0
py
ln p (b) Idealization by double-logarithmic linear relation of volume and pressure. Extension to negative pressure region
v V Vy
ln v
Swelling line
V Vy
Normal-consolidation line 1 ∼
V V
v vy
1
Swelling line
∼
v vy
1 (c) Isotropic consolidation of soils with negative pressure strength
(d) Extended double-logarithmic linear relation.
Fig. 13.1 Linear relation between double-logarithms of volume and pressure for isotropic consolidation of soils
368
13 Constitutive Equations of Soils
state corresponds to the intermediate configuration described in Sect. 17.1. Further, ~ ~ are the material constants prescribing the slopes of normal-consolidation k and j line and swelling line, respectively, in the ðln v; ln pÞ plane, while in the range of infinitesimal deformation they approximately coincide to the values of k and j in the e ln p linear relation (e: void ratio) used in the Cam-clay model as will be described subsequently in this section. Equation (13.7) is extended to the following equation so as to be applicable to the negative pressure range as shown in the lower part in Fig. 13.1, 9 Normal consolidation line: ln vy ¼ ~k ln py þ pe > > > Vy py0 þ pe = ðElastoplastic stateÞ v p þ pe > Swelling line: ln ¼ ~ j ln > > p 0 þ pe ; V ðElastic stateÞ
ð13:8Þ
where pe ð 0Þ is the material constant prescribing the negative pressure for which the volume becomes infinite, i.e. v ! 1 for p ! pe . The logarithmic volumetric strain in Eq. (4.141) and its elastic and plastic parts are given from Eq. (13.8) as follows: v v V v Vy vy V ¼ eev þ epv ¼ ln þ ln ¼ ln þ ln þ ln þ ln V Vy V V vy V V p þ pe py0 þ pe ~ py þ pe p0 þ pe ~ ln ¼ ~ j ln þ ~ j ln k ln j p0 þ pe p0 þ pe py0 þ pe py þ pe p þ pe p þ p y e ~Þ ln ð~ kj ¼ ~ j ln p0 þ pe py0 þ pe
ev ¼ ln
ð13:9Þ
leading to v p þ pe j ln eev ¼ ln ¼ ~ p0 þ pe V V py þ pe py ~Þ ln ~Þ ln kj epv ¼ ln ¼ ð~ ffi ð~ kj py0 þ pe py0 V
9 > > = > > ;
ð13:10Þ
noting pe py . Here, choosing py to coincide with the isotropic hardening function F, i.e. py ¼ FðHÞ; py0 ¼ F0 ðH0 Þ
ð13:11Þ
Equation (13.10) leads to p þ pe 9 > = p0 þ pe F ; ~Þ ln > epv ¼ ð~ kj F0 eev ¼ ~ j ln
ð13:12Þ
13.1
Isotropic Consolidation Characteristics
369
It follows from Eq. (13.12)2 that
epv ~ ~ kj
Fðepv Þ ¼ F0 exp
ð13:13Þ
which is represented in terms of the general symbol H for the isotropic hardening as FðHÞ ¼ F0 exp
H
~ ~ kj
ð13:14Þ
The strain rate is given from Eq. (13.12) as follows: 9 > p v > F > ~Þ > ev ¼ ¼ ~ j ð~ kj > v p0 þ pe F> > > > = p v V j eev ¼ ¼ ~ > > v V p0 þ pe > > > > > V p > F > ~ ; ~Þ ev ¼ ¼ ðk j F V
ð13:15Þ
Adopting Eq. (13.15) and limiting to the infinitesimal strain with the Hooke’s law in Sect. 7.5 hereinafter, let the elastic bulk modulus K and the elastic shear modulus G in Eq. (7.84) of the elastic modulus tensor E in Eq. (7.84) with Eq. (7.101) be assumed as follows:
rm
p
1 K ¼ ¼ ¼ ðp þ pe Þ; e e ~ j ev ev
0 1 jjr jj p þ pe n G¼ ¼ G0 p0 þ pe 2 jje e0 jj
ð13:16Þ
where the facts that the shear modulus G depends on the pressure are taken into account. G0 is the reference value of G at p ¼ p0 and nð 1; n ffi 0:5 for common soils) is the material constant. The shear modulus in Eq. (13.16) increases depending on the pressure and its dependence can be described by the power function as verified by Tatsuoka et al. (1978) through the comparison to various test data. In addition, it is formulated to be applicable up to the negative pressure range. The hyperelastic relation based on Eq. (13.16) will be described in Sect. 17.11.2. In the case that there is no information for the determination of G0 , i.e. G-value in p ¼ p0 , we may calculate it by G0 ¼
3 1 2m 3 1 2m p0 þ pe ; K0 ¼ ~ j 2 1þm 2 1þm
with an appropriate value of the Poisson’s ratio m, noting Table 7.1. (b) Linear relation between void ratio and logarithm of pressure
ð13:17Þ
370
13 Constitutive Equations of Soils
The following e ln p linear relation (e: void ratio) for the isotropic consolidation has been widely adopted for constitutive equation of soils after the Cam-clay models (Roscoe and Burland 1968; Schofield and Wroth 1968), although it possesses a lot of serious deficiencies described later in detail but unfortunately these deficiencies have not been recognized definitely. Normal consolidation line: ey ey0 ¼ k ln Swelling line: e e ¼ j ln
p p0
py 9 > = py0 > ;
ð13:18Þ
where the material constants k and j are the slopes of normal-consolidation and swelling lines, respectively, in the (e; ln p) plane as shown in Fig. 13.2, where ðe0 ; p0 Þ, ðey0 ; py0 Þ and ðey ; py Þ are the void ratios and the pressures in the initial, the initial yield and the current yield states, respectively. In addition, e is the void ratio in the unloaded state to the initial pressure p0 . However, the e ln p linear relation has the following physical impertinence. 1. The change of void ratio induced during a change of pressure from a certain pressure to other certain pressure (e.g. p0 to py0 in Fig. 13.2) along the swelling line is identical in spite of the plastic decrease of void ratio by the increase of a pre-consolidation pressure py . This defect is caused from the fact that the void ratio itself is adopted in vertical axis in the e ln p linear relation, although the logarithm of volume is adopted the ln v ln p linear relation. 2. The void ratio becomes negative if the pressure becomes large as py [ py0 expðey0 =kÞ or p [ p0 expðe=jÞ, noting Eq. (13.18). 3. The void ratio becomes infinitely large when the pressure approaches zero because of e ! 1 due to ln p ! þ 1 for p ! 0 in Eq. (13.18)2 . This defect causes the serious problem in the deformation analysis in which pressure Fig. 13.2 Linear relation between void ratio and logarithm of pressure
13.1
Isotropic Consolidation Characteristics
371
decreases as seen in the footing-settlement analysis: Pressure in soils in the periphery of footing decreases to zero in non-cohesive soils and even to negative in cohesive soils and the cyclic mobility analysis for liquefaction in which an accurate prediction of deformation under a quite low pressure is required. 4. The pressure is not related to the volume but to the void ratio or specific volume (void ratio plus unity), although the volumetric strain is not defined by the change of void ratio but by the change of volume. Note here that the ratio of the current volume to the initial volume can be transformed to the ratio of the current specific volume to the initial specific volume as v=V ¼ ð1 þ eÞvs ðpÞ=½ð1 þ e0 Þvs ðp0 Þ ¼ ð1 þ eÞ=ð1 þ e0 Þ under the assumption of the incompressibility of soil particles themselves, i.e. vs ¼ const: where vs designates the volume occupied by the soil particles. 5. The e ln p linear relation is not formulated by the logarithm of ratio of volumes but it is given by the difference of void ratios. Therefore, it does not fit to the logarithmic strain but it fits to the nominal strain, so that the nominal strain is obliged to be adopted in the derivation of volumetric strain from the e ln p linear relation. However, the nominal volumetric strain does not coincide with
the time-integration of trace of the strain rate e used in the elastoplastic constitutive equations in the subsequent chapters as known from Eq. (4.141) and it is impertinent to the description of large deformation, because (a) the strain is merely minus one even when the material vanishes completely by extreme volumetric compression as known from Eq. (4.141)2 , (b) it does not satisfy the superposition rule as known in Eq. (4.140)2 , (c) the sum of three longitudinal strains in the orthogonal directions does not coincide with the volumetric strain as shown in Eq. (4.141)2 . Because of the deficiency in 5, the e ln p linear relation is obliged to be described by the nominal volumetric strain ev in Eq. (4.141)2 which is limited to the description of infinitesimal deformation as was delineated in Sect. 4.5. The nominal volumetric strain ev is additively decomposed into the elastic and the plastic parts as follows: ev ¼ eev þ epv
ð13:19Þ
where ev ¼
vV e vV p V V ; ev ¼ ; ev ¼ V V V
ð13:20Þ
Here, the italic letter e is used for the logarithmic strain but the roman letter e is used for the nominal strain applicable only to the infinitesimal deformation in order to distinguish them as was described in Sect. 4.5. It should be noted that
the nominal strain cannot be related to the strain rate e in the exact sense as known by ev ¼ tr e ¼ v =v 6¼ v =V ¼ ev . The following relations are used by substituting the specific volume instead of the volume into Eq. (13.20) on the approximation at the sacrifice of the above-mentioned deficiency 4.
372
13 Constitutive Equations of Soils
9 e e0 > > > 1 þ e0 > > = e e e ev ffi 1 þ e0 > > > e e0 ðey0 e0 Þ þ ðey ey0 Þ þ ðe ey Þ > > ; epv ffi ¼ 1 þ e0 1 þ e0 ev ffi
ð13:21Þ
where the nearly equal symbol ffi is used to specify the approximation by the incompressibility of soil particles, noting that the symbols v and V in Eq. (13.20) are not the specific volumes but the volumes. Substituting Eq. (13.18) into Eq. (13.21), one has 9 j p k j py > ev ffi ln ln > > > 1 þ e0 p0 1 þ e0 py0 > = j p e ev ffi ln ð13:22Þ 1 þ e 0 p0 > > > j py0 k py j p0 k j py > > ; epv ffi ln ln ln ¼ ln 1 þ e0 p0 1 þ e0 py0 1 þ e0 py 1 þ e0 py0 from which the nominal volumetric strain rate is given by
9 e j p k j py > > > ev ¼ ffi ¼ > 1 þ e0 1 þ e 0 p 1 þ e0 p y > > > > = p e e j v V eev ¼ V V ffi ¼ > 1 þ e0 1 þ e0 p > > > > > e k j py p > > V ; ¼ ev ¼ V ffi 1 þ e0 1 þ e0 p y
v V
ð13:23Þ
Equation (13.23) for the rates of volumetric strain rate and pressure is adopted widely for elastoplastic constitutive equations of soils after the Cam-clay model which is quite primitive model, while its basic concept was already clarified by Drucker and Prager (1952) and Drucker et al. (1957). Nevertheless, it causes the further deficiencies as follows: 6. It is not derived exactly but approximately from the e ln p linear relation. 7. It cannot be adopted to describe finite deformation since it is derived based on the definition of nominal strain. 8. The tangent elastic bulk modulus is given by
K
p
eev
¼
1 þ e0 p j
ð13:24Þ
from Eq. (13.23). Here, the tangent elastic bulk modulus K depends on the initial void ratio e0 . It has a crucial physical impertinence: “The larger the initial
13.1
Isotropic Consolidation Characteristics
373
void ratio, the larger is the elastic bulk modulus. In other words, the looser the soil, the more difficult to be compressed”. Eventually, it is concluded that the e ln p linear relation is inadequate physically and mathematically for formulation of constitutive equations for finite deformation of soils. On the other hand, all the deficiencies in the e ln p linear relation can be excluded in the ln v ln p linear relation. The nominal volumetric strain and its elastic and plastic parts are related to the strain rate and its elastic and plastic parts used in the elastoplastic constitutive equation as follows:
ð13:25Þ
which are of complicated forms containing the Jacobian and its elastic and plastic parts. Equation (13.25) have been used for constitutive equations under the approximation of J e ffi 1; J ffi J p by some workers (cf. Asaoka et al. 1997; Zhang et al. 2007). However, it is merely the approximation and thus the physical property cannot be remedied, so that it is limited to the description of infinitesimal deformation in spite of the complexity. Then, this modification causes no good and much harmful. A lot of constitutive equations of soils adopt the e ln p linear relation, so that unfortunately they are applicable to the description of infinitesimal deformation. On the other hand, the ln v ln p linear relation is applicable to the description of finite deformation. It has been adopted not only in hypoelastic-based plastic constitutive equations of soils (cf. Hashiguchi 1974, 1978; Hashiguchi and Ueno 1977; Hashiguchi and Chen 1998) but also in hyperelastic-based plastic constitutive equations of soils based on the additive decomposition of infinitesimal strain into elastic and plastic parts (e.g. Houlsby 1985; Collins and Hilder 2002; Coombs and Crouch 2011; Coombs et al. 2013). Further, it has been adopted in hyperelastic-based plastic constitutive equations of soils under the multiplicative decomposition of deformation gradient (e.g. Borja and Tamagnini 1998; Yamakawa et al. 2010, 2021) as will be described in Sect. 20.9.2. The material ~ in the ln v ln p linear relation are approximately related to k and constants ~ k and j ~ ffi j=ð1 þ e0 Þ as described in j in the e ln p linear relation by ~ k ffi k=ð1 þ e0 Þ; j Appendix F, while numerous test data have been accumulated for the latter.
374
13.2
13 Constitutive Equations of Soils
Yield Conditions
Yield conditions for soils are described in this section.
13.2.1
Yield Functions
Various yield conditions of soils have been proposed to date. The functions f ðrÞ of the yield condition f ðrÞ ¼ FðHÞ in Eq. (8.16) can be reduced to the following common form for soils which depend on the stress ratio and the Lode’s angle because of the frictional material. f ðrÞ ¼ pgðg=MÞ
ð13:26Þ
f ðrÞ ¼ pgðgm Þ
ð13:27Þ
or
where g
r0 ; p
g jjgjj;
gm
g M
ð13:28Þ
0
M is the stress ratio jjgjj in the maximum state of jjr jj in the fixed yield surface, i.e. the critical state and is called the critical state stress ratio. It is premised in Eq. (13.27) that the yield surface passes through the isotropic compression state and the null stress point. Then, the function gðgm Þ fulfill the following conditions g ¼ 1 for g ¼ 0 g ! 1 for g ! 1
ð13:29Þ
Furthermore, note that the following equality holds in the critical state (cf. Appendix G). g0 ð1Þ ¼ gð1Þ
ð13:30Þ
which is illustrated in Fig. 13.3. Denoting the p-value in the critical state as pcr , the following equation is obtained by substituting Eq. (13.27) into Eq. (8.16) at gm ¼ 1. pcr ¼ F=gð1Þ
ð13:31Þ
13.2
Yield Conditions
375
Fig. 13.3 Function gðgm Þ in yield surface of soils
The function gðgm Þ for the yield surfaces proposed in the past are given as follows: (1) Original Cam-clay model (Schofield and Wroth 1968)
jjr0 jj=p g ¼ expðgm Þ; i:e: p exp M
¼F
pcr ¼ F=e
ð13:32Þ ð13:33Þ
where e is the Napier’s constant, i.e. e ¼ 2:71828. (2) Modified Cam-clay model (Burland 1965; Roscoe and Burland 1968) "
g¼
1 þ g2m ;
0 # jjr jj=p 2 i:e: p 1 þ ¼F M pcr ¼ F=2
ð13:34Þ ð13:35Þ
(3) Hashiguchi model (Hashiguchi 1972, 1985a) " # g2m 1 jjr0 jj=p 2 ; i:e: p exp ¼F g ¼ exp 2 2 M
pffiffiffi pcr ¼ F= e
ð13:36Þ ð13:37Þ
376
13 Constitutive Equations of Soils
The functions in (1) and (3) are unified as follows:
0 gnm 1 jjr jj=p n ; i:e: p exp ¼F g ¼ exp n n M
ð13:38Þ
pcr ¼ F= expð1=nÞ
ð13:39Þ
where Eqs. (13.32) and (13.36) are given for n ¼ 1 and 2, respectively. The above-mentioned three yield surfaces are shown for the axisymmetric stress states in Fig. 13.4, where 9 p 13 trr ¼ 13 ðra þ 2rl Þ > >
3 > 1 3 0 > 2 ra 3 ðra þ 2rl Þ ¼ 2 ra = q rl ra ¼ 0 ð13:40Þ 3 rl 13 ðra þ 2rl Þ ¼ 3rl > qffiffi qffiffi > > > ; jjr0 jj ¼ 2jrl ra j ¼ 2jqj 3
3
where ra and rl are the axial and the lateral stress, respectively, in the triaxial compression/extension state, while the fact that the value of M in the axisymmetric compression is larger than that in the axisymmetric extension is taken into account in the following. Among the above-mentioned yield surfaces, only the yield surface of the modified Cam-clay model in Eq. (13.34) does not include any corner so that the singularity of the normal direction of the surface is not induced. Equation (8.16) with Eqs. (13.27) and (13.34) is rewritten as
p ðF=2Þ F=2
2
jjr0 jj 2 þ ¼1 MF=2
ð13:41Þ
Axisymmetric compression
q
Hashiguchi (1972) model Modified Cam-clay model Original Cam-clay model
Mc
0
1
F Me e
F 2
F
e
F
p
Axisymmetric extension
Fig. 13.4 Various yield surfaces of soils
13.2
Yield Conditions
377
leading to jjr0 jj ¼ M
F F for p ¼ ðcritical stateÞ 2 2
Therefore, the magnitude of the deviatoric stress jjr0 jj depends on the material function M which is largest and smallest in the triaxial compression and the triaxial extension, respectively, in experiments. The pertinent equation of M will be formulated in the next subsection.
13.2.2
Critical State Surface Taken Account of Third Deviatoric Invariant
The simple equation of the critical state surface taken account of the third deviatoric invariant will be described in this subsection (Hashiguchi 2002). The Coulomb-Mohr criterion is given as follows (see Fig. 13.5):
ra rl ra þ rl ¼ sin /c ðupper sign: Compresstion; lower sign: ExtenstionÞ 2 2 ð13:42Þ
where ra ð\0Þ and rl ð\0Þ are the axial stress and the lateral stress, respectively, and /c is the angle of the internal friction. Equation (13.42) is generalized to the three-dimensional state:
Fig. 13.5 Coulomb-Mohr yield criterion
378
13 Constitutive Equations of Soils 3 2 Y ri þ rj 2 ri þ rj sin /c ¼0 2 2 i;j¼1
ð13:43Þ
which is shown by the hexagon with the unequal radius length in the deviatoric stress plane as shown in Fig. 13.6. The expression of the Coulomb-Mohr criterion by the stress invariants is rather complicated as was given by Hashiguchi (1972). Equation (13.42) is rewritten in terms of the stress ratio Ral
ra rl
[ 1 for triaxial compression \1 for triaxial extenssion
ð13:44Þ
as follows: sin /c ¼
Ral 1 ðupper: Compresstion; lower: ExtenstionÞ Ral þ 1
ð13:45Þ
i.e. Ral ¼
1 sin /c ðupper: Compresstion; lower: ExtenstionÞ 1 sin /c
ð13:46Þ
On the other hand, the ratio of the second deviatoric invariant to the pressure is given from Eq. (13.42) as follows: pffiffiffi Ral 1 jjr0 jj ðupper: compresstion; lower: extenstionÞ ¼ 6 p Ral þ 2
ð13:47Þ
noting qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0 r0a 2 þ 2rl 2 ½ra ðra þ 2rl Þ=3 2 þ 2½rl ðra þ 2rl Þ=3 2 jjr0 jj ¼ ¼ p ðra þ 2rl Þ=3 ðra þ 2rl Þ=3 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ½ra ðra þ 2rl Þ=3 2 þ 2½rl ðra þ 2rl Þ=3 2 =ðrl Þ ¼ ½ðra =rl Þ þ 2Þ=3 =ðrl Þ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ½Ral ðRal þ 2Þ=3 2 þ 2½1 ðRal þ 2Þ=3 2 ¼ ðRal þ 2Þ=3 pffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi2 6 ðRal 1Þ ¼ Ral þ 2
13.2
Yield Conditions
379
Substituting Eq. (13.46) into Eq. (13.47), we have 8 pffiffiffi > 2 6 sin /c p ffiffi ffi > 0 Mc for triaxial compresstion jjr jj 2 6 sin /c < 3 sin /c pffiffiffi ¼ ¼ > 2 6 sin /c 3 sin /c p > : Me for triaial extension 3 þ sin /c
ð13:48Þ
Now, in order to formulate the appropriate smooth critical state surface fulfilling the convexity, we consider the conical surface given by 0
jjr jj ¼ Mðcos 3hr Þ p
ð13:49Þ
where pffiffiffi 0 cos 3hr 6trtr 3 ¼
1 for hr ¼ p=3 ðtriaxial compressionÞ 1 for hr ¼ 0 ðtriaxial extensionÞ
ð13:50Þ
0
0
tr
r 0 ðjjtr jj ¼ 1Þ 0 jjr jj
ð13:51Þ
noting Eq. (1.324) for the Lode angle hr . Now, we will find the function Mðcos 3hr Þ in Eq. (13.49) by which a smooth convex conical surface is depicted. Then, assume the following simple equation which fulfills Eq. (13.48) in the triaxial-compression. pffiffiffi 2 6 sin /c Mðcos 3hr Þ ¼ 3 ð1 cÞ sin /c þ c sin /c cos 3hr 8 compresstion ð cos 3hr ¼ 1Þ > < Mc forptriaxial ffiffiffi
¼ Mc for c ¼ 0 2 6 sin /c > ¼ for triaxial extention ðcos 3hr ¼ 1Þ : 3 ð1 2cÞ sin /c Me for c ¼ 1
ð13:52Þ where cð0 c 1Þ is the material parameter. The variable Mðcos 3hr Þ in Eq. (13.52) coincides with Mc in the triaxial compression state but it is smaller than Me for c [ 1 and coincides with Me for c ¼ 1 in the triaxial extension state. In what follows, we will find the value of the material constant c which leads to the smallest value M in the triaxial extension state, satisfying the convexity of the conical surface. In order that the surface in Eq. (13.49) with Eq. (13.52) is convex, the curve formed by cutting it by the p plane (p constant plane) must satisfy following convexity condition (cf. Eq. (A.46) in Appendix H).
380
13 Constitutive Equations of Soils
2 1 d 1 1 þ 2 ¼ pffiffiffi ½3 ð1 cÞ sin /c þ c sin /c cos 3hr 9c sin /c cos 3hr M dhr M 2 6 sin /c 1 ¼ pffiffiffi ½3 ð1 cÞ sin /c 8c sin /c cos 3hr 0 2 6 sin /c
ð13:53Þ noting 9 ! d 1 d 3 ð1 cÞ sin /c þ c sin /c cos 3hr 3c sin 3hr > > pffiffiffi pffiffiffi > ¼ ¼ > = dhr M dhr 2 6 sin /c 2 6 2 > d 1 9c cos 3hr > > pffiffiffi > ¼ ; 2 M dhr 2 6 In addition, Mðcos 3hr Þ takes the maximum value at hr ¼ p=3 (triaxial compression) and the minimum value at the triaxial extension at hr ¼ 0 (triaxial extension) (see Fig. 13.6), noting pffiffiffi dM 6 6c sin2 /c ¼ sin 3hr dhr ½3 ð1 cÞ sin /c þ c sin /c cos 3hr 2
Fig. 13.6 Sections of the critical state surfaces of soils in p–constant plane
ð13:54Þ
13.2
Yield Conditions
381
9 8 þ for hr ¼ p=3 0 = > > > > 0 for hr ¼ p=3 maximal for triaxial compression > > ; < for hr ¼ p=3 þ90 ¼ > for hr ¼ 0 = > > > > minimal for triaxial extension 0 for hr ¼ 0 > : ; þ for hr ¼ þ 0
ð13:55Þ
Therefore, the radius of the p section of the surface is the maximum and the minimum in the triaxial compression and the extension, respectively. Then, by setting hr ¼ 0 in Eq. (13.53), the following inequality must hold for the convexity of the conical surface in Eq. (13.49) with Eq. (13.52). c
3 sin /c 7 sin /c
ð13:56Þ
Equation (13.52) is reduced to the following equations for the particular values of the material constant c. 8 < Mc for pffifficffi ¼ 0 2 6 sin /c ð13:57Þ M¼ for c ¼ 1 : 3 þ sin /c cos 3hr Equation (13.57)1 leads to the Drucker-Prager yield condition and (13.57)2 was proposed by Satake (1972) and later by Gudehus (1973) and Argyris et al. (1973) as the failure criterion of soils. However, the frictional angle /c in the triaxial compression is limited to the following inequality in order to fulfill the convexity in Eq. (13.56) for c ¼ 1. 3 sin /c ; i.e:/c /ch 22:02 8
ð13:58Þ
Therefore, Eq. (13.57)2 is inappropriate to the material with the high frictional angle /c 22:02 . Now, let the rigorous function of Mðcos 3hr Þ be formulated, which satisfies the convexity condition in Eq. (13.53) for a whole range of frictional angle /c . Then, by choosing c to be the largest value satisfying the inequality in Eq. (13.56) for the convexity condition, i.e. c¼
3 sin /c 7 sin /c
ð13:59Þ
382
13 Constitutive Equations of Soils
M is finally determined p asffiffiffi pffiffiffi 2 6 sin /c 7 2 6 sin /c ¼ M¼ /c /c 8 þ cos 3hr 3 sin /c 3 1 3sin sin /c þ 3sin 7 sin / 7 sin / sin /c cos 3hr c
c
i.e. pffiffiffi 7 14 6 sin /c Mðcos 3hr Þ ¼ Mc ¼ ð8 þ cos 3hr Þð3 sin /c Þ 8 þ cos 3hr
ð13:60Þ
which was proposed by Hashiguchi (2002) (see Fig. 13.6). Equation (13.60) is reduced to the following equation in the triaxial extension. pffiffiffi 7 2 6 sin /c 7 7 3 þ sin /c M¼ ¼ Mc ¼ Me for hr ¼ 0 ðtriaxial extensionÞ 9 3 sin /c 9 9 3 sin /c ð13:61Þ Therefore, M for Eq. (13.60) is smaller or larger than Me for the Coulomb-Mohr criterion in the triaxial extension state ðh ¼ 2nðp=3ÞÞ as known from the inequality 8 > < [ Me forp/ffiffiffic [ /ch M ¼ Me ¼ 2 6=9 for /c ¼ /ch ¼ 22:02 ð sin /c ¼ 3=8Þ for hr ¼ 0 ðtriaxial extensionÞ > : \Me for /c \/ch
ð13:62Þ The difference of the variable Mðhr Þ in the tri-axial compression and -extension states in most of test data is not so large as given by the Coulomb-Mohr criterion but it would be expressed approximately by Eq. (13.60). The other functions for M with some physical meanings are referred to Matsuoka and Nakai (1974), Nakai and Mihara (1984) and Matsuoka et al. (1999). However, their applications to general deformation analysis possess difficulties because they are formulated by the cubic equation of the radius of the failure surface in the deviatoric stress plane. Various modifications of the Coulomb-Mohr criterion have been proposed hitherto, e.g. Argyris et. al. (1973), William and Warnke (1974), Lin et al. (1987), Klisinski et al. (1991), Bardet (1990), Andrade and Borja (2006) but the most rigorous modification would be given by Eq. (13.49) with Eq. (13.60) only which the fulfillment of the convexity condition is verified.
13.3
13.3
Subloading Surface Model for Soils
383
Subloading Surface Model for Soils
The initial subloading surface model with the isotropic hardening is described in Chap. 9. In what follows, the simple initial subloading surface model with the modified Cam-clay yield surface will be formulated and compared with the other soil models. The following functions are adopted based on Eqs. (13.14), (13.34) and (13.60). "
0 # g 2 jjr jj=p 2 ¼ p 1þ f ðrÞ ¼ p 1 þ M M H dF F FðHÞ ¼ F0 exp ¼ ; F0 ~j dH ~k j ~ ~ k
9 > > > > > > > > > =
> > > H ¼ epv ; H ¼ tr ep ; fHn ð¼ H = kÞ ¼ trn > > > > > 7 > ; Mðcos 3hr Þ ¼ Mc 8 þ cos 3hr
ð13:63Þ
p
The plastic modulus M in Eq. (9.24) is given for Eq. (13.63) as p
M
trn UðRÞ n :r þ ~ R ~ kj
ð13:64Þ
and thus the plastic strain rate is described by p
e ¼
n :r n :r n p n¼ trn UðRÞ M þ n :r ~ R ~ kj
ð13:65Þ
and thus
n :r n trn UðRÞ n :r þ ~ R ~ kj
e ¼ E1 : r þ
ð13:66Þ
Here, note that the subloading hardening, i.e. n : r [ 0 in Eq. (9.31) can be induced over the critical state line, fulfilling tr n 0 since the positive quantity p UðRÞ=R is contained in the plastic modulus M . The partial derivatives of the yield function in Eq. (13.63) are given as follows:
g 2 g 2 @f ðp; g; MÞ @f ðp; g; MÞ @ pg ¼ 1þ ¼ P 1þ ¼ 2 2 ð13:67Þ ; @p @p M M @g M
384
13 Constitutive Equations of Soils
@f ðp; g; MÞ p g 2 ¼ 2 @M M M @p @ðtrrÞ=3 1 ¼ ¼ I @r @r 3 0 @g @jjr jj=p 1 1 0 1 0 1 0 ¼ ¼ 2 ptr þ jjr jjI ¼ tr þ gI @r @r p 3 p 3 pffiffiffi @M 14 6 sin /c M ¼ ¼ 2 @ cos 3hr 8 þ cos 3hr ð3 sin /c Þð8 þ cos 3hr Þ 3 sin /c ¼ pffiffiffi M2 14 6 sin /c pffiffiffi @ cos 3hr 3 6 0 2 1 1 0 ¼ 0 tr pffiffiffi cos 3ht jjtr jj2 I @r jjr jj 3 6
ð13:68Þ ð13:69Þ ð13:70Þ
ð13:71Þ
ð13:72Þ
@f ðrÞ @f ðp; g; MÞ @f ðp; g; MÞ @p @f ðp; g; MÞ @g ¼ ¼ þ @r @r @p @r @g @r @f ðp; g; MÞ @M @ cos 3hr þ @M @ cos 3hr @r g 2 1 pg 1 0 1 Iþ2 2 tr þ gI ¼ 1þ 3 M M p 3 ! pffiffiffi p g 2 3 sin /c 3 6 0 2 1 1 0 2 tr pffiffiffi cos 3ht jjtr jj2 I pffiffiffi M2 0 jjr jj M M 3 14 6 sin /c 6 pffiffiffi g 3 sin /c 0 2 1 g 2 g 0 1 1 0 2 pffiffiffi ¼ 1 t pffiffiffi cos 3ht jjtr jj I I þ 2 2 tr þ 3 6 3 M M M 7 6 sin /c r 3 6
ð13:73Þ noting the partial derivative formulae in Sect. 1.16. The predictions of the drained triaxial compression behavior of soils under the constant lateral stress by the subloading surface model in Eq. (13.66) and the conventional Cap model (Roscoe and Burland 1968; Schofield and Wroth 1968) are depicted in Fig. 13.7, where ea is the axial strain. The stress path is given by q = 3(p + rl) (rl = const.) because of p = –(ra + rl) / 3, q = rl – ra. Here, the curves of axial stress and volumetric strain vs. the axial strain in the loading from the heavily and the lightly over-consolidated states, i.e. points o and o0 , respectively are shown in this figure. In the loading from the lightly over-consolidated state o0 , the volume contraction proceeds and the axial stress increases monotonically up to the critical state. The abrupt transition from the elastic to the plastic state is predicted by the conventional Cam-clay model. On the other hand, the smooth behavior is predicted always by the subloading surface model as observed in experiments.
%
+
+
+
+
+
≥0
− t r n + U ) n:σ M p ≡( % R { λ − κ% { ≥0
%
%
+
Subloading Surface Model for Soils
Fig. 13.7 Comparison of predictions of triaxial compression behavior under constant lateral stress by the conventional Cam-clay model and the subloading surface model
%
13.3 385
386
13 Constitutive Equations of Soils
Next, consider the loading from the heavily over-consolidated state o. The elastic behavior is predicted until the stress reaches the yield surface and then the stress is predicted to decrease abruptly toward the critical state, exhibiting the intense softening by the Cam-clay model. On the other hand, the following realistic deformation behavior is predicted by the subloading surface model. (1) The first term in the plastic modulus in Eq. (13.64) decreases from the positive value to zero because of trn\0 in the process from the initial state up to the critical state line. On the other hand, the second term is always positive. Therefore, the plastic modulus is kept to be positive in this process, i.e. p ~Þ [ 0 ! 0, U=R [ 0, M [ 0 for o ! c. Both of the normal trn=ð~ kj
hardening F [ 0 in Eq. (9.30) and the subloading hardening n : r [ 0 in Eq. (9.31) proceed in this process. (2) The first term becomes zero but the second term is positive at the point on the p ~Þ ¼ 0, U=R [ 0, M [ 0 at the point c. The critical state line, i.e. tr n=ð~ kj normal and the subloading hardenings proceed continuously rising up along this line. (3) The first term tends to be negative but the second term is positive, so that the plastic modulus is kept to be positive until the stress reaches the peak, i.e. p ~j ~Þ\0, U=R [ 0, M [ 0 for c ! p. The normal softening F \0 trn=ðk
but the subloading hardening n : r [ 0 proceed in this process. ~Þ ð\0Þ reaches the minimum value but the second (4) The first term trn=ð~ kj term U=Rð [ 0Þ is always positive so that the sum of them becomes zero, p canceling each other, resulting in M ¼ 0 at the peak stress point p. The normal softening proceeds continuously but the subloading hardening ceases at this
point. The fact that ðev =j ea jÞmax : an the peak stress qmax : are induced concurrently has been recognized after (Taylor 1948). (5) The first term is negative but tends to increase toward zero while the second term is always positive but decreases continuously so that the sum of them ~Þ\0, becomes negative, exhibiting the subloading softening, i.e. tr n=ð~k j p p U=R [ 0, M \0 and then reducing to the critical (residual) state ðM ¼ 0Þ as
p ! c. Both of the normal softening F \0 and the subloading softening n : r \0 proceed in this process. (6) The first and the second terms reach zero at the residual (critical) state, i.e. p ~Þ ¼ 0, U=R ¼ 0, M ¼ 0 at the final point c. Both of the normal tr n=ð~ kj and the subloading softening cease in this state. (7) The sign of the plastic volumetric strain rate is identical to that of tr n, whilst the sign of the elastic volumetric strain rate is opposite to that of the rate of pressure. The volume contraction is induced by the elastic volume contraction due to the increase of pressure in the initial stage of loading. Thereafter, the rate of plastic volume expansion tends to be larger than the rate of elastic volume contraction after passing through the critical state line, so that the rate of volume expansion
13.3
Subloading Surface Model for Soils
387
n DP Plastic potential surface assumed in non-associated Drucker-Prager model
n Normal-yield surface
σ Subloading surface
Drucker-Prager yield surface
0
Fig. 13.8 Outward-normal of subloading surface coinciding approximately with plastic potential surface assumed in Drucker-Prager model, nDP : unit normal to Drucker-Prager’s yield surface, n: unit normal to subloading surface and plastic potential surface used for Drucker-Prager’s yield surface (trn < trnDP)
e
p proceeds. However, reaching the peak stress ðn : r ¼ M ¼ ev ¼ 0; p ¼ 0Þ at which the subloading surface expands at most and thus tr n becomes maximum
p
leading to ev ¼ ev ¼ max: by virtue of the “associated flow rule”, the maximum
ratio of volume expansion strain rate to axial strain rate, i.e. ev =j ea j ¼ max: are induced. This fact was indicated by Taylor (1948) based on the experimental evidence. Eventually, these typical deformation behavior in normally-consolidated and over-consolidated states can be described pertinently by the initial subloading surface model. Here, it should be emphasized that these realistic predictions are attained by the subloading surface model without resort to the non-associated flow rule. As described above, the conventional Cam-clay model predicts unrealistically high yield stress in the over-consolidated state. Then, the Cap model in which the over-consolidated side of Cam-clay yield surface is replaced by the Drucker-Prager yield surface (Drucker and Prager 1952) in Fig. 13.8 is widely used. However, the Cap model possesses various drawbacks described in the following. 1. The Cap model falls within the framework of conventional plasticity with the yield surface enclosing the purely elastic domain. Therefore, it predicts the stress–strain curve which rises up steeply (elastically) to the peak stress, and subsequently the stress decreases suddenly exhibiting a strong softening. In contrast, the subloading surface model can describe the realistic stress–strain curve with the smooth elastic–plastic transition since the plastic strain rate develops gradually as the stress approaches the yield surface. 2. The Cap model necessitates the yielding judgment, i.e. the judgment whether or not the stress lies on the yield surface in addition to the judgment on the direction of strain rate in the loading criterion as shown in Eq. (8.57). In contrast, the yielding judgment is not required in the subloading surface model since the stress lies always on the subloading surface playing the role of the loading surface as shown in Eq. (9.29).
388
13 Constitutive Equations of Soils
3. The Cap model requires the particular computer algorithm for pulling-back the stress to the yield surface when the increments of stress or strain with finite magnitudes are input in numerical calculations. Otherwise, it is obliged to perform numerical calculations with quite infinitesimal loading increments. In contrast, the subloading surface model enables numerical calculations with finite loading increments without incorporation of particular computer algorithm since it possesses an automatic controlling function to attract the stress to the yield surface in the loading process as was described in Sect. 9.3. 4. The cap model additionally adopts the plastic potential surface of the conical shape to predict a dilatancy angle lower than that predicted by applying the associated flow rule to the Drucker-Prager yield surface, avoiding an unrealistically large plastic volume expansion. Then, the formulation is obliged to adopt the non-associated flow rule which causes the serious physical drawbacks as described in Subsect. 8.6.4 (Hashiguchi et al. 1991). In contrast, the subloading surface model can use the associated flow rule, whereas the outward-normal n of the subloading surface in the current stress is approximately identical to the outward-normal nDP of the plastic potential surface adopted in the Drucker-Prager model as shown in Fig. 13.8. 5. The cap model is obliged to adopt the non-associativity for the Drucker-Prager yield surface as described in 4. Therefore, it is accompanied with the asymmetry of the elastoplastic stiffness modulus tensor Kep as shown in Eq. (8.40). This fact engenders the complexity in the formulation of variational principle and thus the difficulty in the analysis of boundary value problems. In contrast, the subloading surface model adopts the associativity leading to the symmetry of the elastoplastic stiffness modulus tensor. 6. The cap model predicts the failure surface which is determined uniquely by the Drucker-Prager yield surface itself, independent of the loading paths, because the interior of the yield surface is assumed to be a purely elastic domain and only the softening is induced when the stress reaches the Drucker-Prager yield surface. However, the surface depicted by connecting the peak stresses depends on the loading paths and its meridian section for the constant Lode angle is not straight but curved in real soils. In contrast, these facts can be described pertinently by the subloading surface model combined with the modified Cam-clay yield surface (cf. Hashiguchi et al. 2002). 7. The cap model is required the tension cut for the Drucker-Prager yield surface, which runs out sharply into the negative pressure range, when it is applied to the description of deformation in vicinity of zero pressure. In contrast, the subloading surface model does not require the tension cut because it adopts the normal-yield surface passing through the vicinity of the null stress state. 8. The cap model is accompanied with the singularity in the direction of outward-normal of the yield surface, i.e. plastic strain rate on the intersecting lines of the Drucker-Prager yield surface with the Cam-clay and the tension-cut yield surfaces. It results in unrealistic description of deformation behavior and would induce the computational difficulty in deformation analysis. In contrast,
13.3
Subloading Surface Model for Soils
389
200 Subloading surface model 150
||σ' || (kPa)
100 Drucker-Prager model (Non-AFR) Drucker-Prager model (AFR) 50
: Test data
0 0
䠉5
䠉10 a (%)
䠉15
䠉20
2.0 Drucker-Prager model (AFR) 1.5 1.0 v
(%)
0.5
Subloading surface model Drucker-Prager model (Non-AFR)
0.0 : Test data
䠉0.5 䠉1.0
0
䠉5
䠉10 a (%)
䠉15
䠉20
Fig. 13.9 Comparison of the calculated results by the Drucker-Prager and the subloading surface models with the test data (after Skempton and Brown 1961) of Weald clay for the drained triaxial compression with the constant lateral pressure
the subloading surface model adopts a single smooth normal-yield surface so that it is not accompanied with the singularity of the plastic modulus. 9. The cap model predicts the simultaneous occurrence of the peak stress and the maximum volume compression in over-consolidated clays and dense sands, in contradiction to experimental facts. In contrast, the subloading surface model provides the realistic prediction that the peak stress and the maximum ratio of volume expansion strain rate vs. axial strain rate occur simultaneously as observed in real soils: Figs. 13.9 and 13.10. 10. The cap model is required to incorporate at least two more material constants describing the inclinations of yield and plastic potential surfaces in addition to the material constants in the Cam-clay model. In contrast, the subloading surface model is required to incorporate only one more material constant u in the evolution rule of the normal-yield ratio despite the distinctively accurate description. In what follows, some comparisons of the simulations of typical triaxial test data by the Cap model and the subloading surface model are shown (Hashiguchi et al. 2002).
390
13 Constitutive Equations of Soils 4000 Subloading surface model
3000 ||σ' || (kPa) 2000
Drucker-Prager model (Non-AFR) Drucker-Prager model (AFR)
1000
: Test data
0 0
䠉5
䠉10
䠉15
䠉20
䠉25
a (%)
3 Drucker-Prager model (AFR)
2 1
v
(%)
0 Subloading surface model Drucker-Prager model (Non-AFR)
䠉1
: Test data 䠉2 0
䠉5
䠉10
䠉15
䠉20
䠉25
a (%)
Fig. 13.10 Comparison of the calculated results by the Drucker-Prager and the subloading surface models with the test data (after Stark et al. 1994) of kaolinite-silt mixtures for the drained triaxial compression with the constant lateral pressure
The simulations of the test data measured by Skempton and Brown (1961) for Weald clay subjected to the drained triaxial compression with a constant lateral stress are shown in Fig. 13.9 where the material constants and the initial value are selected as follows: ~ ~ ¼ 0:002; m ¼ 0:37; Mc ¼ 1:2; ð/c ¼ 36:17 Þ k ¼ 0:045; j My ¼ 0:574; Mp ¼ 0:071 for the Drucker Prager yield surface; u ¼ 33 for the subloading surface model; F0 ¼ 330:0 kPa; whilst the initial stress state is r0 ¼ 67:0I kPa. Here, the Mc -value in Eq. (13.48) is given for the value of M, since the simulation is limited to the triaxial compression state. My and Mp are the inclinations of yield and the plastic potential 0 surface, respectively, of the Drucker-Prager model in the ðp; jjr jjÞ plane. The associated flow rule and the nonassociated flow rule are abbreviated as AFR and Non-AFR, respectively, in this figure. The similar simulations for the test data of kaolinite-silt mixtures measured by Stark et al. (1994) are shown in Fig. 13.10 where the material constants and the initial value are selected as follows:
13.3
Subloading Surface Model for Soils
391
~ ~ ¼ 0:006; m ¼ 0:3; Mc ¼ 1:051; k ¼ 0:1; j My ¼ 0:528; Mp ¼ 0:093 for the Drucker Prager model; u ¼ 35 for the subloading surface model; F0 ¼ 6; 000:0 kPa; whilst the initial stress state is r0 ¼ 1275:0I kPa. The subloading surface model gives rise to the clearly better prediction than the Drucker-Prager model for both the axial stress-axial strain and the volumetric strain-axial strain curves. The curves predicted by the Drucker-Prager model are not smooth, which are formed by the three segments, i.e. the elastic, the elastoplastic and the critical state segments, whilst the former two form the concave curves of the ‘Eiffel-tower’ shape. Intense softening is induced rapidly lowering to the critical state immediately after the stress reaches the Drucker-Prager yield surface. However, note that the adoption of the non-associated flow rule in the Drucker-Prager model does not lead to the substantial improvement in simulation, whilst the subloading surface model adopting the associated flow rule gives the realistic prediction even for the volumetric strain. The parameter u is determined such that the stress–strain curve fit to the gentleness in the elastic–plastic transition. 1200 1000
||σ' || (kPa)
Drucker-Prager model Yield surface in Drucker-Prager model
800 600
Test data p0 (kPa) 104 553
400 200
Subloading surface model 0
0
200
400
600 p (kPa)
800
1000
1200
1200 1000
||σ' || (kPa)
Test data p0 (kPa) 104 553
Drucker-Prager model =0.30 =0.45
800 600 400 200
Subloading surface model 0
0
1
2
3
4
5
a (%)
Fig. 13.11 Comparison of the calculated results by the Drucker-Prager and the subloading surface models with the test data (after Bishop et al. 1965) for the undrained triaxial compression with the constant lateral pressure
392
13 Constitutive Equations of Soils
The simulations of the stress paths and the stress–strain curves to the test data measured by Bishop et al. (1965) for London clay subjected to the undrained triaxial compression are shown in Fig. 13.11 where the material constants and the initial value are selected as follows: ~ ~ ¼ 0:0063; Mc ¼ 0:82; k ¼ 0:022; j m ¼ 0:3 and m ¼ 0:45; My ¼ 0:62; Mp ¼ 0:21 for the Drucker Prager model; m ¼ 0:3; u ¼ 70:0 for the subloading surface model; F0 ¼ 1; 700:0 kPa: The similar simulations for the test data of red clay measured by Wesley (1990) for red clay are shown in Fig. 13.12 where the material constants and the initial value are selected as follows:
400 Drucker-Prager model Yield surface of Drucker-Prager model
300
||σ' || (kPa) 200
Subloading surface model
100
0
Test data p0 (kPa) 50 100 250
100
0
400
200 p (kPa)
300
400
Drucker-Prager model =0.30 =0.43
300
||σ' || (kPa)
Test data 200
p0 (kPa) 50 100 250
100 Subloading surface model 0
0
2
4
6
8
10
a (%)
Fig. 13.12 Comparison of the calculated results by the Drucker-Prager and the subloading surface models with the test data (after Wesley 1990) for the undrained triaxial compression with the constant lateral pressure
13.3
Subloading Surface Model for Soils
393
~ ~ ¼ 0:012; Mc ¼ 1:015; k ¼ 0:035; j m ¼ 0:3 and m ¼ 0:43; My ¼ 0:767; Mp ¼ 0:24 for the Drucker Prager model; m ¼ 0:3; u ¼ 20:0 for the subloading surface model; F0 ¼ 300:0 kPa: The test data are predicted fairly well by the subloading surface model. On the other hand, both the stress paths and the stress–strain curves predicted by the Drucker-Prager model are quite different from the test data, which are not smooth being formed by the three segments, where the Poisson’s ratio is selected two levels of m ¼ 0:30 and 0.45 in Fig. 13.11 and m ¼ 0:30 and 0.43 in Fig. 13.12. The simple (initial) subloading surface model has been widely applied to the analyses of soil deformation behavior (e.g. Hashiguchi and Ueno 1977; Hashiguchi 1978; Topolnicki 1990; Kohgo et al. 1993; Asaoka et al. 1997; Hashiguchi and Chen 1998; Chowdhury et al. 1999; Hashiguchi et al. 2002; Khojastehpor and Hashiguchi 2004a, b; Nakai and Hinokio 2004; Hashiguchi and Mase 2007; Wongsaroj et al. 2007 and many others).
13.4
Extension of Material Functions
Material functions contained in constitutive equation of soils formulated in the last section will be extended in order to describe the deformation behavior more realistically for the negative pressure range and the isotropic and anisotropic hardening behaviors.
13.4.1
Yield Surface with Tensile Strength
Consider the extended yield surface fulfilling the following conditions. (1) It includes not only positive but also negative pressure ranges. Here, note that the subloading surface is indeterminate at the null stress point when the stress reaches the null stress and thus the singular point of plastic modulus is induced since the normal-yield and the subloading surfaces pass through the null stress point which is thus to be the similarity-center of these surfaces in the initial subloading surface model described in Sect. 13.3. On the other hand, this problem is not induced in the extended subloading surface model because the similarity-center of the normal-yield and the subloading surfaces, i.e. elastic-core is not fixed at the null stress point and thus the subloading surface does not pass through the null stress point in general. The exclusion of the singularity of plastic modulus at the null stress point is of importance for the engineering design of soil structures because soils near the side edges of
394
13 Constitutive Equations of Soils
footings, soils at the pointed ends of piles, etc. are exposed to the null or further negative stress state. In addition, the incorporation of tensile yield strength is of importance for the engineering design of structures of natural soils such as soft rocks and cement-treated soils widely used recently, which have the tensile yield strength. Further, it should be noticed that the shift of the yield surface is of the importance also in the computational aspect. (2) The yield surface expands/contracts maintaining a similarity with respect to the origin of stress space so that the yield stress increases/decreases in all directions in the space. (3) For the sake of mathematical simplicity, the yield condition is described by a separate form consisting of the function of the stress, i.e. f ðrÞ and the function including the isotropic hardening variable, i.e. the isotropic hardening function FðHÞ which describes the size of the yield surface. Here, the function f ðrÞ must be a homogeneous function of the stress tensor r in order to fulfill the above-mentioned conditions (2) and (3). Equation (13.41) becomes the following equation through the translation of the yield surface to the negative pressure range by nh F ðp ! p þ nh FÞ (Hashiguchi 2007d; Hashiguchi and Mase 2007).
p ðð1=2Þ nh ÞF F=2
2
2 0 jjr jj þ ¼1 MF=2
Fig. 13.13 Yield surface of soils extended to negative pressure range
ð13:74Þ
13.4
Extension of Material Functions
395
leading to ð13:740 Þ
ð1 nh Þnh F 2 þ ð1 2nh ÞpF ðp2 þ q2 Þ ¼ 0 where 0
q
jjr jj M
ð13:75Þ
nh is the material constant, while it must fulfill nh 1=2 since the tensile yield stress is lower than the compression yield stress and further the inequality nh \pe =F is required, since the volume does not become infinite by the elastic deformation inside the yield surface, i.e. for p [ nh F. The yield surface in Eq. (13.74), i.e. ~ is given from (13.75) is depicted in Fig. 13.13 for the axisymmetric stress state. M 0 e 2nh ÞF=2 as follows: the relation jjr jj ¼ MF=2 ¼ Mð1 e ¼ M
1 M 1 2nh
ð13:76Þ
The frictional angle /c in the triaxial compression is described from Eqs. (13.61) and (13.76) as follows: /c ¼ sin1
3Mc pffiffiffi 2 6 þ Mc
¼ sin1
ec 3ð1 2nh Þ M pffiffiffi ec 2 6 þ ð1 2nh Þ M
! ð13:77Þ
e c are the values of M and M e in the axisymmmetric compression where Mc and M stress state. Equation (13.75) can be expressed in the separated form of the function f ðp; qÞ of the stress and internal variable and the hardening function F, i.e. 8 2 > < p½1 þ ðq=pÞ for nh ¼ 0 f ðp; qÞ ¼ F; f ðp; qÞ ¼ 1 > : ~ ðpq nh pÞ for nh 6¼ 0 nh
ð13:78Þ
where ~ nh 2ð1 nh Þnh ; nh 1 2nh ; pq
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi p2 þ 2~nh q2
ð13:79Þ
396
13 Constitutive Equations of Soils
The partial-derivative of the function @f ðp; qÞ=@r for nh 6¼ 0 is given as follows: @f ðp; qÞ 1 @pq @p nh ¼ ~ @r @r nh @r 2
3 16 1 @p @q @p 7 þ 4~ nh q nh 5 ¼ 4 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2p ~ @r @r @r nh 2 p 2 þ 2 ~ nh q2 1 3 0 2 1 6 1B p 2n~h @q7 C ¼ 4 @qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi nh AI þ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi q 5 ~ 3 @r nh p2 þ 2 ~ nh q 2 p2 þ 2~nh q2
ð13:80Þ
where @p @ 1 1 ¼ trr ¼ I @r @r 3 3
ð13:81Þ
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi p2 þ 2~nh q2
1 @ðp2 þ 2~nh q2 Þ 1 @p @q ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2p þ 4~nh q @r @r @r @r 2 p2 þ 2~nh q2 2 p2 þ 2~nh q2 p ffiffi ffi 1 2 1 0 3 6 1 1 0 0 ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pI þ 4~nh q ðtr þ tr 2 pffiffiffi cos 3htr I 3 M 8 þ cos 3hr 3 6 2 p2 þ 2~nh q2
@pq @ ¼ @r
ð13:82Þ noting pffiffiffi @ cos 3hr 3 6 0 2 1 1 0 ¼ 0 t pffiffiffi cos 3htr I @r jjr jj r 3 6 @M @ 7 7@ cos 3hr =@r Mc ¼ Mc ¼ @r @r 8 þ cos 3hr ð8 þ cos 3hr Þ2 pffiffiffi 0 0 21 6 tr 2 p1ffiffi6 cos 3htr 13 I Mc ¼ ð8 þ cos 3hr Þjjr0 jj 8 þ cos 3hr pffiffiffi 0 0 3 6 tr 2 p1ffiffi6 cos 3htr 13 I M ¼ ð8 þ cos 3hr Þjjr0 jj
ð13:83Þ
ð13:84Þ
13.4
Extension of Material Functions
397
0 @jjr0 jj 0 M jjr jj @M @q @ jjr jj @r ¼ ¼ @r M2 @r @r M 1 0 pffiffiffi 0 2 0 p1ffiffi cos 3ht 1 I 3 6 t r r 3 1 0 6 ¼ 2 @t r M þ MA 8 þ cos 3hr M pffiffiffi 1 0 3 6 1 1 0 0 Þ tr 2 pffiffiffi cos 3htr I ¼ ðtr þ M 8 þ cos 3hr 3 6
ð13:85Þ
In the above, the yield surface of soils is formulated so as to fulfill the conditions (1)–(3) based on the modified Cam-clay model. It is difficult to derive the other yield surface fulfilling the conditions (1)–(3). For instance, consider the translation of the original Cam clay model to the negative pressure range by p ! p þ nh F.
0
jjr jj =M ðp þ nh FÞ exp p þ nh F
¼F
ð13:86Þ
However, a separated form into the function of stress and internal variables and the hardening function cannot be derived from this equation. On the other hand, the translation of the yield surface to the negative pressure range by the constant value Cy ðp ! p þ Cy Þ is adopted for constitutive equations for unsaturated soils (e.g. Alonso et al. 1990; Simo and Meschke 1993; Borja 2004). The modified Cam-clay model, for instance, is described by this translation as follows:
p ð1=2ÞF þ Cy F=2
2
þ
2 0 jjr jj ¼1 MF=2
ð13:87Þ
In this equation, the yield surface expands/contracts from/to the fixed point r ¼ Cy I ðp ¼ Cy Þ on the hydrostatic axis and thus it does not fulfill the condition (2). The incorporation of this yield condition into the subloading surface model leads to the physical impertinence that the unloading without a plastic deformation is induced against the fact that a large plastic deformation (expansion) is induced when the stress translates towards the negative pressure direction from the origin of stress space. The computer program for soil deformation analysis based on the formulations described so far is given in the Appendix L(b).
13.4.2
Rotational Hardening
The inherent anisotropy represented in the orthotropic anisotropy described in Sect. 12.7 cannot be ignored in metals and woods. On the other hand, the induced
398
13 Constitutive Equations of Soils
Fig. 13.14 Inadequacy of kinematic hardening for description of anisotropy of soils: Stress can never return to null state
anisotropy is more dominant in soils since soils are assemblies of particles with weak cohesions between them and thus the rearrangement of soil particles is induced easily. Here, the yield surface of soils must always include the origin of stress space but does very slightly because of the weak cohesion. Besides, the remarkable softening (contraction of the yield surface) by the plastic volume expansion is induced as the stress approaches the null stress state. Then, the stress can never return to the origin of stress space, once the yield surface translates so as not to include the origin, as illustrated on the (p, q) plane for the axisymmetric stress state in Fig. 13.14. Therefore, the kinematic hardening in Eq. (8.89) is not applicable to soils. General speaking, the stress in pressure-independent materials would correspond to the stress ratio, i.e. the ratio of the deviatoric stress versus pressure in pressure-dependent, i.e. frictional materials such as soils and further the translation of yield surface, i.e. the kinematic hardening in the former would correspond to the rotation of yield surface in the latter. The description of anisotropy of soils by the rotation of the Cam-clay yield surface was proposed by Sekiguchi and Ohta (1977), 0 0 replacing the deviatoric stress r to the novel variable r pb ðtr b ¼ 0Þ in the yield condition. This concept was called the rotational hardening in contrast to the kinematic hardening for pressure-independent materials and the second-order deviatoric dimensionless tensor b is referred to as the rotational hardening variable by Hashiguchi (1977). Then, the yield condition in Eq. (13.74) or (13.78) is extended as follows (Hashiguchi and Mase 2007):
p ½ð1=2Þ nh F F=2
2
2
32 _0 jjr jj 5 ¼1 þ4_ M F=2
ð13:88Þ
13.4
Extension of Material Functions
399
Fig. 13.15 Rotated yield surface in the (p, q) plane
i.e.
( _
_
f ðp; qÞ ¼ F; f ðp; qÞ ¼
_
p½1 þ ðq=pÞ2 for nh ¼ 0 1 _ ~n ðpq nh pÞ for nh 6¼ 0
ð13:89Þ
h
where _0
r r0 pb
ð13:90Þ
_0
_
q
jjr jj
ð13:91Þ
_
M qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _2 _ _2 nh q pq p þ 2~
ð13:92Þ
pffiffiffi 14 6 sin /c M ðcos 3h_ Þ ¼ r ð3 sin /c Þð8 þ cos 3h_ Þ _
ð13:93Þ
r
cos 3h_ r
_0 pffiffiffi 0 3 0 r 6trt _ ; t _ _0 r r kr k
ð13:94Þ
The yield surface in Eq. (13.88), i.e. Equation (13.89) is depicted in Fig. 13.15 for the axisymmetric stress state. The evolution rule of rotational hardening tensor b is given below (Hashiguchi and Chen 1998; Hashiguchi 2001a). The following assumptions are adopted for the formulation of the evolution rule. (1) Rotation of the yield surface is induced only by the deviatoric component of the plastic strain rate independent of the mean component. (2) The rotation ceases when the central axis of yield surface reaches the surface, called the rotational limit surface, which exhibits the conical surface having the summit at the origin of stress space. Let the rotational hardening limit surface be given by
400
13 Constitutive Equations of Soils
Normal-yield surface 3
Central axis η= β
Hydrostatic axis η=0
Rotational limit surface ||η|| = M r
2
1
Fig. 13.16 Yield surface and rotational limit surfaces (illustrated in the principal stress space)
kgk ¼ Mr
ð13:95Þ
where Mr is the stress ratio in the rotational hardening limit surface, called the rotational limit stress ratio, and let it be given following Eq. (13.60) by pffiffiffi 14 6 sin /r ð13:96Þ Mr ðcos 3h_ Þ ¼ r ð3 sin /r Þð8 þ cos 3hrÞ _
/r being the material constant, called the rotational limit angle. (3) The central axis of yield surface g ¼ b rotates towards the conjugate line g ¼ Mr t_ on the rotational limit surface, where the conjugate line is the genr
erating line of the rotational limit surface which is observed from the hydrostatic axis in the same direction observed from the central axis g ¼ b of the yield surface to the current stress (see Fig. 13.16). Based on the above-mentioned assumptions, let the following evolution rule of rotational hardening be postulated in the form in Eq. (8.89) with the replacements of a ! b; ck ! br ; bk F ! Mr as follows: 1 p0 p0 b ¼ br e jj e jjb Mr
ð13:97Þ
where br is the material constant. Here, it is noteworthy that the rotational hardening is not induced since the deviatoric strain rate is not induced when the stress
13.4
Extension of Material Functions
401
Fig. 13.17 Relation of axial component of rotational hardening variable versus axial plastic strain in the axisymmetric stress state
lies on the central axis of the subloading surface in the modified Cam-clay model, p0
i.e. e ¼ O ! b ¼ O for g ¼ b as illustrated in Fig. 13.15. Here, needless to say, the deviatoric plastic strain rate is adopted in Eq. (13.97) since the anisotropic hardening is independent of plastic volumetric strain rate. Equation (13.97) is described in one-dimensional state as follows:
ba ¼ br
ba p 1 ea Mr
ð13:98Þ
which is shown in Fig. 13.17.
13.5
Extended Subloading Surface Model
Elastoplastic constitutive equation for describing cyclic loading behavior of soils is formulated below by incorporating the rotational hardening instead of the kinematic hardening into the extended subloading surface model described in Chap. 11. Further, the supertield surface proposed by Asaoka et al. (2000) is incorporated in order to describe the degradation of soil skeleton structure by which the deformation behavior of soils with different void ratios may be described uniformly.
13.5.1
Superyield, Normal-Yield and Subloading Surfaces
Setting a ¼ O in Eqs. (11.1) and (11.2) and incorporating the rotational hardening b, the normal-yield and the subloading surfaces for soils are given as follows (see Fig. 13.18):
402
13 Constitutive Equations of Soils
Fig. 13.18 Rotated normal-yield, subloading, elastic-core and limit elastic-core surfaces shown in the (p, q) plane
f ðr; bÞ ¼ FðHÞ
ð13:99Þ
f ðr; bÞ ¼ RFðHÞ
ð13:100Þ
where the following relation holds by virtue of the similarity of the subloading surface to the normal-yield surface. _
r r að¼ Rry Þ ¼ r þ Rc
ð13:101Þ
a ð1 RÞc ðac ¼ Rða cÞ; a ¼ 0Þ
ð13:102Þ
with
leading to
a ¼ ð1 RÞ c R c
ð13:103Þ
where _
r rc
ð13:104Þ
a is the conjugate (similar) point in the subloading surface to the reference point að¼ 0Þ in the normal-yield surface as shown in the ðp; qÞ plane in Fig. 13.18. ry is the conjugate point on the normal-yield surface to the current r on the subloading surface.
13.5
Extended Subloading Surface Model
403
Now, introduce the superyield surface (Asaoka et al. 2000) (Fig. 13.19(a)): f ðr; bÞ ¼ RFðHÞ
ð13:105Þ
where Rð 1Þ is the super-normal-yield ratio designating the ratio of the size of the superyield surface to the size of the normal-yield surface, which represents the degree of the soil skeleton structure (larger in loose soils). The subloading surface in Eq. (13.100) is described by f ðr; bÞ ¼ RRFðHÞ
ð13:106Þ
where Rð R 1Þ is the super-normal-yield ratio designating the ratio of the size of the subloading surface to the size of the superyield surface. The subloading surface model combined with the superyield surface is called the superyield-subloading surface model, abbreviated as SYS model, by Asaoka et al. (2000, 2002) which provides the primitive formulation: (1) The extended subloading surface model is not incorporated but the initial subloading surface model is incorporated, (2) The yield surface is limited to the positive pressure range and obeys the irrational rotational (rate-linear) hardening rule and (3) The isotropic hardening function is formulated based on the e ln p linear relation limited to the infinitesimal deformation as was described in Sect. 13.1(b). Let the evolution rules of the super-normal yield ratio R based on the assumption that the super-yield surface shrinks by the deviatoric plastic strain rate and shrinks/ expands by the plastic volumetric contraction/expansion and the super-yield ratio R based on the assumption that the subloading surface approaches the super-yield surface in the plastic loading process be given by ð13:107Þ where ð13:108Þ cv ð 1Þ; u; a and u are the material constants and u is the material function which will be formulated in Eq. (13.130). The material constant cv would be larger in clays than sands. The following relation for the normal-yield ratio R in terms of the super-normal yield ratio R and the super-yield ratio R holds from Eq. (13.107) with Eq. (13.108) noting Eq. (13.106).
ð13:109Þ
404
13 Constitutive Equations of Soils
q
Subloading surface f (σ, β) = RF ( H )
Normal-yield surface f (σ, β) = F ( H )
superyield surface f (σ, β) = RF ( H )
σ
p
0
(a) Various basic surfaces in the (p, q) plane.
q
q
Stress path
Stress path
p
0 Loose sands
0
Dense sands
p
(b) Stress path under constant volume or undrained condition.
Fig. 13.19 Various basic surfaces and undrained stress path based on super-yield subloading surface model proposed by Asaoka et al. (2000)
The undrained stress paths in loose and dense sands can be predicted by the SYS model with one same set of material parameters as shown in Fig. 13.19(b) without resorting to the plastic deviatoric hardening (Nova 1977; Wilde 1977) as will be known from the plastic modulus shown in Eq. (13.129). The material-time derivative of Eq. (13.100) reads: @f ðr; bÞ @f ðr; bÞ @f ðr; bÞ :r :a þ :b ¼ RF þRF @r @r @b
ð13:110Þ
which can be rewritten as "
# 1 @f ðr; bÞ F R :b r ¼ 0 n : r n : a þ r þ r F R RF @b
ð13:111Þ
13.5
Extended Subloading Surface Model
405
where n
@f ðr; bÞ @r
@f ðr; bÞ @r ðjjnjj ¼ 1Þ
ð13:112Þ
noting the following equation based on the Euler’s theorem for the function of r in homogeneous degree-one. 1 ¼ @f ðr; bÞ @r
13.5.2
@f ðr; bÞ :r f ðr; bÞ RF @r n¼ n¼ n n:r n:r n:r
ð13:113Þ
Evolution Rules of Internal Variables
The evolution rules of the internal variables are formulated in the following. (a) Rotational hardening variable The rate of rotational hardening is given by extending Eq. (13.97) as follows: 0 1 p0 1 p0 ð13:114Þ b ¼ br @e a e bA ¼ f bn k Mr where
pffiffiffi 14 6 sin /r M r ðcos 3hrÞ ¼ ð3 sin /r Þð8 þ cos 3hrÞ _
_
_
_0
_0 0
0
r r pb; tr _
r
_0
jjr jj 0
; cos 3hr
f bn br @n0
_
pffiffiffi 0 3 6tr tr _
ð13:115Þ
ð13:116Þ
1 1 a
kn0 kbA
ð13:117Þ
Mr /r is the angle designating the limit of the rotation of the normal-yield surface. (b) Elastic-core To avoid the unlimited approach of the elastic-core to the normal-yield surface, first let the following surface, called the elastic-core surface, be introduced as shown in Fig. 13.18, which passes through the elastic-core c and possesses a similar shape and orientation to the normal-yield surface with respect to the null stress ðr ¼ a ¼ 0Þ
406
13 Constitutive Equations of Soils
ð13:118Þ . where the variable is the ratio of the size of the elastic-core surface to that of the normal-yield surface, called the elastic-core yield ratio. It plays the role of a measure for the approaching degree of the elastic-core to the normal-yield surface. Since the elastic-core must lie inside the normal-yield surface as described above, the elastic-core yield ratio has to be less than unity. Then, the inequality ð13:119Þ must hold, where vð\1Þ is a material constant exhibiting the maximum value of > > > < p4 1 þ
_!2
q p
3 5
> > 1 _ > > : ð p q nh pÞ ~ nh
for nh ¼ 0
ð13:132Þ
for nh 6¼ 0
where p ðtr rÞ=3 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _2 _ pq p2 þ 2~ nh q
ð13:133Þ ð13:134Þ
_0
_
q
jjr jj
ð13:135Þ
_
M _
M ðcos 3hrÞ ¼ _
pffiffiffi 14 6 sin /c ð3 sin /c Þð8 þ cos 3hrÞ _
ð13:136Þ
_0
pffiffiffi r cos 3hr 6tr tr0 3 ; tr0 _0 ðjjt0_ jj ¼ 1Þ r jjr jj _
13.5.5
_
_
ð13:137Þ
Partial Derivatives of Subloading Surface Function
The partial derivatives of the function in Eq. (13.132) are shown below. 8 _ _2 > > 1 þ ðq=pÞ2 þ pð2p33 Þq for nh ¼ 0 > _ 1 @f ðp; qÞ < 0 ¼ 1 1 2p > @p @ > nh A for nh 6¼ 0 > _ :~ nh 2 pq 8 !2 _ ð13:138Þ > > q > > 1 for n ¼ 0 > h > < p 0 1 ¼ > > 1 p > > @ nh A for nh 6¼ 0 > > _ :~ nh p q
410
13 Constitutive Equations of Soils
_
@f ðp; qÞ _
@q
¼
8 _ > q > > >2 < p
for
nh ¼ 0 ¼
_
> 1 1 4~ nh q > > > _ :~ 2 n pq
nh 6¼ 0
for
8 _ > q > > >2 < p
for
nh ¼ 0 ð13:139Þ
_
> q > > > :2_ pq
for
nh 6¼ 0
@p 1 ¼ I @r 3
ð13:140Þ
@r0 0 ¼I @r
ð13:141Þ
0
@rij @ðrij þ pdij Þ 1 1 ¼ ¼ ðdik djl þ dil djk Þ dij dkl @rkl @rkl 2 3
!
_0
@r 1 0 ¼I þ bI @r 3
0
ð13:142Þ
1 _0 0 @rij @ðrij pbij Þ 1 1 1 @ ¼ ¼ ðdik djl þ dil djk Þ dij dkl þ bij dkl A @rkl 2 3 3 @rkl pffiffiffi 14 6 sin /c
_
@M _
@ cos 3hr
¼
_
ð3 sin /c Þð8 þ cos 3hrÞ2 _
3 sin /c _ 2 pffiffiffi ¼ ¼ M _ 14 6 sin /c 8 þ cos 3hr M
@tr0 _
_
¼
@ r0 0
0 _
B _0 B B @ t ij B 0 ¼ B _ B@r @ kl
¼
0 _
r ij ffi @ qffiffiffiffiffiffiffiffiffi 0 0
@ r ij 0 _
_ _
r rs r sr
_
0
¼
@ r kl
1 2 ðdik djl
_
jjrjj
0
0
_
0
0
0
ðI tr tr Þ _
qffiffiffiffiffiffiffiffiffiffiffi ffi 0 0 0 _ _ _ @ rrs rsr rij _
@rkl
1
ð13:143Þ
!
ð13:144Þ
_
qffiffiffiffiffiffiffiffiffi ffi 0 0 _ _
r rs r sr 0 _
@ r kl
rrs rsr
qffiffiffiffiffiffiffiffiffiffiffi ffi 0 0 _ _ ^0ij t0 kl þ dil djk Þ rrs rsr r _ _
0
_
0
rrs rsr
r
1
C 1 1 0 0 ffi ¼ qffiffiffiffiffiffiffiffiffiffiffi ðdik djl þ dil djk Þ tr ij tr kl C 0 0 A _ _ 2 rrs rsr _
_
13.5
Extended Subloading Surface Model
411
@ðtrt0 r3 Þ ¼ 3t0 r2 @t0 r _
ð13:145Þ
_
_
0
1 0
0
B@tr rs tr st tr @ 0 _ @ t ij _
_
_
0
tr
0
¼ dir djs tr st tr _
_
0
_
@ cos 3hr _
0
¼
tr
_
3 _
_
0
tr
0 0 0 0 C þ tr rs tr st dit djr ¼ 3trir trrj A _
_
_
pffiffiffi 0 _ 0 ð 6tr 2 cos 3hrtr Þ _
0
jjr jj
@r
0
þ tr rs dis djt tr
_
ð13:146Þ
_
0 pffiffiffi 1 _ _ 0_ 0_ 0_ 0_ 0_ 0_ pffiffiffi @t0 r pffiffiffi 0 _ 0 _ @t0 r r r r r r r @ 6 t t t t t @t lm mn nl lm mn nl rs rs @ sm t r nr _0 A ¼ 6 ¼ 3 6t r _ _0 _0 rs @t0 r @ rij @ rij @ rij
pffiffiffi 0 _ 0 _ 1 1 0_ 0_ sn t r nr qffiffiffiffiffiffiffiffiffiffiffiffiffi rs t r ij dri dsj þ drj dsi t r ¼ 3 6t r _ _ 2 pq r qp r _ _ pffiffiffi _ _ _ _ 1 in t0 r nj t0 r sn t0 r nr t0 r rs t0 r ij ¼ 3 6 qffiffiffiffiffiffiffiffiffiffiffiffiffi t0 r 0 0 _ _ qp pq r r pffiffiffi _ _ _ 3 0 0 0 t cos 3h r t ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r r r 6 t in nj ij pq t0 r qp t0 r _
@q
_0
@r
0
¼ _ t0_ þ 1
r
M
_
_
@ @q ¼ q@q ffiffiffiffiffiffiffiffiffiffiffiffiffi _0 _0 _0 @rij @ rpq rqp
3 8 þ cos 3h_
pffiffiffi 0 ð 6t0_ 2 cos 3h_ t_ Þ
qffiffiffiffiffiffiffiffiffiffiffiffiffi _0 _0 @ rpq rqp _0
_
þ
@q _
ð13:147Þ
r r
r
r
_
@M @ cos 3h_
@ cos 3h_ _0
r
@rij @rij @M r qffiffiffiffiffiffiffiffiffiffiffiffiffi 0 0 _ _ _ @ rpq rqp @M @ cos 3h_r rij 1 q ffiffiffiffiffiffiffiffiffiffiffiffiffi ¼ _ _0 _2 _0 _0 @ cos 3h_ @r M @ rpq rqp ij M r 2 31 pffiffiffi 0 0 14 3 ¼ _ t_ ij þ ð 6t_ ir t_ rj cos 3h_ t_0 ij Þ5A 8 þ cos 3h_ r r r r r M r ( " _ #) _ _ _ _ @f ðp; qÞ 1 @f ðp; qÞ @f ðp; qÞ @q 1 @q Iþ tr _0 ðI bÞ I ¼ _ _0 @r 3 @p 3 @q @r @r _0
412
13 Constitutive Equations of Soils _
_
_
_
_
0
@f ðp; qÞ 1 @f ðp; qÞ @p @f ðp; qÞ @q @rrs ¼ þ 0 _ _ @rij 3 @p @rij @q @rrs @rij _ _ _ 1 @f ðp; qÞ @p @f ðp; qÞ @q 1 1 1 ðdri dsj þ drj dsi Þ drs dij þ brs dij ¼ þ 0 _ _ 3 @p @rij 3 3 @q @rrs 2 _ _ _ 1 @f ðp; qÞ @f ðp; qÞ @q 1 1 1 dij þ ¼ d þ d d Þ d þ d ðd d b ri sj rj si rs ij ij 0 _ _ 3 @p 3 3 rs @q @rrs 2 8 2 3 " _ #) _!2 _( _ > > 1 q q @ q 1 @ q > > 5I þ 2 > 41 0 :ðI bÞ I > _0 > p p 3 3 _ > > _ @ r @ r > @f ðp; q Þ < for nh ¼ 0 0 1 ¼ " _ #) _ ( _ > @r > q @q 1 @q > 11@p > A > 0 :ðI bÞ I _ nh I þ 2 _ > _0 > 3~ 3 _ nh p q > p > @ r @ r q > : for nh 6¼ 0 ð13:148Þ _
_
_
@f ðp; qÞ @f ðp; qÞ @q ¼ p _ _0 @b @q @ r _
_
_
_
0
_
ð13:149Þ
_
@f ðp; qÞ @f ðp; qÞ @q @rrs @f ðp; qÞ @q ¼ ¼ 0 0 ðpI rsij Þ _ _ _ _ @bij @q @rrs @bij @v @rrs _
¼ p
_
_
_
1
@f ðp; qÞ @q 1 @f ðp; qÞ @q A ðdrj dsj þ drj dsi Þ ¼ p 0 0 _ _ _ _ 2 @q @rrs @q @rrs
8 _ > _ @q > > > _ > 2q _0 for nh ¼ 0 @f ðp; q Þ < @r _ ¼ > @b > p _ @q > > > : 2 _ q _0 for nh 6¼ 0 pq @ r 8 _ < pc f1 þ ðqc =pc Þ2 g for nh ¼ 0 _ 1 _ f ðc; bÞ ¼ f ðpc ; qc Þ ¼ : ~ ðpqc nh pc Þ for nh 6¼ 0 nh
ð13:150Þ
ð13:151Þ
13.5
Extended Subloading Surface Model
413
where 1 pc tr c; 3 _
0
0
c c þ pc I
ð13:152Þ
0
c c pc b
ð13:153Þ
_0
_
qc _
M c ðcos 3h_ Þ ¼ c
jjc jj
ð13:154Þ
_
Mc
pffiffiffi 14 6 sin /c ð3 sin /c Þð8 þ cos 3h_ Þ
ð13:155Þ
c
0
_ pffiffiffi 0 3 c 0 6trt _ ; t_ _0 c c c jjc jj qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _2 _ nh qc pqc p2c þ 2~
cos 3h_
"
_
@ qc _0
@c
¼
1 _
M _c
pffiffiffi 3 0 t0_c þ 6t02 _ cos 3h_ t_ c c c 8 þ cos 3h_c
8 _ > _ @ qc > > 2q > c _0 for nh ¼ 0 @f ðc; bÞ < @c _ ¼ > pc _ @ p c @b > > p c _0 for nh 6¼ 0 > : 2 _ p qc @c @f ðc; bÞ 11 ¼ @c 3 ~nh
13.5.6
! p _
p qc
_
qc
nh I þ 2 _ p qc
(
ð13:156Þ
ð13:157Þ # ð13:158Þ
ð13:159Þ
" _ #) 1 @ qc ð13:160Þ 0 :ðI bÞ I _0 3 @_ @c c _
@ qc
Calculation of Normal-Yield Ratio
The normal-yield ratio R must be calculated from the equation of the subloading surface in the unloading process. It can be calculated directly from R ¼ f ðr; bÞ=F by solving f ðr; bÞ ¼ RF in the initial subloading surface model. In the extended subloading surface model, however, it is difficult to solve Eq. (13.100) with Eq. (13.132) for R analytically. Then, we have to calculate R by the semi-analytical
414
13 Constitutive Equations of Soils
or numerical methods. The super-normal yield ratio R does not change and thus the
superyield ratio R can be calculated by R ¼ R=R in the unloading process ðe p ¼ OÞ (a) Semi-analytical method In general, R is calculated numerically by solving the nonlinear equation obtained by substituting the current known values of r; b and c into Eq. (13.100) with _ Eq. (13.132), noting r ¼ r þ Rc in Eq. (11.6). The Newton–Raphson method would be useful for the calculation. The other numerical method is shown here. First, one has 1 _ _0 _ p ¼ trðr þ RcÞ ¼ ðrm þ Rcm Þ; r0 ¼ r þ Rc0 3
ð13:161Þ
from Eq. (11.6). Substituting Eq. (13.161) into Eq. (13.135), one has _0
_0
r ð¼ r0 pbÞ ¼ r þ Rc0 þ
1 _ _0 _ ½trðr þ RcÞ b ¼ r þ Rc0 þ ðrm þ Rcm Þb ð13:162Þ 3
_0
_
q
jjr jj _
M
_0
¼
jjr þ Rc0 þ ðrm þ Rcm Þbjj _
_
ð13:163Þ
M
Further, substituting Eq. (13.161)–(13.163) into Eq. (13.100) with Eq. (13.132) of the extended subloading surface, one has the following equation and can transform it in turn. 2
12 3 0 _0 _ r þ Rc0 þ rm þ Rcm b 6 B C 7 _ B C 7 6 6 _ B C 7 M 6 ¼ RF for nh ¼ 0 rm þ Rcm 61 þ B C 7 _ B C 7 6 7 r þ Rc m m @ A 4 5
9 > > > > > > > > > > > > > > > > =
9> 8vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi > 32ffi u 2 > > > _0 > _ > > u > 0 = < > r þ Rc þ r þ Rc b m m 2 > 1 u _ _ > t rm þ Rcm þ 2~ 4 5 > þ nh rm þ Rcm nh > _ > ~nh > > > > > ;> : M > > > ; ¼ RF or nh 6¼ 0 ð13:164Þ
13.5
Extended Subloading Surface Model
415
9 32 _0 0 _ > > jjðr þ Rc Þ þ ðr þ Rc Þbjj _ m m > 5 ¼ ðrm þ Rcm ÞRF > ð~ rm þ Rcm Þ 4 > > ^ > > M > > > = ¼ 0 for n h vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi 2 3 u 2 0 _ 0 u > _ > u _ > > ~h 4jjr þ Rc þ ðrm þ Rcm Þbjj5 > t½ðrm þ Rcm Þ 2 þ 2n > _ > > > M > > ; _ ¼~ nh RF nh ðrm þ Rcm Þ for nh 6¼ 0 2
2
_
2
_
_0
2
0 2
_
_0
0
M ðrm þ Rcm Þ þ jjr þ Rc jj þ 2ðrm þ Rcm Þ½ðr þ Rc Þ: b _
_
2
_
þ ðrm þ Rcm Þ2 jjbjj2 þ M ðrm þ Rcm ÞRF ¼ 0 for nh ¼ 0
9 > > > > > > =
> _0 _0 _ _ > > M ðrm þ Rcm Þ2 þ 2~ nh jjr þ Rc0 jj2 þ 4~ nh ðrm þ Rem Þ½ðr þ Rc0 Þ: b > > > _ ; _ _ 2 2 2 ~ m þ Rcm Þ jjbjj fM ½~ þ 2nðr nh RF nh ðrm þ Rcm Þ g ¼ 0 for nh 6¼ 0 _2
9 _0 2 _0 0 0 2 2> > M rm þ 2 M rm cm R þ M þ r þ 2 r : c R þ kc k R > > > _0 > > _ _ > 2 0 > þ 2 rm þ Rcm r : b þ 2 rm R þ cm R ðc : bÞ > > > > 2 > _2 > _ _ _ 2 > 2 2 2 > þ rm þ 2Rrm cm þ cm R kbk þ M rmFR þ cm FR ¼ 0 > > > > > for nh 6¼ 0 > = _2 _2 _2 2 0 _ _2 _ ~h r M rm þ 2 M rm cm R þ M c2m R2 þ 2n > > _0 _0 > > _ 2 0 0 2 > > þ 4~ n r : c R þ 2~ nh rm þ cm R r : b nh kc k R þ 4 ~ > > 2 > _ _ _ 2 > 2 0 2 2 ~ ~ > þ 4n rm R þ cm R ðc : bÞ þ 2nh rm þ 2rm cm R þ cm R kbk > > > > > 2 > _2 > 2 2 _ _ 2 ~ ~ > M nh F nh cm R 2nh rm nh F nh cm R þ nh rm ¼ 0 > > > > ; for nh 6¼ 0 0h i2 h i2 1 _ _ ~ ~ ¼ nh F ncm R nh rm C B nh RF nh rm þ Rcm 2 A @ _ _2 nh F nh cm R þ nrm ¼ ~ nh F nh cm R2 2nh rm ~ _
2
_2
_
2
_
_
2
c2m R2
416
13 Constitutive Equations of Soils
_
M
2
0 2 2
0
þ kc k R þ 2cm ðc : bÞR þ c2m kbk2 R2 þ _0 _0 _2 _ þ 2 M rm cm R þ 2 r : c0 R þ 2cm r : b R c2m R2
_
2
2
_
2
M cm FR
2
þ 2rm ðc0 : bÞR þ M rm FR þ 2rm cm kbk2 R _0 2 _0 _2 _2 _ _2 þ M rm þ r þ 2rm r : b þ rm kbk2 ¼ 0 for nh ¼ 0 _
_
_
_
2
9 > > > > > > > > > > > > > > > > > > > > > > > > > =
M c2m R2 þ 2~ nh kc0 k2 R2 þ 4~ nh cm ðc0 : bÞR2 þ 2~ nh c2m kbk2 R2
> > > > > 2 > 2 ~ 0 ~ > M R ðnh F nh cm Þ þ 2M rm cm R þ 4nh ðr : c ÞR > > > _0 _ _ 2 > 0 ~ ~ ~ > þ 4nh cm r : b R þ 4nh rm ðc : bÞR þ 4nh rm cm kbk R > > > > > _2 _2 2 0 0 > _ _ _ _2 _ > þ 2 M nh rm ~ nh F nh cm R þ M rm þ 2~nh r þ 4~nh rm r : b > > > > > 2 > _ ; 2 2 2 _ _ 2 ~ þ 2nh rm kbk M nh rm ¼ 0 for nh 6¼ 0 _
2
_
2
_
_0
9 _2 _2 > 2 > 2 0 2 0 2 2 > M cm þ kc k þ 2cm ðc : bÞ þ cm kbk þ M cm F R > > > > _2 > _0 _0 > > _ _ 0 0 > > þ 2 M rm cm þ 2 r : c þ 2cm r : b þ 2rm ðc : bÞ > > > > > _2 _2 > 2 0 > _ _ _2 2 > > ~m F R þ M rm þ r þ 2rm cm kbk þ M r > > > _0 > > _ _2 2 = þ 2rm r : b þ rm kbk ¼ 0 for nh ¼ 0 _2 2 > _ 2 > > R2 > M c2m þ 2~ nh k c 0 k 2 þ 4 ~ nh c m ð c 0 : b Þ þ 2 ~ nh c2m kbk2 M ~nh F nh cm > > > h _0 _0 > > 2_ _ > 0 0 ~ ~ ~ > þ 2 M rm cm þ 2nh r : c þ 2nh cm r : b þ 2nh rm ðc : bÞ > > > > > _2 _ 2 > 2 _2 _ _ > 2 ~ ~ > þ 2nh rm cm kbk þ M nh rm nh F nh cm R þ M 1 nh rm > > > > _0 2 > _0 > _ _2 > 2 ~ ~ ~ ; þ 2nh r þ 4nh rm r : b þ 2nh rm kbk ¼ 0 for nh 6¼ 0 Solving this quadratic equation, the normal-yield ratio R is expressed as follows: 8 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi B2 AC B > > < for nh ¼ 0 A R ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ~2 ~ ~ ~ > > : B AC B for nh ¼ 6 0 ~ A
ð13:165Þ
13.5
Extended Subloading Surface Model
417
where A
_
2
_2 þ kc k þ 2cm ðc : bÞ þ c2m kbk2 þ M cm F _0 _0 _ _ rm cm þ 2tr r c0 þ 2cm r : b þ 2rm ðc0 : bÞ
M c2m _2
B 2M
0 2
0
_
_
2
_
þ 2rm cm kbk2 þ M rm F _0 2 _0 _2 _2 _ _2 C M rm þ r þ 2rm r : b þ rm kbk2 _
2
9 > > > > > > > > > > > > > > > > > > > > > > > > > > > =
~ M c2 þ 2~ A nh k c 0 k 2 þ 4 ~ n c ð c 0 : bÞ þ 2 ~ nh c2m kbk2 m > 2 h m _ > 2 > > > M ~ nh F nh e m > > > > _0 _0 _2 > > _ _ 0 0 > ~ ~ ~ ~ B M rm cm þ 2nh r : c þ 2nh cm r : b þ 2nh rm ðc : bÞ > > > > > _2 > _ _ > 2 ~ ~ > þ 2nh rm em kbk þ M nh rm nh F nh cm > > > > > _ 2 2 > 0 0 _ _ 2 _2 _ _2 > 2 ~ ~ ~ M 1 n r þ 2~ ; n r þ 4 n r r : b þ 2 n r b C k k h h m h m h m
ð13:166Þ
Explicit numerical calculation processes: (1) First step (beginning of calculation): Calculate the normal-yield ratio R by Eq. (13.165), substituting the trial value _ _ pffiffiffi pffiffiffi M ¼ 2 6 sin /c =3 which is the average of M ¼ 2 6 sin /c =ð3 þ sin /c Þ and pffiffiffi 2 6 sin /c =ð3 sin /c Þ. (2) Second step: Recalculate R by substituting the value 0
1 pffiffiffi pffiffiffi 14 6 sin / 14 6 sin /c c A¼ pffiffiffi M ðcos 3h_ Þ @¼ ð3 sin /c Þð8 þ cos 3h_ Þ r ð3 sin /c Þð8 þ 6trt3 Þ _
r
_
r
ð13:167Þ into Eq. (13.165), while the value of R obtained in the former step is used in Eq. (13.167). (3) Repeat the process (2) until R will reaches the convergence within a prescribed tolerance. (b) Numerical method In what follows, the numerical calculation of R by the Newton–Raphson method is described below.
418
13 Constitutive Equations of Soils
Consider the following equation of R given from Eq. (13.164)2 for the subloading surface. 3 2vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi !2 u u 2 16 rc _ _ 7 nh _ þ nh rm þ Rcm 5 RF ¼ 0 gðRÞ ¼ 4t rm þ Rcm þ 2~ ~ nh M ð13:168Þ where
_ 0 _ rc r þ Rc0 þ rm þ Rcm b
ð13:169Þ
Differentiation of Eq. (13.168) leads to 2 _3 !2 31=2 2 _ @rc 2 @M M r @gðRÞ 1 4 _ r r c @R _ c ~ c @R 42 r 5 rm þ Rcm þ 2~ nh _ 5 ¼ m þ Rcm cm þ 4nh _ _2 @R 2~nh M M M nh þ cm F n~h
ð13:170Þ
0
g ðRÞ ¼
1 ~nh
2 _3 " 2 #1=2 @rc rc @ M _ 2 @R _ @R _ rc M ~ ~ 4 r m þ Rcm cm þ 2nh rc _ 5 þ r m þ Rcm þ 2nh _ 2 M
M
3 nh c ~ nh m
F5
ð13:171Þ where h_0 _ i @rc 0 ¼ v1 r þ Rc þ r þ Rc b : ð c 0 þ cm bÞ m m c @R @ k Tk T @T ¼ : @R kTk @R 2 _ @M 3 sin /c _ 2 4 ¼ pffiffiffi M @R 14 6 sin /c
3
kr k _0
3 pffiffiffi 6t2_ cos 3h_ t_ 5 : ðc0 btrcÞ ð13:173Þ r r
r
noting _
ð13:172Þ
_0
_
@M @M @ cos 3hr @ r ¼ : _0 @R @ cos 3hr @R @r _
_
13.5
Extended Subloading Surface Model _0
@r ¼ @R
419
h_ i _ @ ðr þ RcÞ0 trðr þ RcÞb @R
¼ c0 b tr c
ð13:174Þ
Here, noting g0 ðRn Þ ¼
gðRn þ 1 Þ gðRn Þ 0 gðRn Þ ¼ Rn þ 1 Rn Rn þ 1 Rn
we have Rn þ 1 ¼ Rn
gðRn Þ g0 ðRn Þ
ð13:175Þ
from which R is obtained by repeating calculation until Eq. (13.168) converges within a given tolerance. In addition, we adopt the following “reduced” Newton– Raphson iteration instead of Eq. (13.175) if the convergence is not obtained. Rn þ 1 ¼ Rn ln
gðRn Þ g0 ðRn Þ
ð13:176Þ
Calculate first by setting ln ¼ 1. If the inequality nþ1 ln n gðR gðR Þ Þ \ 1 4
ð13:177Þ
is fulfilled, continue the calculation. However, if Eq. (13.177) is not fulfilled, calculate by reducing ln to the half and repeat again the calculation by reducing ln to the further half until Eq. (13.177) is fulfilled. The constitutive equation of soils based on the extended subloading surface model is described in detail so far. The tangential-inelastic strain rate has to be incorporated for the analyses of non-proportional loading behavior as shown in the analyses by Hashiguchi and Tsutsumi (2001, 2003, 2006) and Khojastehpour and Hashiguchi (2004a, b).
13.6
Simulations of Test Results
Some simulations of test data are shown below in order to show the capability of the subloading surface model to reproduce the real deformation behavior of soils (Hashiguchi and Chen 1998). The simulation of the test data (after Saada and Bianchini 1989) for Hostun sand subjected to the drained triaxial compression with a constant lateral stress, which
420
13 Constitutive Equations of Soils
Fig. 13.20 Drained behavior of Hostun sand (data from Saada and Bianchini 1989)
includes the unloading–reloading process, is shown in Fig. 13.20 where the material constants and the initial values are selected as follows: Material constants: Yield surface ðellipsoid:Þ/c ¼ 27 ; ( 8 > ~ ¼ 0:003; # ¼ 0:025; k ¼ 0:008; j < isotropic volumetric: ~ Hardening=softening deviatoric: l ¼ 0:6; / d d ¼ 25 ; > : rotational: br ¼ 10; /r ¼ 20 ; Evolution of normal yield ratio: u1 ¼ 1:5; m1 ¼ 3:8; Translationon of elastic core : ce = 0.05, Elastic shear modulus: G = 200 000 kPa,
13.6
Simulations of Test Results
421
Initial values: Hardening function: F0 ¼ 400 kPa; Rotational hardening variable: b0 ¼ O; Elastic core: c0 ¼ 50I kPa; Stress: r0 ¼ 100I kPa where the hyperbolic equation UðRÞ ¼ u1 ð1=Rm1 1Þ ðu1 ; m1 : material constantsÞ is used for the evolution rule of the normal-yield ratio. Measured and calculated results are shown by the dashed and solid line, respectively. The simulation of the test data (after Saada and Bianchini 1989) for Hostun sand subjected to the drained proportional loading with b ð¼ ðr2 r3 Þ=ðr1 r3 ÞÞ ¼ 0:666 ðhr ¼ 19 090 Þ from r0 ¼ 500I kPa by the true triaxial test apparatus is shown in Fig. 13.21. The material parameters are the same as those for the above-mentioned Fig. 13.21 Drained proportional loading behavior of Hostun sand (data from Saada and Bianchini 1989). Measured and calculated results are shown by the dashed and solid line, respectively
422
13 Constitutive Equations of Soils
drained triaxial compression, while the sample was preliminarily loaded the isotropic compression from r ¼ 100I kPa to 500I kPa before the test. The simulations of the test data (after Castro 1969) for Banding sand subjected to the undrained triaxial compression with a constant lateral stress are shown in Fig. 13.22 where the material constants and the initial values are selected as follows:
Fig. 13.22 Undrained behavior of Banding sand (data from Castro 1969). Calculated results are shown by the solid lines
13.6
Simulations of Test Results
423
Material constants: Yield surfaceðellipsoid:Þ/c ¼ 26; 30; 31; 32 ; 8 8 8 ~ > > > > < k ¼ 0:025; 0:018; 0:014; 0:010; > > > > > > > > ~ ¼ 0:0067; 0:0065; 0:0060; 0:0058; > < volumetric> j > > : < isotropic # ¼ 0; 0:021; 0:058; 0:138; > Hardening=softening
> > > > l ¼ 1:00; 0:65; 0:30; 0:10; > > > > deviatoric d > : > > /d ¼ 40; 33; 30; 20 ; > > : rotational: br ¼ 10; /r ¼ 20 ;
u1 ¼ 0:1; 0:3; 0:5; 1:0; 33:0; Evolution of normal yield ratio m1 ¼ 0:1; 0:4; 0:5; 0; 7; Translationon of elastic core : ce = 0.049, 0.038, 0.027, 0.014;
Elastic shear modulus: G = 18 000, 23 000, 25 000, 35 000 kPa, Initial values: Hardening function: F0 = 410, 480, 520, 580 kPa; Rotational hardening variable : b0 = O, Elastic core : c0 = - 200, - 110, - 100, - 80 I kPa, Stress: r0 = 67:0I kPa where the four values for each of the above parameters correspond to the initial relative densities Dr ¼ 0:29; 0:44; 0:47; 0:64, respectively, in this order. The different values of material parameters depending on the initial relative densities are used in this simulation because the deviatoric strain hardening (Nova 1979; Wilde 1979) in addition to the plastic volumetric strain hardening is adopted. The simulation would be performed using the unique set of material parameters by the SYS model as described in Sect. 13.5.1, while it was shown explicitly by Asaoka et al (2002) whose formulation is the immature based on the initial subloading surface model. The simulation of the cyclic mobility under the liquefaction to the test data by the subloading surface model was shown by Hashiguchi et al. (2021) although not illustrated here.
13.7
Numerical Analysis of Footing Settlement Problem
Numerical analysis of footing settlement problem will be shown in this section (Mase and Hashiguchi 2009). The prediction of peak load and post-peak behavior for the footing-settlement problem on sands having the high friction and dilatancy cannot be attained in fact by the usual implicit finite element method requiring the repeated calculations of total stiffness equation which needs quite large calculation
424
13 Constitutive Equations of Soils
time. On the other hand, it can be attained by the explicit dynamic relaxation method in which the dynamic equilibrium equation is solved directly without solving the total stiffness equation so that the calculation time is drastically reduced. The FLAC3D (Fast Lagrangian Analysis of Continua in 3 Dimensions; Cundall and Board 1988; Itasca Consulting Group 2006) based on the explicit dynamic relaxation method, i.e. the finite difference method (FDM) is adopted in the present analysis, in which the initial subloading surface model of soils with the automatic controlling function to attract the stress to the normal-yield surface is implemented as the constitutive equation. The calculation is executed by the forward Euler method without iteration calculation for convergence in this program by adopting small incremental steps so as not to influence on the calculation results, while this fact is examined prior to the calculation. The finite elements are composed of eight-node cuboidal elements. Each cuboidal element is divided into the two kinds of overlays, i.e. assembly of five tetrahedral sub-elements having different directions. Then, the deviatoric variables are analyzed using individual values in each tetrahedral sub-element. On the other hand, isotropic variables are analyzed using averaging values in five tetrahedral sub-elements in order to avoid the over-constraint problems common in finite element calculations for dilatant materials, i.e. the dilatancy locking. Test data used for numerical simulation The test data of footing settlement phenomenon on sand layers under the plane strain condition are simulated in the present analysis. The sizes of the test apparatus of type A (Tatsuoka et al. 1984) and type B (Tani 1986) have the same height 49 cm and depth 40 cm and the different widths 122 cm and 183 cm, respectively. The size of type C (Okahara et al. 1989) has the height 400 cm and the depth 350 cm and the widths 700 cm. The footings width, denoted as B, is taken 10 cm for the types A and B and 50 cm for the type C. The sand layers has been prepared carefully by the air-pluviation method for the dried Toyoura sand in order to obtain the same homogeneous layers but the test data exhibit dispersion more or less test by test despite of the laborious preparation work. Numerical analysis and comparison with test data The finite element meshes in the present analyses for the simulations of the test data are shown in Fig. 13.23. The nodal points of soil layer contacting with the footing and the bottom of soil bin are fixed to them, respectively. On the other hand, the nodal points at the side walls can move freely in the vertical direction. The right half of soil layer is analyzed in order to reduce the calculation time as has been done widely even for searching the localized deformation (cf. e.g. Sloan and Randolph 1982; Pietruszczak and Niu 1993; Stallebrass and Taylor 1997; Borja and Tamagnini 1998; Siddiquee et al. 1999; de Borst and Groen 1999; Sheng and Soloan 2000; Borja et al. 2003). First, the analysis of deformation caused by the gravity force is performed. Then, the vertical displacement of footing is imposed by incremental steps of 105 5 104 cm.
13.7
Numerical Analysis of Footing Settlement Problem
425
The material parameters in the subloading surface model are selected as F0 = 350 kPa;/c = 30 , nh = 0.001, # = 0.00003, ~ = 0:0015, j ~ = 0:00015 k m = 0:3,
/d = 29 , ld = 0.2 u = 15:0: The values of material parameters listed above are used for all the following numerical calculations because Toyoura sands having the same initial void ratio 0.66 are used in these tests. (Fig. 13.23) The comparisons of test and calculated results are shown in Fig. 13.24, where the simulation by Siddiquee et al. (1999) which is calculated by the finite element method adopting the Mohr–Coulomb model is also depicted in (c). In this figure qm is the average footing pressure, cd is the unit dry weight, Nc is the normalized footing pressure corresponding to the bearing capacity coefficient associated with frictional property of soil and S is the settlement. The qualitative trends of test results and the quantitative simulation to some extent are captured and the ultimate loads, i.e. bearing capacities are predicted well by the present analyses, although the analyses are performed for the sand with the high friction and dilatancy. Here, the post-peak behavior, i.e. the increase of load after exhibiting once the minimal value is also predicted well qualitatively. It would be provided by the adoption of the up-dated Lagrangian calculation realizing the accumulation of displacements by updating the positions of nodal points, which results in the upsurge of soils around the footing and thus the increase of footing load. However, the quantitative prediction of post-peak behavior would require the further study taking account of the tangential inelastic strain rate due to the stress rate tangential to the loading surface (Hashiguchi and Tsutsumi 2001) and the gradient effect (cf. Hashiguchi and Tsutsumi 2006) by introducing the shear-embedded model (cf. Pietruszczak and Mroz 1983; Tanaka and Kawamoto 1988; Tanaka and Sakai 1993) for example, which will be described in Sect. 20.3. The displacements of nodal points from the initiation of settlement are shown in Fig. 13.25 at the settlement 11, 15, 80 mm for Type A, B and C, respectively, which are the final stage of calculation. The Prantdl’s slip line solution with the triangle wedge, the logarithmic spiral zone and the passive Rankine zone is observed clearly in this figure (Fig. 13.25). On the other hand, the soils in the periphery of footing inevitably experience the null or further negative pressure since they are pushed upward as the footing settlement proceeds (see Fig. 13.26). It causes the singularity of plastic modulus for the normal-yield surface passing through the origin of stress space at which the normal-yield and the subloading surfaces contact with each other. This defect is improved in the present model by making the normal-yield surface translate to the
426
13 Constitutive Equations of Soils
C.L.
56cm
49cm
5cm
(a) Type A (B: 10cm) C.L.
86cm
49cm
5cm
(b) Type B (B: 10cm) C.L.
325cm
400cm
25cm
(c) Type C (B: 50cm) Fig. 13.23 Finite element meshes
13.7
Numerical Analysis of Footing Settlement Problem
427
300
300
FDM by subloading surface model Test (Tatsuoka et al., 1984): e=0.66
FDM by subloading surface model Test (Tani, 1986): e=0.669,BC2 Test (Tani, 1986): e=0.669,BC3
Test (Tatsuoka et al., 1984): e=0.66
200
200
100
100
0
0.00
0.05
0.10
0
0.15
0.00
0.05
0.10
Relative settlement S B
Relative settlement S B
(a) Type A (B: 10cm)
(b) Type B (B: 10cm)
300 FDM by subloading surface model FEM (Siddiquee et al., 1999)
Test (Okahara et al., 1989): e=0.66 Test (Okahara et al., 1989): e=0.66
200
0.15
qm : Average footing pressure : Unit dry weight S : Settlement B: Footing width
100
0
0.00
0.05
0.10
0.15
Relative settlement S B (c) Type C (B: 50cm)
Fig. 13.24 Comparisons of test and calculated results for footing settlement phenomenon
region of negative pressure as shown in Fig. 13.13, whilst the numerical difficulty can be avoided although the translation was taken quite small as 1/1000 in size of the normal-yield surface. In addition, the impertinence of infinite elastic volume expansion is avoided by shifting the isotropic consolidation characteristic into the negative range of pressure as shown in Fig. 13.1. It should be emphasized that the stable analysis cannot be executed without these improvements. The pertinent result for the footing-settlement problem on the sand with a high friction, one of the difficult problems in soil mechanics, is obtained in the present study as described above. Here, the peak, the subsequent reduction and the final
428
13 Constitutive Equations of Soils
Displacement (cm) 0.0䡚0.1 0.1䡚0.2 0.2䡚0.3 0.3䡚0.4 0.4䡚0.5 0.5䡚0.6 0.6䡚0.7 0.7䡚0.8 0.8䡚0.9 0.9䡚1.0 1.0䡚1.1
(a) Type A (B: 10cm) Displacement (cm) 0.0䡚0.2 0.2䡚0.4 0.4䡚0.6 0.6䡚0.8 0.8䡚1.0 1.0䡚1.2 1.2䡚1.4 1.4䡚1.5
(b) Type B (B: 10cm) Displacement (cm) Liv
0.00䡚1.00 1.00䡚2.00 2.00䡚3.00 3.00䡚4.00 4.00䡚5.00 5.00䡚6.00 6.00䡚7.00 7.00䡚8.00
(c) Type C (B: 50cm) Fig. 13.25 Deformed finite element meshes at final step
13.7
Numerical Analysis of Footing Settlement Problem
429
C.L.
p (kPa) 0.0001䡚0 0 䡚10 10䡚20 20䡚30 30䡚40 40䡚50 50䡚60 60䡚
I t
Fig. 13.26 Distribution of mean pressure for type A at final step
increase in footing load are predicted well qualitatively and quantitatively to some extent. The reasons for succession are summarized as follows: (1) The subloading surface model applied in the present analysis has the advantages: (i) It is furnished with the automatic controlling function to attract the stress to the yield surface, whilst all other elastoplastic constitutive models are required to incorporate a return-mapping algorithm to pull back the stress to the yield surface in the plastic deformation process: (ii) It is capable of describing the softening behavior and dilatancy characteristics quite realistically, predicting the simultaneous occurrence of the peak load and the highest dilatancy rate as was found experimentally by Taylor (1948). (iii) By virtue of the model extension to accommodate negative pressure, it has the full regularity since the normal-yield surface does not pass through the zero stress point and thus the subloading surface is always determined uniquely. In addition, the elastic property is improved such that the elastic bulk modulus does not become zero even for null stress or negative pressure inside the normal-yield surface. (2) The finite difference program FLAC3D adopted in the present study is based on the explicit-relaxation method which enables us to shorten the calculation time drastically since it is not required to solve the total stiffness matrix.
430
13 Constitutive Equations of Soils
13.8
Hyperelastic Equation of Soils
The isotropic hyperelastic constitutive equation for soils independent of the third invariant of stress and strain is given by r¼
@wðeev ; eed Þ @wðeev ; eed Þ @eev @wðeev ; eed Þ @eed ¼ þ @ee @ee @ee @eev @eed
ð13:178Þ
where ee is the infinitesimal elastic strain tensor and eev tree ; eed kee 0 k leading to @eev @tr ee ¼ ¼ I; @ee @ee
0
0
@eed ee ee ¼ ¼ 0 @ee kee k eed
ð13:179Þ
Equation (13.178) is rewritten by Eq. (13.179) as 0 @w eev ; eed @w eev ; eed re Iþ r¼ eed @eev @eed
ð13:180Þ
The strain energy function of soils was first proposed by Houlsby (1985) and subsequently modified by Borja and Tamagnini (1998), Tamagnini et al. (2002), in which the ln v ln p linear relation in Eq. (13.8) under pe ¼ 0 in Sect. 13.1 is incorporated and the dependence of the shear modulus on the pressure is taken account. Mechanical and mathematical properties of this model have been studied by Callari et al. (1998), Niemunis and Cudny (1998), Houlsby et al. (2005), Amorosi et al. (2007), etc. In what follows, the hyperelastic equation of soils (Hashiguchi 2018c) is shown in the following, which possesses the identical forms for the infinitesimal and the finite strains. Firstly, we introduce the following strain energy function based on Eqs. (13.12)1 and (13.16)1. wðeev ; eed Þ
¼
pe eev
e e ev e ~ðp0 þ pe Þ exp þ G0 exp n v ee2 þj d ~ ~ j j
ð13:181Þ
It follows for Eq. (13.181) that e h e i 8 e e < @wðeve;ed Þ ¼ pe ðp0 þ pe Þ exp ev n G0 exp n ev eed 2 ~ ~ ~ j j j @ev h i e e : @wðeve;ed Þ ¼ 2G0 exp n eev ee d ~ j @e d
ð13:182Þ
13.8
Hyperelastic Equation of Soils
431
Substituting Eq. (13.182) into Eq. (13.180) leads to e e e n e r ¼ fpe ðp0 þ pe Þ exp v G0 exp n v eed 2 gI ~ ~ ~ j j j e e 0 þ 2G0 exp n v re ~ j
ð13:183Þ
The hydrostatic and deviatoric parts in Eq. (13.183) are given by e p þ pe e ¼ exp v for ee 0 ¼ O ðeed ¼ 0Þ ~ j p0 þ pe i.e.
eev
p þ pe ¼ ~ j ln p0 þ pe
for ee 0 ¼ O ðeed ¼ 0Þ
ð13:184Þ
and e ev p þ pe n e 0 e0 e ¼ 2G0 e r ¼ 2G0 exp n ~ j p0 þ pe 0
ð13:185Þ
which conform to Eqs. (13.12)1 and (13.16)2 in the rate form, respectively, while Eq. (13.184) is used for deriving Eq. (13.185) from the second term in the right-hand side of Eq. (13.183). The general assessment of the various elastic constitutive equations for soils can be referred to Yamakawa (2021).
Chapter 14
Viscoplastic Constitutive Equations with Subloading Surface Concept
The plastic deformation depends on the rate of deformation in general. The time (or rate)-dependent inelastic deformation is called the viscoplastic deformation. The viscoplastic constitutive models are classified into the creep model and the overstress model. The viscoplastic deformation is induced depending on the ratio of the magnitude of the stress to the magnitude of the yield stress in the creep model and thus the creep model is not reduced to the elastoplastic constitutive equation in the quasi-static deformation process. On the other hand, the viscoplastic deformation is induced by the overstress from the yield surface and thus the overstress model is reduced to the elastoplastic constitutive equation in the quasi-static deformation process. Therefore, the creep model is unacceptable since it is unreasonable such that the creep model inappropriately predict a viscoplastic deformation for any low stress level and thus it cannot describe both of the rate-independent and the rate-dependent deformation behaviors by a single equation. On the other hand, the overstress model is capable of describing both of the rate-independent and the rate-dependent deformation behaviors by a single constitutive equation. However, the existing overstress model possesses the defects: (1) The plastic deformation during the cyclic loading process by the rate of stress within the yield surface cannot be described, (2) The purely elastic deformation is described for the impact loading, resulting in the inapplicability to the deformation behavior in a high strain rate. The rational overstress model is formulated by incorporating the extended subloading surface model in this chapter.
14.1
Rate-Dependent Deformation of Solids
The elastic and the inelastic deformations of solids are induced by the deformation of solid particles themselves (crystals in metals, soil particles in soils, etc.) and the mutual slips between them, respectively. Therefore, © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_14
433
434
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
1. The elastic tangent modulus is high, since the elastic deformation is induced by the deformations of solid particles themselves, the rete-dependent deformation of which is described by the viscoelastic constitutive equation, 2. High stress has to be applied to solids in order that inelastic deformation is induced, which overcomes the friction resistance between solid particles. The stress inducing the inelastic rate-independent deformation is macroscopically the so-called yield stress. 3. The rate-dependent inelastic deformation is induced when an overstress from the yield surface is applied to solids, 4. The tangent modulus in the inelastic deformation process lowers from that in the elastic deformation process by the inelastic relaxation. Consequently, the rate-dependent elastic deformation induced in the stress lower than the yield stress is called the viscoelastic deformation. On the other hand, the rate-independent inelastic deformation induced by the yield stress and the rate-dependent inelastic deformation induced over the yield stress are called the plastic deformation and the viscoplastic deformation, respectively. Then, they are classified as follows: 8 Deformation of Rate-independence: Elastic constitutive equation > > > < soild particles Rate-dependence: Viscoel astic constitutive equation Deformation of soilds > Mutual slips of Rate-independence: Plastic constitutive equation > > : solid particles Rate-dependence: Viscoplastic constitutive equation
which is illustrated in Fig. 14.1. The strain e is additively decomposed into the elastic strain ee and the viscoplastic strain rate evp , i.e. e ¼ ee þ evp ;
e ¼ ee þ evp
ð14:1Þ
Elasto-viscoplastic deformation
Elastoplastic deformation induced in the quasi-static rate
Yield stress Range of viscoelastic deformation
Elastic response
Yield stress
Fig. 14.1 Rate-dependent deformation of solids
14.1
Rate-Dependent Deformation of Solids
435
The stress versus elastic strain relation is described by Eq. (8.4). When the strain energy function is given by the quadratic equation in Eq. (8.3), the following relations hold. r ¼ E : ee ¼ E : ðe evp Þ; ee ¼ E1 : r; e ¼ E1 : r þ evp r ¼ E : e e ¼ E : e e vp ; e e ¼ E1 : r; e ¼ E : r þ e vp
ð14:2Þ ð14:3Þ
The stress r is calculated by substituting the plastic strain evp calculated by the plastic constitutive equation into Eq. (14.2). Here, let the viscoplastic strain evp be vp decomposed into the storage part evp ks and the dissipative part ekd for the kinematic vp hardening and let it be decomposed into the storage part ecs and the dissipative part evp cd for the elastic-core, i.e.
14.2
vp evp ¼ evp ks þ ekd ;
vp e vp ¼ e vp ks þ e kd
ð14:4Þ
vp evp ¼ evp cs þ ecd ;
vp e vp ¼ e vp cs þ e cd
ð14:5Þ
History of Viscoplastic Constitutive Equations
The most pertinent viscoplastic model would be the overstress model. The development of this model is reviewed in this section, while the overview of the history is portrayed simply by the one-dimensional deformation with the rheology models in Fig. 14.2. The elastic constitutive equation extended so as to describe the rate-dependence is called the viscoelastic constitutive equation and one of the typical models is the Maxwell model, in which the spring and the dashpot are connected in series. Therefore, the strain rate e is additively decomposed into the elastic strain rate e e ¼ E 1 r and the viscous strain rate e v ¼ l1 r, where r designates the stress, E is the elastic modulus and l is the viscous coefficient, leading to
e ¼ e e þ e v ¼ E 1 r þ l1 r
ð14:6Þ
This model is concerned with the rate-dependent deformation at the low stress level below the yield stress. On the other hand, the elastoplastic constitutive equation for a quasi-static deformation can be schematically expressed by the Prandtl model in which the dashpot is replaced with the slider in the Maxwell model, whereas the slider begins to move in the state that the stress r reaches the yield stress ry , by which the plastic
436
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Fig. 14.2 Development of viscoplastic model
strain rate is induced (see Fig. 14.2). Then, the strain rate is additively composed of the elastic and the plastic strain rates, i.e. ( e e ¼ E 1 r for r\ry e¼ e e þ e p ¼ E 1 r þ M p1 r for r ¼ ry
ð14:7Þ
where M p is the plastic modulus. Furthermore, the model which describes the rate-dependent plastic strain rate e vp induced for the state of stress over the yield stress, called the viscoplastic strain rate, was introduced by Bingham (1922), combining the above-mentioned Maxwell model and Prandtl model so as to connect the dashpot and the slider in parallel as
14.2
History of Viscoplastic Constitutive Equations
437
shown in Fig. 14.2, where l is the viscoplastic coefficient and n is the material constant, while n is chosen to be 4 8 in practice. Then, the strain rate is given by
e ¼ e e þ evp ¼ E 1 r þ l1 \r ry [ n
ð14:8Þ
The Bingham model for the elasto-viscoplastic deformation is the origin of the overstress model based on the concept that the viscoplastic strain rate is induced by the overstress, i.e. stress over the yield stress. The above-mentioned Bingham model for the one-dimensional deformation was extended by Hohenemser and Prager (1932) and Prager (1961) to describe the three-dimensional deformation of metals, adopting the Mises yield condition for the slider as shown in Fig. 14.2. The viscoplastic strain rate e vp is given by vp
e where eeqvp
n 1 r eq r0 1 ¼ eqvp l F ðe Þ kr0 k
ð14:9Þ
pffiffiffiffiffiffiffiffi R 0 2=3 e vp dt is the equivalent viscoplastic strain given by
replacing the plastic strain rate e p to the viscoplastic strain rate e vp in the plastic pffiffiffiffiffiffiffiffi 0 equivalent strain eeqp 2=3ep dt in Eq. (8.43). The viscoplastic coefficient l depends on stress, internal variables and temperature in general. Furthermore, the viscoplastic strain rate in the Prager’s overstress model was extended by Perzyna (1963, 1966) for materials having the general yield condition unlimited to the Mises yield condition as
e vp ¼
n 1 f ðrÞ 1 n; l FðHÞ
n
@f ðrÞ @f ðrÞ = @r @r
ð14:10Þ
Then, substituting Eqs. (8.5) and (14.10) into Eq. (14.1), we have n 1 f ðrÞ 1 n l FðHÞ
ð14:11Þ
n 1 f ðrÞ 1 E:n r ¼ E : e l FðHÞ
ð14:12Þ
e ¼ E1 : r þ and thus
where the hardening variable H evolves by H ¼ H r; H; e vp
ð14:13Þ
438
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
by replacing the plastic strain rate e p to the viscoplastic strain rate e vp in the evolution rule of the isotropic hardening variable in Eq. (8.23) for the plastic con stitutive equation. Therefore, H_ is the homogeneous function of e vp in degree-one. In what follows, the isotropic yield condition f ðrÞ ¼ FðHÞ in Eq. (8.16) is used below for the sake of simplicity in explanation up to Sect. 14.4. It should be noted that the response of the overstress model is reduced to that of the elastoplastic constitutive equation because f ðrÞ=FðHÞ 1 ffi 0, i.e. f ðrÞ ffi FðHÞ holds in the quasi-static deformation: dr=dt ffi O and de=dt ffi O. In other words, the overstress model is capable of describing the elastoplastic deformation behavior so that the elastoplastic constitutive equation can be disused if the overstress model is used.
14.3
Irrationality of Creep Model
Based on a concept different from the overstress model, the creep model, which also aims at describing the viscoplastic deformation, has been studied widely. The typical one is the Norton law (Norton 1929) in which the creep strain rate is given as follows:
e c ¼ d0c krkm n
ð14:14Þ
where d0c and mð 1Þ are the material constants. Further, it has been modified by Rice (1970, 1971), Bonder and Partom (1975), Lemaitre and Chaboche (1990), Chaboche (1997, 2008), Chaboche et al. (1979), Ohno et al. (2018, 2021), etc. (see also de Sauza Neto et al. 2008) for the viscoplastic deformation and Nakada and Keh (1966), Hutchinson (1976), Pan et al. (1983), Peirce et al. (1983), etc. for the crystal plasticity as follows:
m c c f ðrÞ n e ¼ e n ¼ d0 FðHÞ
c
ð14:15Þ
where the rate of the isotropic hardening variable is given by H ¼ H r; H; e c
ð14:16Þ
instead of Eq. (14.13). The strain rate is given by
e ¼ e e þ e c ¼ E1 : r þ d0c krkm n for Eq. (14.14) and
ð14:17Þ
14.3
Irrationality of Creep Model
439
P
l0
f (σ )/ F ( H ) = 1 yield state
Over - run
Elastically decrease
0
Overstress model
0
Creep model
(a) Unloading behavior by reduction of stress
f (σ )/ F ( H ) = 1 yield state
P
f (σ )/ F ( H ) = 1 yield state
P
Stop
Unlimitedly proceed
l
Overstress model
0
Creep model
∞
0
Bar is stretched endlessly only by dead weight in creep model.
f (σ )/ F ( H ) = 1 P yield state
(b) Creep deformation behavior under constant stress
Fig. 14.3 Comparison of overstress model and creep model: The latter exhibits unrealistic behavior predicting always creep strain rate
e ¼ e e þ e c ¼ E1 : r þ d0c
f ðrÞ FðHÞ
m n
ð14:18Þ
for Eq. (14.15). The creep model is regarded to be merely the nonlinearization of the Maxwell model whose rheology model does not possess the slider which is the basic element for the description of irreversible deformation. The creep model described above has different structures from the overstress model because Eqs. (14.14) and (14.15) possess no threshold value for the generation of the creep strain rate so that the creep strain rate is induced for any low stress level. Then, its deformation behavior is not reduced to that of the elastoplastic constitutive equation at quasi-static deformation. Therefore, this model cannot describe appropriately the deformation behavior at a low strain rate. In fact, the creep strain rate does not diminish even if the stress decreases into the inside of yield surface, exhibiting the overrunning stress–strain curve and the creep deformation proceeds unlimitedly under a constant stress state in the creep model as shown in Fig. 14.3. As a concrete example, it is quite unnatural that a bar subjected to any low tension continues to be elongated endlessly as a time elapses. In addition, it is incapable of describing appropriately the impact loading behavior as it describes an elastic deformation behavior with an infinite strength. On the other hand, the viscoplastic strain rate diminishes immediately after the stress decreases into the inside of yield surface in the overstress model as is observed in real materials.
440
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Consequently, the creep model is not physically accepted but the overstress model is appropriate for the description of the rate-dependent plastic behavior. As described above, the creep model is physically irrational possessing the mechanical structure different basically from the elastoplasticity because the inelastic strain rate is induced even inside the yield surface. Nevertheless, Ohno et al. (2018), Wu et al. (2018), etc. incorporated the creep model into their cylindrical yield surface model in order to describe the rate-dependent deformation behavior, resulting in the ad hoc model by which the rate-independent and the rate-dependent deformation behaviors cannot be described in a unified equation. Needless to say, the deformation behavior in a general rate varying from the quasi-static to the non-static rate cannot be analyzed by such ad hoc method. Furthermore, various rate-dependent constitutive models including the time itself have been proposed to date (cf. e.g. Conway 1967; Penny and Marriott 1971; Boyce et al. 1988). Here, however, note that the determination of time when the creep deformation starts and stops is accompanied with an ambiguity depending on the subjectivity of observers, especially in a fluctuating rate of deformation. Therefore, these models are impertinent, lacking the objectivity.
14.4
Mechanical Response of Past Overstress Model
The development of rate-dependent elastoplastic constitutive equation is reviewed above and it is described that the overstress model would have a pertinent basic structure. Here, let the mechanical responses at the infinitesimal and the infinite rates of deformation be examined in order to clarify the basic property of this model. The past overstress model advocated by Bingham (1922) and extended by Hohenemser and Prager (1932), Prager (1961) and Perzyna (1963, 1966), Guo et al. (2013) (see also de Souza Neto et al., 2008) describes the elastoplastic deformation in the quasi-static loading since the viscous resistance of the dashpot diminishes but the elastic deformation in the impact loading since the viscous resistance of the dashpot becomes infinite. Therefore, it cannot be applied to the deformation behavior in a high rate as an impact loading. In what follows, we will show this fact on the past overstress model. Equations (14.11) and (14.12) are expressed in the following equations for the incremental forms. n 1 f ðrÞ 1 n dt l FðHÞ
ð14:19Þ
n 1 f ðrÞ 1 E : n dt l FðHÞ
ð14:20Þ
de ¼ E1 : dr þ dr ¼ E : de
14.4
Mechanical Response of Past Overstress Model
441
Equation (14.20) is reduced to the following relation for the infinitesimal rate of deformation (quasi-static deformation) fulfilling dt ! 1: dr=dt ! O and de=dt ! O. n 1 f ðrÞ 1 E: n OffiO l FðHÞ
ð14:21Þ
f ðrÞ 1!0 FðHÞ
ð14:22Þ
leading to
Then, the stress changes fulfilling the yield condition f ðrÞ ¼ FðHÞ, i.e. obeying the plastic constitutive relation, while the elastic deformation is given by the change of stress, so that Eqs. (14.11) and (14.12) exhibit the response of the elastoplastic constitutive relation in the quasi-static deformation as shown in Fig. 14.4. Then, the overstress model is the extension of the elastoplastic constitutive equation to the non-zero rate of deformation. In fact, the elastoplastic constitutive relation is reproduced by the quasi-static deformation in the overstress model. Here, it is reproduced easily for the non-viscous material, i.e. the material with a small viscoplastic coefficient l ffi 0 causing f ðrÞ ¼ FðHÞ, noting that the Bingham model is reduced to the Prandtl model for l ffi 0 in Fig. 14.2. On the other hand, in the infinite rate of deformation fulfilling dt ffi 0: dr=dt ! 1 and de=dt ! 1, Eq. (14.12) is reduced to
dr ¼ E : de O; i.e:r ¼ E : e O
ð14:23Þ
Elastic
Impact loading ||ε ||
|| ε || increase
Overstress
0
R =1
Quasi-static loading || ε || 0
ε
Fig. 14.4 Past overstress model which is inapplicable to prediction of deformation behavior at high rate
442
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
approaching the elastic response as shown in Fig. 14.4. Therefore, it predicts the unrealistic response that the material can bear an infinite load. The past overstress model is inapplicable to the prediction of deformation at high rate in general. In order to modify this defect, Lemaitre and Chaboche (1990) and Chaboche (2008) proposed to add the creep strain rate in addition to the elastic and the viscoplastic strain rate. This modification is irrational spoiling the rational advantage of the overstress model. The rigorous modification of the overstress model so as to describe the elasto-viscoplastic behavior in the general rate ranging from the quasi-static to the impact loading can be attained by the subloading-overstress model described in the next section. Eventually, the existing formulation of overstress model in Eq. (14.11), i.e. (14.12) is inapplicable to the prediction of deformation at a high rate. The material constant n included as the power form in Eq. (14.11), i.e. (14.12) is usually selected to be larger than five, but the fitting to the test data for impact load is impossible even if n is selected as one hundred which, needless to say, results in the inappropriate prediction of deformation in a slow loading process. In addition, the inclusion of a high power in the equation induces difficulty in numerical calculations. The so-called flow stress model was proposed by Johnson and Cook (1983) in which the yield stress, i.e. flow stress depends not only on the viscoplastic strain rate but also on the viscoplastic strain rate. It is the empirical model without a generality, which is concerned with the fast loading behavior but inapplicable to the slow loading behavior. However, unfortunately it is used widely for the deformation analyses in the fast loading behavior by adopting the commercial FEM software, e.g. Abaqus and LD-DYNA. Hereinafter, it is desirable to adopt the subloading-overstress model for the accurate analyses of the general loading behavior ranging from the quasi-static to the impact loadings, which will be described in the next section.
14.5
Subloading Overstress Model: Extension to Description of General Rate of Deformation
The subloading surface model will be extended to the subloading-overstress model in this section, which is capable of describing the deformation in general rate from quasi-static (elastoplastic) to the impact loading. All the aforementioned defects in the traditional overstress model are excluded in the subloading-overstress model.
14.5.1
Static and Limit Subloading Surfaces
The static subloading and the limit subloading surfaces (see Fig. 14.5) are incorporated in addition to the subloading, the normal-yield, the elastic-core and the limit
14.5
Subloading Overstress Model: Extension to Description …
443
Fig. 14.5 Limit subloading, subloading, normal-yield, static-subloading surface, limit elastic-core, elastic-core and static-subloading surfaces in subloading-overstress model
elastic-core surfaces in the subloading surface model for the rate-independent elastoplastic deformation described in Chap. 11. Here, the subloading surface on which the current stress lies may become larger than the normal-yield surface in general. The normal-yield ratio R which may become larger than unity in general is calculated by the identical equation for the rate-independent elastoplastic deformation, i.e. Eq. (11.42) in general and by Eq. (12.32) for metals and by Eq. (13.164) for soils. The viscoplastic strain rate is induced by the overstress from the static subloading surface expressed in the following equation which is given by replacing the current stress r, the center a and the normal-yield ratio R to their conjugate points rs , as and the static normal-yield ratio Rs ð 1Þ, respectively, in the subloading surface in Eq. (11.2) for the elastoplastic deformation. f ðrs Þ ¼ Rs FðHÞ
ð14:24Þ
with rs ¼ rs as ¼
Rs rs as r a rc ; rs c ¼ Rs ðry cÞ ¼ Rs r ¼ R Rs R R ð14:25Þ
444
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Fig. 14.6 Function
in the evolution rule of static-normal-yield ratio: (a) Re = 0, (b) Re > 0
where Rs evolves by the following equation which is identical to Eq. (9.9) with Eq. (11.53) for the extended subloading surface model with the replacement of the
plastic strain rate e p to the viscoplastic strain rate e vp (see Fig. 14.6). O
ð14:26Þ
with ð14:27Þ Rs ð 1Þ is called the static normal-yield ratio because the quasi-static deformation proceeds in the state R ¼ Rs . It can be analytically calculated by Eq. (9.15) with the replacement of the plastic strain rate to the viscoplastic strain rate for = const. as follows:
under the initial condition Rs ¼ Rs0 for e ¼ vp
evp 0 ,
where e ¼ vp
R
ð14:28Þ jje jjdt. The vp
inequality R s \0 holds for Rs [ 1 in Eq. (14.26)1 with Eq. (14.27), bringing about the high numerical efficiency.
14.5
Subloading Overstress Model: Extension to Description …
14.5.2
445
Viscoplastic Strain Rate
Now, let the flow rule of the viscoplastic strain rate be given by incorporating the crucially important variable in the subloading surface model, i.e. the normal-yield ratio R into the past overstress model (Prager 1961; Perzyna 1963, 1966) as follows: vp
e ¼ C n
ð14:29Þ
where C is the positive viscoplastic multiplier given by C
1 hR Rs in lv
ð14:30Þ
or exp[
]
,
ð14:31Þ
sinh
lv (viscoplastic coefficient) and n are the material constants (Hashiguchi 2022a). The smooth elastic-viscoplastic transition is described by adopting the overstress due to R RS instead of R 1, because RS increases gradually up to 1 with the viscoplastic deformation, while RS was fixed as RS ¼ 1 in the past overstress model. The function U for the evolution rule of the normal-yield ratio R and the positive _ are modified to Eqs. (11.53) and (11.54) by incorporating proportionality factor k ^c : n Þ in the extended subloading surface and Cn ð n the variables model for the rate-independent elastoplastic deformation. Concurrently, the function Us for the evolution rule of the static normal-yield ratio Rs is given by Eq. (14.27). Analogously, let Eqs. (14.30) and (14.31) for viscoplastic multiplier C be modified to the following equation (see Fig. 14.7). ð14:32Þ or ð14:33Þ where uc is the material constant. Thus, the rising and the lowering of the stress are and , respectively. It leads to the description enforced in the state of the global Masing effect in the viscoplastic deformation process as shown in Fig. 14.7.
On uniaxial deformation at constant strain rate ðea ¼ constantÞ
The axial stress rate ra increases rapidly (almost elastically) at the beginning of the
deformation because the variable R Rs is small and thus e vp a is small as
e e vp a e a ffi e a . Thereafter, as the elastic-core increases toward
, the
446
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
normalyield ratio R increases rapidly for a certain increase of stress and thus R Rs increases rapidly so that the viscoplastic strain rate develops rapidly leading to
e e vp a ¼ e a ðe a ¼ 0Þ resulting in r a ¼ 0 (peak stress state). The further increase of induces the large viscoplastic strain rate R Rs caused by the increase of
e vp [ e leading to e e ð¼ e e p Þ\0, i.e. ra \ 0 if we use Eq. (14.30) or (14.31). Then, the unrealistic stress-strain curve exhibiting the peak point and the decrease of stress is predicted by these equations. This defect can be remedied by sup
. It can be realized by pressing the viscoplastic strain rate e vp a with the increase of leading to dividing the positive viscoplastic multiplier C by the variable the suppression of the viscoplastic strain rate with the increase of R c as shown in Eq. (14.32) or (14.33) (see Fig. 14.7). Further, let Eqs. (14.32) and (14.33) be extended such that the infinite viscoplastic strain rate is induced for R ! cm Rs ðcm : material constantÞ leading to C ! 1 in the impact loading process as follows: ð14:34Þ
Fig. 14.7 Modification of stress vs. viscoplastic strain curve by incorporation of variables Cn leading to expression of generalized Masing effect
and
14.5
Subloading Overstress Model: Extension to Description …
447
or ð14:35Þ The variable cm Rs ð cm Þ is called the limit normal-yield ratio. Here, the surface to which the stress can reach at most is given by setting R ¼ cm Rs in the subloading surface in Eq. (11.2) is called the limit subloading surface. On the other hand, the elastic response is described unrealistically for the impact loading in the past overstress models. Further, the quite irrational model was proposed by Lemaitre and Chaboche (1990) incorporating the extra creep strain rate in addition to the elastic and the overstress strain rates. The rational description of deformation in the general rate ranging from the quasi-static to the impact loading is attained by the above-mentioned subloading-overstress model in which the crucially-important variables, i.e. the and the limit normal-yield ratio R, the static normal-yield ratio normal-yield ratio cm Rs are incorporated. The rates of the internal variables H; a and c are given from Eqs. (11.12), (11.13)
nÞ as follows: and (11.25) with the replacement of e p to e vp ð¼ C
H ¼ CfHn
a ¼ Cf kn
c ¼ Cf cn :
ð14:36Þ ð14:37Þ ð14:38Þ
The isotropic hardening stagnation can be incorporated by the identical equa
tions formulated in Sect. 12.2 with the replacement of the plastic strain rate e p to
the viscoplastic strain rate e vp .
14.5.3
Strain Rate Versus Stress Rate Relation
The strain rate versus stress rate relations are given from Eqs. (14.3) and (14.29) as follows: e ¼ E1 : r þ Cn r ¼ E : e CE : n
ð14:39Þ
which is represented in the incremental form as follows:
de ¼ E1 : dr þ Cndt dr ¼ E : de CE : ndt
ð14:40Þ
Then, it follows from Eq. (14.40) with Eq. (14.30) or (14.31) in the quasi-static deformation that
448
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Fig. 14.8 Stress–strain curve predicted by subloading-overstress model
C ¼ 0 leading to R ¼ RS for jjdejj=dt ! 0 and jjdrjj=dt ! 0
ð14:41Þ
so that the stress changes along the static subloading surface given by Eq. (14.24). Consequently, the response of the subloading–overstress mode is reduced to that of the original subloading surface model in the rate–independent elastoplastic deformation behavior described in Chaps. 8–13. The subloading–overstress model is no more than the generalization of the subloading surface model to the description of the elasto-viscoplastic deformation in the general strain rate. Irreversible deformations in any rate from the static to the impact loading can be described by the subloading— overstress model. Eventually, the elastoplastic constitutive equation with the plastic modulus for the rate–independent elastoplastic deformation which is derived by consistency condition of the yield condition (subloading surface for the subloading surface model) can be disused by adopting only the subloading–overstress model. In the overstress model, we do not need to calculate the plastic modulus which possesses often a complex form as seen in Eq. (11.45) but instead we have only to perform the update calculation of the viscoplastic internal variables involved in the positive viscoplastic multiplier C in Eq. (14.29) and n. The stagnation of isotropic hardening is incorporated by replacing the plastic
strain rate e p to the viscoplastic strain rate e vp in the formulation described in Sect. 12.2. The stress–strain curve described by the subloading-overstress model is illustrated in Fig. 14.8. The smooth elastic-viscoplastic transition and the cyclic loading behavior can be described appropriately. The stress is given from Eqs. (8.4) and (14.29) by ð14:42Þ
14.5
Subloading Overstress Model: Extension to Description …
449
The second term in the right-hand sides is induced by the viscoplastic deformation caused from the mutual viscoplastic slips of material particles, which is called the viscoplastic (stress) relaxations. It should be emphasized that the distinguished advantages in the subloading-overstress model is realized by the incorporation of the basic variable in the subloading surface model, i.e. the normal-yield ratio R in which various important information on the hardening/softening, the cyclic loading behaviors, etc. are incorporated. On the other hand, the overstress models other than the subloading-overstress model can never succeed to describe the plastic and the viscoplastic deformation behaviors rigorously, because the normal-yield ratio R has never been incorporated in them, although various complicated exponential-type, hyperbolic sine-type, etc. functions have been incorporated for the positive viscoplastic multiplier C as reviewed by Chaboche (2008). The computer program for the subloading-overstress model for metals is given in the Appendix L(c).
14.6
Comparison with Test Data
The capability of the subloading-overstress model for describing the deformation behavior of metals in the general rate of deformation including the quasi-static deformation has been verified performing the simulation of various test data by Hashiguchi et al. (2023). Some of the simulation results are shown in this section.
14.6.1
Dynamic Loading Process Inducing Elastic-Viscoplastic Deformation
The monotonic loading behaviors at various constant strain rates at 550° C in the test data for Modified 9Cr-1 Mo steel after Abel-Karim and Ohno (2000) and their simulations are shown in Fig. 14.9, where the material parameters are chosen as follows: Material constants: Elastic moduli: E ¼ 150; 000 MPa; m ¼ 0:3; ( isotropic: sr ¼ 0:23; cH ¼ 200; Hardening kinematic: ck ¼ 10 MPa; bk ¼ 0:1; Evolution of normal yield ratio: u ¼ 1; 000; uc ¼ 6; Re ¼ 0:2; Translation of elastic core: ce ¼ 5; v ¼ 0:7; c ¼ 5; n ¼ 4; cm ¼ 1:9 Viscoplastic deformation: lv ¼ 2; 000; u
450
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Fig. 14.9 Uniaxial loading behavior of Modified 9Cr-1 Mo steel under various axial strain rates at 550°C (test data after Abdel-Karim and Ohno 2000)
Initial values: Isotrophic hardening function: F0 ¼ 230MPa: The quite close simulations are also seen in Fig. 14.9. The accuracy of the present simulations are same levels of the simulations by the experimenters Abdel-Karim and Ohno (2000) based the creep model, who have provided the test data, while the creep model is physically irrational lacking the generality, i.e. incapable of describing the rate-independent elastoplastic deformation behavior as explained in Sect. 14.3. The deformation behavior at the variable strain rate between 0.001/s and 0.00001/s for 304 stainless steel at room temperature 20°C in the test data after Krempl (cf. Lemaitre and Chaboche 1979) is shown in Fig. 14.10, where the material parameters are chosen as follows:
14.6
Comparison with Test Data
451
Fig. 14.10 Uniaxial loading behavior under various axial strain rate between 0.1 and 0.001 (%/s) of 304 stainless steel at room temperature 20°C (test data after Krempl: cf. Lemaitre and Chaboche 1990)
Material constants: Elastic moduli: E ¼ 200; 000 MPa; m ¼ 0:3; ( isotropic: sr ¼ 0:2; cH ¼ 200; Hardening kinematic: ck ¼ 500 MPa; bk ¼ 0:5; Evolution of normal yield ratio: u ¼ 1; 000; uc ¼ 6; Re ¼ 0:2; Translation of elastic core: ce ¼ 200; v ¼ 0:7; c ¼ 10; n ¼ 4; cm ¼ 2 Viscoplastic deformation: lv ¼ 5; u Initial values: Isotrophic hardening function: F0 ¼ 160 MPa: The quite close simulation is shown in Fig. 14.10.
14.6.2
Quasi-static Loading Process Inducing Elastoplastic Deformation Behaviors
It was verified in the previous subsection that the subloading-overstress model is capable of describing the elasto-viscoplastic deformations induced in the dynamic
452
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Fig. 14.11 Mechanical ratcheting during pulsating loading between 0 and +830MPa of 1070 steel at room temperature (Test data after Jiang and Zhang 2008)
loading process. Further, it will be shown below that the subloading-overstress model is capable of describing the elastoplastic deformations induced in the quasi-static loading process at the room temperature by the simulations of test data. The comparison with the test data for the 1070 steel under the cyclic loading of axial stress between 0 and +830MPa at room temperature after Jiang and Zhang (2008) is depicted in Fig. 14.11, where the material parameters and the strain rate are selected as shown in the following, while the axial strain state is chosen to be
quite low, i.e. e a ¼ 107 =s, although the strain rate is not written in the paper of Jiang and Zhang (2008). Material constants: Elastic moduli: E ¼ 160; 000 MPa; m ¼ 0:3; ( isotropic: sr ¼ 0:61; cH ¼ 155; Hardening kinematic: ck ¼ 3; 000 MPa; bk ¼ 0:5; Evolution of normal yield ratio: u ¼ 90; uc ¼ 6; Re ¼ 0:5; Translation of elastic core: ce ¼ 7; 000; v ¼ 0:7; c ¼ 6; n ¼ 3; cm ¼ 5 Viscoplastic deformation: lv ¼ 5; 000; u
14.6
Comparison with Test Data
453
Fig. 14.12 Cyclic loading behavior under increasing strain amplitude of 316 steel at room temperature (Test data after Chaboche et al. 1979): (a) Isotropic hardening stagnation is considered, (b) Isotropic hardening is ignored
Initial values: Isotrophic hardening function : F0 ¼ 471 MPa: Quite close simulation is shown in Fig. 14.11, while the isotropic hardening stagnation is hardly relevant to the accuracy of simulation. Furthermore, examine the uniaxial cyclic loading behavior under the constant symmetric strain amplitudes to both postive and negative sides. Comparison with the test data of the 316 steel under the cyclic loading with the increasing axial strain amplitudes ±1.0, ±1.5, ±2.0, ±2.5, ±3.0% after Chaboche et al. (1979) is depicted in Fig. 14.12 where the material parameters and the strain rate are selected as shown in
the following, while the axial strain rate is chosen to be quite low, i.e. e a ¼ 108 =s, although the strain rate is not written in the paper of Chaboche et al. (1979). Material constants: Elastic moduli: E ¼ 170; 000 MPa; m ¼ 0:3; ( isotropic: sr ¼ 1:2; cH ¼ 8; Hardening kinematic: ck ¼ 2; 000 MPa; bk ¼ 0:4; Evolution of normal yield ratio: u ¼ 20; uc ¼ 3; Re ¼ 0:5; Translation of elastic core: ce ¼ 283; v ¼ 0:7; c ¼ 3; n ¼ 2; cm ¼ 10 Viscoplastic deformation: lv ¼ 5; 000; u
454
14 Viscoplastic Constitutive Equations with Subloading Surface Concept
Initial values: Isotrophic hardening function: F0 ¼ 290 MPa: The quite close simulation is shown in Fig. 14.12, while the accuracy of the simulation decreases without the isotropic hardening stagnation since the full inverse loading is repeated mant tymes in this test data. As shown in this section, not only the rate-dependent but also the rate-independent deformation behaviors can be described accurately by the subloading-overstress model and thus the elastoplastic constitutive equation can be disused by adopting the subloading-overstress model.
14.7
Temperature Dependence of Elasto-Viscoplastic Deformation Behavior
The elasto-viscoplastic deformation behavior is influenced by the temperature. However, the dependence of material functions and constants in the elasto-viscoplastic constitutive equation would be complex so that they are specified for each temperature resulting in the piecewise-linear relation in the deformation analysis as seen in the research papers (e.g. Ohno et al. 2018, 2021) and the commercial software (e.g. Marc and Abaqus). Such method for incorporation of the piecewise-linear relation is easy-going but requires the large computational load. Therefore, it is desirable to formulate material functions and constants as functions of temperature. The simple equations of the isotropic hardening function and the Young’s modulus as the functions of temperature are given in this section. The isotropic hardening function FðHÞ would decrease with the elevation of temperature. The following equation taken account of the influence of temperature would be able to be postulated referring to the test data (Poh 2001; Vu and Staat 2007; Cao et al. 2014; Li et al. 2019; Lee et al. 2020) (see Fig. 14.9). h i F ðH; hÞ ¼ F0 f1 þ h1 ½1 expðh2 H Þ g exp Ch ðh hS Þ2
ð14:43Þ
F ðH; hÞ ¼ F0 f1 þ h1 ½1 expðh2 H Þ gsech2 ½Ch ðh hS Þ
ð14:44Þ
or
where h is the temperature (centigrade). hs is the material constant designating the temperature below which FðH; hÞ is highest and constant, while it is lower than the room temperature. Ch is the material constant designating the intensity of the temperature-dependence. Equation (14.43) is the slight modification of the normal (Gaussian) distribution function. Equation (14.44) is invoked from the crystal plasticity (Peierce et al. 1983). It should be noticed that the consideration of the
14.7
Temperature Dependence of Elasto-Viscoplastic Deformation Behavior
455
temperature-dependence at the temperature much higher than hs , i.e. h hS or h hS is required in many engineering cases. The viscoplastic multiplier C in Eq. (14.30) or (14.31) would be extended to the temperature dependence by multiplying the Arrhenius type term (cf. e.g. Conrad 1964; Skrzypek and Hetnarski 1993; Rusinek et al. 2007) i.e. C ! C exp
Q RT
ð14:45Þ
where Q is the activation energy, R is the gas constant and C is the absolute temperature. Besides, the following equation for the temperature dependence of the Young’s modulus E was proposed by Watchman et al. (1961) (see also Rusinek et al. 2007) based on the experimental data.
CT dE CT CT ; exp ¼ B 1 þ E ¼ ET0 BT exp T T T dT
ð14:46Þ
where ET0 is the Young’s modulus at the absolute temperature zero T ¼ 0, B and CT being the material constants, while expðCT =T Þ is the Boltzmann factor. Equation (14.46) leads to the linear relation between E and T, i.e. dE=dT ¼ B for the above-room temperature, i.e. T [ 273ðKÞ as shown in Fig. 14.10. Therefore, the following linear relation would hold for the above room temperature from Eq. (14.46). E ¼ E0 Bh
ð14:47Þ
where E0 is the Young’s modulus at h ¼ 0. On the other hand, the Poisson’s ratio m would not be much affected by the temperature. The strain eh induced by the temperature elevation would be given by eh ¼ aðh h0 ÞI
ð14:48Þ
where a is the material constant. The formulations of the temperature-dependence of material constants other than the isotropic hardening function and the Young’s modulus are also required in the elastoplastic deformation analysis. In particular, the temperature-dependences of the material constants involved in the kinematic hardening and the translation of the elastic-core will have to be formulated in the near future.
Chapter 15
Continuum Damage Model with Subloading Surface Concept
In the brittle materials such as rocks, concrete, glass, and ceramics, the growth of small cracks and their coalescence leading to the macroscopic failure are caused when the stress reaches the yield surface and thereby a plastic deformation is induced. On the other hand, in the ductile materials such as metals, the growth of small voids and their coalescence leading to the failure are caused when a large plastic deformation is induced. These phenomena are called the damage. The phenomenological mechanics describing the damage phenomena within the framework of the continuum mechanics without resorting to a microscopic description is called the continuum damage mechanics. The continuum damage mechanics for the brittle materials has been developed by Kachanov (1958), Rabotnov (1969), Lemaitre and Chaboche (1990), Lemaitre (1992), Lemaitre and Desmoral (2005), de Sauza Neto et al. (2008), Murakami (2012), etc. and the one for the ductile materials has been developed by Gurson (1977), Needleman and Rice (1978) and others. The incorporation of the subloading surface model would be inevitable in the constitutive equation of the damage, because the damage progresses remarkably under the cyclic loading even in the low stress level inside the yield surface and the remarkable softening occurs. The elastoplastic-damage and elasto-viscoplasticdamage models for the brittle and the ductile materials and their extensions by the incorporation of the extended subloading surface model to describe the cyclic loading behavior are explained in this chapter.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_15
457
458
15
15.1
Continuum Damage Model with Subloading Surface Concept
Basic Hypothesis of Strain Equivalence in Constitutive Equation with Brittle Damage
The hypothesis of strain equivalence (Lemaitre 1971) insists “The deformation of the damaged actual material subjected to the actual stress is identical to the deformation of the undamaged material subjected to the effective stress”, i.e. “The strain in the current damaged state is given by the constitutive equation for the undamaged state with the replacement of the stress to the effective stress”. Let this hypothesis be applied also to the elastic and the plastic strains, i.e. ep ðr 6¼ r Þ e ¼ e ; ee ¼ ee; ep ¼
ð15:1Þ
εe = −1: σ = (D)−1: σ (σ, D) % % %
ð15:2Þ
where the variables in the undamaged effective configuration are specified by adding the under tilde ( ). The hypothesis of strain equivalence holds based on the fact that only the microcracks without thickness are induced. Therefore, it holds for the brittle damage but does not hold for the ductile damage in which microvoids possessing the volumes are induced.
15.2
Hyperelastic Equation in Undamaged Variables
The hyperelastic constitutive equation is described by the elastic strain energy function wðee Þ and the Gibbs’ free energy function /ðr Þ noting Eq. (7.74) as follows: r ¼
@wðee Þ ; @ee
ee ¼
@/ðr Þ @r
ð15:3Þ
Now, adopt the quadratic forms of wðee Þ and /ðr Þ as follows: 1 e 1 e e wðe Þ ¼ e : E :e ¼ r :e ; 2 2 e
1 1 e 1 /ðr Þ¼ r :E : r ¼ r :e 2 2
ð15:4Þ
leading to : ee ; r ¼ E
1 ee ¼ E : r
is the elastic modulus tensor in the undamaged state. where E
ð15:5Þ
15.2
Hyperelastic Equation in Undamaged Variables
459
Here, let E be given by the Hooke’s law in Eq. (7.91) under the assumption that the Young’s modulus is influenced but the Poisson’s ratio m is not influenced by the damage, i.e. ; v ¼ v E E
ð15:6Þ
leading to E ⎧ = % (S + ν T ), ⎪E 1 − 2ν % 1 +ν ⎪ ⎨ −1 1 ⎪E = [(1 +ν )S −νT ], E ⎪ ⎩% %
E ijkl = E% [ 1 (δ ik δ jl + δ ilδ jk ) + ν δ ijδ kl ] 1 +ν 2 1 − 2ν % −1 E ijkl = 1 [ 1 (1 + ν )(δ ik δ jl + δ ilδ jk ) −νδ ijδ kl ]
%
ð15:7Þ
E 2 %
Then, the elastic strain energy function wðee Þ and the elastic complementary strain energy function /ðr Þ in the undamaged state are given for Eq. (15.7) referring to Eq. (7.98) as follows: 8 1 E m > e e e 2 e > > < wðe Þ ¼ 2 1 þ m ½eij eij þ 1 2m ðekk Þ 1 > 2 > > : /ðr Þ ¼ 2 ½ð1 þ mÞr ij r ij mðr kk Þ E
ð15:8Þ
The effective stress r and the elastic strain ee are derived from Eq. (15.3) with (15.8) as follows: ⎧ E ∂ψ (ε e ) ν e e e ⎪σ% = ∂ ε e = E: ε = 1 +% ν [ε + 1 − 2ν (tr ε ) I ] % ⎪ ⎪⎛ ⎞ e ⎪⎜ σ ij = ∂ψ (εe ) = E ijkl ε e = E% (ε ij + ν ε kk δ ij ) ⎟ kl 1 +ν 1 − 2ν ∂ε ij ⎟ ⎪⎜ % % ⎪⎝ ⎠ ⎨ σ φ ( ) ∂ 1 ⎪ εe = = E−1 : σ = E [(1+ν ) σ −ν (tr σ)I] ∂σ% ⎪ % % % % % % ⎪ ⎛ ⎞ ⎪ ∂φ (σ ) −1 σ kl = 1 [(1+ν )σ ij −νσ kk δ ij ] ⎟ = E ijkl ⎪⎜ ε ije = % E ∂σ ij ⎟ % % % % ⎪⎩⎜⎝ % ⎠ %
ð15:9Þ
The relation of the elastic strain rate and the undamaged fictitious stress rate is given as follows:
1 ee ¼ E :r ;
r ¼E : ee
ð15:10Þ
460
15
15.3
Continuum Damage Model with Subloading Surface Concept
Hyperelastic Equations with Damage
Let the quadratic free energy functions for the damaged state be given by 8 1 e 1 D e e e e > < w ðe ; DÞ ¼ e : EðDÞ: e ¼ rðe ; DÞ: e 2 2 1 > : /D ðr; DÞ ¼ rðee ; DÞ: E1 ðDÞ: rðee ; DÞ 2 where Dð0 D 1Þ is the damage variable, from which one has 8 > @wD ðee ; DÞ > < rðee ; DÞ ¼ ¼ EðDÞ: ee ; @ee > @/D ðr; DÞ > : ee ¼ ¼ E1 ðDÞ: rðee ; DÞ @r
ð15:11Þ
ð15:12Þ
The relation between the actual stress rðee ; DÞ and the effective stress r is given from Eqs. (15.5) and (15.12) as 1 : ee ¼ E : E1 ðDÞ: rðee ; DÞ; rðee ; DÞ ¼ EðDÞ: E : r r ¼ E
15.3.1
ð15:13Þ
Bilateral Damage
We consider the bilateral damage in which the identical damage is induced in both positive and negative principal stress directions. We adopt the Hooke’s law and it is assumed that the Young’s modulus E decreases by the following equation but the Poisson’s ratio V is not influenced by the damage. ; E ¼ ð1 DÞ E
v ¼ const:
ð15:14Þ
Then, the actual damaged elastic modulus tensor in the Hooke’s law is given by taking account of Eq. (15.14) into Eq. (15.7) as follows: (1 − D ) E ν ⎫ δ δ + 1 (δ δ + δ ilδ jk )] ⎪ [ 1 + ν % 1 − 2ν ij kl 2 ik jl ⎪ ⎬ 1 1 (1 + ν )(δ δ + δ δ ) −νδ δ ]⎪ [ = ij kl ik jl il jk (1 − D)E 2 ⎪ ⎭ %
E ijkl ( D) = (1 − D)E ijkl = %
−1
E ijkl ( D) =
−1 1 E (1 − D) % ijkl
ð15:15Þ
15.3
Hyperelastic Equations with Damage
461
which leads to ; E ¼ EðDÞ=ð1 DÞ EðDÞ ¼ ð1 DÞ E EðDÞ1 ¼ E
1
=ð1 DÞ;
; EðDÞ ¼ D E
E
1
¼ ð1 DÞEðDÞ1
EðDÞ1 ¼
ð15:16Þ ð15:17Þ
D E 1 2 ð1 DÞ
ð15:18Þ
Incidentally, the strain energy function and the complementary energy function are given by 8 1 1 1 > D e e e > > r e w ð e ; D Þ ¼ ð ; D Þ: e : ee ¼ ee : EðDÞ : ee ¼ ð1 DÞee : E > > 2 2 2 > > > > > 1 ð1 DÞ E m > e e > ðtr ee Þ2 ¼ ½e : e þ > > > 2 1þm 1 2m > > > h > > 1 ð1 DÞ E v e 2 i > e e e > w e e ð ; D Þ ¼ e þ e > ij ij kk < 2 1þv 1 2v 1 1 1 > D e e e 1 > rðee ; DÞ: E : rðee ; DÞ / ðr; DÞ ¼ rðe ; DÞ: e ¼ rðe ; DÞ: EðDÞ1 : rðee ; DÞ ¼ > > > 2 2 2ð1 DÞ > > > h i > 1 > > > ¼ ð1 þ vÞrðee ; DÞ: rðee ; DÞ vðtrrðee ; DÞÞ2 > > 2ð1 DÞ E > > ! > > h i > > 1 2 > D e e e > ð1 þ vÞrij ðe ; DÞrij ðe ; DÞ vðrkk ðe ; DÞÞ / ðr; DÞ ¼ > : 2ð1 DÞ E
ð15:19Þ Then, the stress tensor in the current damaged state is given from Eq. (15.13) with Eq. (15.16) as follows: 8 ð1 DÞ E > @wD m > e e e e > tre e r ð e ; D Þ ¼ ¼ ð1 DÞ r ¼ ð1 DÞ E : e ¼ þ ð ÞI ; > e > > 1 2v @e 1þm < r ¼ rðee ; DÞ=ð1 DÞ > > > > > 1 1 > : ee ¼ EðDÞ1 :rðee ; DÞ ¼ E 1 : rðee ; DÞ ¼ ½ð1 þ mÞr mðtrrÞI 1D ð1 DÞE
ð15:20Þ from which one has 8 > D < r ¼ ð1 DÞ r þ Dr ¼ 1 r þ r 1D ð1 DÞ2 ð1 DÞ2 > : r ¼ ð1 DÞ r D r
ð15:21Þ
462
15
Continuum Damage Model with Subloading Surface Concept
8 > > >
> > : r ¼ ð1 DÞ E : e e
ð15:22Þ
D r 1D
noting
ee¼
15.3.2
D E ð1 DÞ2
1
:rþ
1 E 1 : r 1D
ð15:23Þ
Unilateral Damage
The cracks on the planes to which the negative (compressive) principal stresses act would close partly, which is called the crack-closure effect, and this phenomenon is called the unilateral damage. Constitutive equations for the unilateral damage will be given in the following. (1) General tensor formulation with stress transformation tensor The stress tensor is represented as r¼
3 X i¼1
ri nri
nri
¼
3 X 3 X
dij ri nri
nrj
i¼1 j¼1
¼
3 X 3 X
! rij ei ej
ð15:24Þ
i¼1 j¼1
where ri ði ¼ 1; 2; 3Þ are the principal stresses and nri are the unit vector base in the principal stress directions of the stress tensor r. It is postulated that the openings of the cracks are induced on the plane to which the positive principal stresses apply. Then, let the stress tensor be decomposed into the stress tensors with the positive and the negative principal values as follows: r ¼ rþ þ r
ð15:25Þ
8 3 2 0 0 hr1 i > > 3 3 > X > 7 X 6 þ r r > > ¼ r ¼ r n ¼ H ðri Þri nri nri 0 0 r h i h in 5 4 2 i > i i > > > i¼1 i¼1 < 0 0 hr 3 i 3 2 > r 0 0 h i > 1 > 3 3 > X X > 7 6 > > r ¼ 4 0 hr2 i 0 5¼ hri inri nri ¼ H ðri Þri nri nri : > > > : i¼1 j¼1 hr3 i 0 0
ð15:26Þ
15.3
Hyperelastic Equations with Damage
463
where h i is the Macaulay’s bracket and Hð Þ is the Heaviside’s step function, i.e. HðsÞ ¼ 1 for s 0 and HðsÞ ¼ 0 for s\0 (s: arbitrary scalar variable), fulfilling
8 1 for a1 ; a2 ; ; an [ 0 > > ð ÞH ð a Þ H ð a Þ ¼ H a < 1 2 n 0
for others 1 for a1 ; a2 ; ; an \0 > > : H ða1 ÞH ða2 Þ H ðan Þ ¼ 0 for others
ð15:27Þ
Further, introduce the coordinate transformation tensor from the fixed base fei g to the positive principal stress bases Hðri Þnri and the negative principal stress base Hðri Þnri , i.e. 3 8 3 P P > rþ r r > H ð r Þn e ¼ H ð r Þd n e Q < i i i i ij i j i¼1 i¼1 3 3 P P > r > r r : Q H ðri Þni ei ¼ H ðri Þdij ni ej i¼1
ð15:28Þ
i¼1
the components of which are given by 8 3 3 P P > rþ > H ðrr Þnrr er ej ¼ H ðrr Þei nrr djr ¼ H rj ei nrj < Qij ei r¼1
r¼1
3 3 P P > > : Qr H ðrr Þnrr er ej ¼ H ðrr Þei nrr djr ¼ H rj ei nrj ij ei r¼1
r¼1
ð15:29Þ The stress tensor r is transformed to r þ and r by applying the novel positive or negative orthogonal projection tensors Pr þ and Pr , respectively, as follows: r þ ¼ Pr þ : r r ¼ Pr : r
ð15:30Þ
where 8 i 3 P 3 P 3 P 3 h P > rþ > > H ðri ÞH rj H ðrk ÞH ðrl Þdik djl nri nrj nrk nrl
r > > H ðri ÞH rj H ðrk ÞH ðrl Þdik djl nri nrj nrk nrl :P i¼1 j¼1 k¼1 l¼1
ð15:31Þ
464
15
Continuum Damage Model with Subloading Surface Concept
noting Pr þ : r ¼
3 X 3 X 3 X 3 h X
3 i X H ðri ÞH rj H ðrk ÞH ðrl Þdik djl nri nrj nrk nrl : rr nrr nrr
i¼1 j¼1 k¼1 l¼1
¼
3 X
r¼1
H ðrr ÞH ðrr ÞH ðrr ÞH ðrr Þnrr
nrr rr
r¼1
¼
3 X
H ðrr Þrr nrr nrr ¼ r þ
r¼1
Pr : r ¼
3 X 3 X 3 X 3 h X
3 i X H ðri ÞH rj H ðrk ÞH ðrl Þdik djl nri nrj nrk nrl : rr nrr nrr
i¼1 j¼1 k¼1 l¼1
¼
3 X
r¼1
H ðrr ÞH ðrr ÞH ðrr ÞH ðrr Þnrr nrr rr
r¼1
¼
3 X
H ðrr Þrr nrr nrr ¼ r
r¼1
The components of the tensor Pr þ in the base fei g are described as σ+ σ σ ⎧P ijkl = Q irσ +Q σjs+Q kr Q ls ⎪ ⎨ σ− σ− σ− σ σ ⎪⎩P ijkl = Q ir Q js Q krQ ls
ð15:32Þ
noting 3
σ
3
3
3
P ijkl+ = e i ⊗ e j : ∑∑∑∑ H (σ p ) H (σ q ) H (σ r) H (σ s )δ prδ qsn σp ⊗ nσq ⊗ nσr ⊗ n σs e k ⊗ e l p =1q =1 r =1s =1
3
3
= ( ∑ H (σ p )(e i • n σp ) )( ∑ H (σ q )(e j • n σq ))(ek • n σp )(el • n σq ) p =1 σ+ σ+ σ
q =1
σ
= Q i p Q jq Q kpQ lq 3
σ−
3
3
3
P ijkl = e i ⊗ e j : ∑∑∑∑ H (−σ p ) H (−σ q ) H (−σ r) H (−σ s )δ prδ qsn σp ⊗ nσq ⊗ nσr ⊗ n σs e k ⊗ e l p =1q =1 r =1s =1
3
3
= ( ∑ H (−σ p )(e i • n σp ) )( ∑ H (−σ q )(e j • n σq ))(ek • n σp )(el • n σq ) p =1 σ− σ− σ
q =1
σ
= Q i p Q jq Q kpQ lq
On the other hand, the stress tensor r is transformed to the positive stress tensor r þ by the application of the following tensor (Lubarda and Krajcinovic 1993; Lubarda et al. 1994; Wu et al. 2006; Cicekli et al. 2007; Voyiadjis et al. 2008; Abu-Rub 2010; Murakami 2012; Wu and Xu 2013).
15.3
Hyperelastic Equations with Damage
er þ P
3 X
465
H ðri Þnri nri nri nri
ð15:33Þ
i¼1
noting e r þ : r ¼ H ðrr Þnr nr nr nr : rs nr nr P r r r r s S
¼ H ðrr Þrs drs nrr nrr ¼ H ðrr Þrr nrr nrr ¼ r þ
The components of the tensor in Eq. (15.33) are given as 3
σ+ P%ijkl ≡ ei ⊗ e j : ∑ H (σ r )n σr ⊗ n σr ⊗ n σr ⊗ n σr ek ⊗ e l
r =1
3
= ∑ H (σ r )(n σr • ei )⊗ (n σr • e j ) ⊗ (n σr • ek ) ⊗ (n σr • el ) r =1 3
= ∑ H (σ r )QirQ jr Qkr Qlr r =1
However, Eq. (15.33) cannot be applied to the transformation of the elastic modulus tensor as known from er þ er þ : E : P P
¼
3 X
H ðri Þnri nri nri nri :
i¼1
:
3 X
3 X 3 X 3 X 3 X
Eijkl nri nrj nrk nrl
i¼1 j¼1 k¼1 l¼1
H ðrr Þnrr nrr nrr nrr
r¼1 3 X 3 X ¼ H ðr1 ÞE11kl nr1 nr1 þ H ðr2 ÞE22kl nr2 nr2 þ H ðr3 ÞE33kl nr3 nr3 nrk nrl k¼1 l¼1
:ðHðr1 Þnr1 nr1 nr1 nr1 þ Hðr1 Þnr1 nr1 nr1 nr1 þ Hðr1 Þnr1 nr1 nr1 nr1 Þ ¼ ½Hðr1 ÞE1111 nr1 nr1 nr1 nr1 þ Hðr2 ÞE2222 nr2 nr2 nr2 nr2 þ H ðr3 ÞE3333 nr3 nr3 nr3 nr3 ¼
3 X
H ðri ÞEiiii nri nri nri nri
i¼1
which represents the tensor only with the diagonal components.
466
15
Continuum Damage Model with Subloading Surface Concept
Further, consider 3 3 3 3 ⎧ σ+ : P E H (σ i ) H (σ j )E ijkl n σi ⊗ nσj ⊗ nσk ⊗ nσl = ∑∑∑∑ ⎪ % i =1 j =1 k =1l =1 % ⎪ ⎨ 3 3 3 3 ⎪P σ − : E = ∑∑∑∑ H (−σ i ) H (−σ j )E% ijkl n σi ⊗ nσj ⊗ nσk ⊗ nσl ⎪⎩ % i =1 j =1 k =1l =1
ð15:34Þ
8 rþ P :E > > > >
> 3 X 3 X 3 X 3 E X > 1 > v > r r r r > ð ÞH ¼ H r d þ d d d þ r d d > i j ik il il jk ij kl ni nj nk nl < 1þv 2 1 2v i¼1 j¼1 k¼1 l¼1 Pr : E > > > >
> 3 X 3 X 3 X 3 E X > 1 > v > r r r r > ¼ H ð r ÞH r d þ d d d þ d d i j ik il il jk ij kl ni nj nk nl > : 1 þ v i¼1 j¼1 k¼1 l¼1 2 1 2v
ð15:35Þ noting 3
3
3
3
3
3
3
P σ + : E = ∑∑∑∑ [ H (σ i ) H (σ j )H ( σ r ) H (σ s ) δ ir δ j s n σi ⊗ nσj ⊗ nσr ⊗ nσs ] % i =1 j =1 r =1 s =1
:
3
E pqkl n σp ⊗ nσq ⊗ nσk ⊗ nσl ∑∑∑∑ p =1q =1 k =1l =1 %
3
3
3
3
3
3
3
3
3
3
3
3
3
3
3
= ∑∑∑∑ H (σ i ) H (σ j )H ( σ i ) H (σ j )E ijkl n σi ⊗ nσj ⊗ nσk ⊗ nσl % i =1 j =1 k =1l =1 = ∑∑∑∑ H (σ i ) H (σ j )E ijkl n σi ⊗ nσj ⊗ nσk ⊗ nσl % i =1 j =1 k =1l =1 P σ − : E = ∑∑∑∑ [ H (−σ i ) H (−σ j )H ( −σ r ) H (−σ s ) δ irδ j s n σi ⊗ nσj ⊗ nσr ⊗ nσs ] % i =1 j =1 r =1 s =1
: 3
3
3
3
3
3
3
3
3
E pqkl n σp ⊗ nσq ⊗ nσk ⊗ nσl ∑∑∑∑ p =1q =1 k =1l =1 %
= ∑∑∑∑ H (−σ i ) H (−σ j )H ( −σ i ) H (−σ j )E ijkl n σi ⊗ nσj ⊗ nσk ⊗ nσl % i =1 j =1 k =1l =1 = ∑∑∑∑ H (−σ i ) H (−σ j )E ijkl n σi ⊗ nσj ⊗ nσk ⊗ nσl % i =1 j =1 k =1l =1
15.3
Hyperelastic Equations with Damage
467
Then, incorporate the following fictitious undamaged elastic modulus tensor possessing only the components related to the positive or negative principal stress components, respectively. 3 3 3 3 ⎧ σ+ σ+ σ+ ⎪ E ≡ P : E : P =∑∑∑∑ [ H (σ i ) H (σ j )H ( σ k ) H (σ l )E ijkl % % i =1 j =1 k =1l =1 ⎪ ⎪ n σi ⊗ nσj ⊗ nσk ⊗ nσl ] ⎪ ⎨ 3 3 3 3 ⎪ Eσ − ≡ P σ − : E : P σ − = ∑∑∑ [ H (−σ i ) H (−σ j )H ( −σ k ) H (−σ l )E% ijkl ∑ ⎪ % i =1 j =1 k =1l =1 ⎪ n σi ⊗ nσj ⊗ nσk ⊗ nσl ] ⎪ ⎩
ð15:36Þ
8 rþ E Pr þ : E : Pr þ > >
> 3 X 3 X 3 X 3 > E X > 1 v > > H r þ ¼ H ð r ÞH r ð ÞH ð r Þ d þ d d d d d > i j k l ik jl il jk ij kl > > 1 þ v i¼1 j¼1 k¼1 l¼1 2 1 2v > > > > < nri nrj nrk nrl r r r :P > >E P :E >
> 3 X 3 X 3 X 3 E X > r 1 v > >¼ > H ri H rj H ðrk ÞH ðrl Þ dik djl þ dil djk þ dij dkl > > 1þv 2 1 2v > i¼1 j¼1 k¼1 l¼1 > > > : nri nrj nrk nrl
ð15:37Þ ⎧ σ + −1 3 3 3 3 −1 σ = ∑∑∑∑ H (σ i ) H (σ j )H (σ k ) H (σ l )E ijkl n i ⊗ nσj ⊗ nσk ⊗ nσl ⎪E % i =1 j =1 k =1l =1 ⎪ ⎨ 3 3 3 3 −1 σ ⎪ Eσ − −1 = H (−σ i ) H (−σ j )H (−σ k ) H ( −σ l )E ijkl n i ⊗ nσj ⊗ nσk ⊗ nσl ∑∑∑∑ ⎪ % i =1 j =1 k =1l =1 ⎩
ð15:38Þ 8
3 X 3 X 3 X 3 > 1X 1þv > r þ 1 > dik djl þ dil djk vdij dkl nri nrj nrk nrl ¼ H ðri ÞH rj H ðrk ÞH ðrl Þ >E > 2 < E i¼1 j¼1 k¼1 l¼1
3 X 3 X 3 X 3 X > 1 1þv 1 > > dik djl þ dil djk vdij dkl nri nrj nrk nrl H ðri ÞH rj H ðrk ÞH ðrl Þ > Er ¼ > : 2 E i¼1 i¼1 k¼1 l¼1
ð15:39Þ
468
15
Continuum Damage Model with Subloading Surface Concept
noting 3
3
3
3
P σ + : E : P σ + = ∑∑∑∑ H (σ i ) H (σ j )E ijpq n σi ⊗ nσj ⊗ nσp ⊗ nσq % % i =1 j =1 p =1q =1 3
3
3
3
: ∑∑∑∑ [ H (σ r ) H (σ s )H ( σ k ) H (σ l ) δ rk δ sl n σr ⊗ nσs ⊗ nσk ⊗ nσl ] r =1s =1 k =1l =1
3
3
3
3
= ∑∑∑∑ H (σ i ) H (σ j )E ijpq n σi ⊗ nσj δ prδ qs % i =1 j =1 p =1q =1 3
3
3
3
: ∑∑∑∑ [ H (σ r ) H (σ s )H ( σ k ) H (σ l ) δ rk δ sl ⊗ nσk ⊗ nσl ] r =1s =1 k =1l =1
P
σ−:
3
3
3
3
E : P σ − = ∑∑∑∑ H (−σ i ) H (−σ j )E ijpq n σi ⊗ nσj ⊗ nσp ⊗ nσq %
%
i =1 j =1 p =1q =1 3
3
3
3
: ∑∑∑∑ [ H (−σ r ) H (−σ s )H ( −σ k ) H (−σ l ) δ rk δ sl n σr ⊗ nσs ⊗ nσk ⊗ nσl ] r =1s =1 k =1l =1
3
3
3
3
= ∑∑∑∑ H (−σ i ) H (− σ j )E ijpq n σi ⊗ nσj δ prδ qs % i =1 j =1 p =1q =1 3
3
3
3
: ∑∑∑∑ [ H (−σ r ) H (−σ s )H ( −σ k ) H (−σ l ) δ rk δ sl ⊗ nσk ⊗ nσl ] r =1s =1 k =1l =1
It should be noted that Er þ
and ðEr þ Þ
1
possess the components
−1 E ijkl and E ijkl , respectively, while Pr þ possesses the components dij dkl . % %
Now, let the actual elastic modulus tensor for the unilateral damage be given as follows (Fig. 15.1): EðDÞ ¼ EDr E DðEr þ þ hDEr Þ
ð15:40Þ
a
E E
Compression
E
(1 hD)E
1
1
Extension
(1 D) E
0
0
Extension 1
(1 D) E
ae
(1 hD)E
a Compression
(a) Relation of principal actual Young’s modulus vs. principal actual damaged stress.
(b) Relation of principal actual damage stress vs. principal elastic strain
Fig. 15.1 Influence of damage on Young’s modulus illustrated in uniaxial loading (ra : axial stress, eea : elastic axial strain)
15.3
Hyperelastic Equations with Damage
469
h i E Dr ¼ DðEr þ þ hEr Þ þ D E r þ h Er
ð15:41Þ
where hð0 h 1Þ is the material constant, called the crack closure parameter or partial microcrack/void closure, which is h ffi 0:2 in many cases (Lemaitre 1996). The elastic stress-strain relation is given from Eq. (15.12) with Eq. (15.40) as r ¼ EDr : ee ee ¼ EDr1 : r
ð15:42Þ
Then, the stress rate is given by h i r ¼ DðEr þ þ hEr Þ þ D E r þ þ h E r : ee þ EDr : e e
ð15:43Þ
The elastic strain rate is given from Eq. (15.43) as i o 1 n h r þ r þ DðE þ hEr Þ þ D E r þ þ h E r : ee ee¼ EDr
ð15:44Þ
(2) Principal stress formulation The unilateral model in terms of the principal stresses has been formulated by Lemaitre (1996), Lemaitre and Desmoral (2005), de Neto Sauza et al. (2008), etc. It is explained along the formulation and the description by de Sauza Neto et al. (2008) in the following. The stress–strain relation in the uniaxial loading is given in the partial unilateral model as follows: ee ; r ¼ ð1 DÞ E r ¼ ð1 hDÞ E ee ;
9 r for r 0 > = ð1 DÞ E r ee ¼ for r\0 > ; ð1 hDÞ E
ee ¼
ð15:45Þ
The uniaxial stress can be described as r ¼ r þ þ r
ð15:46Þ
where rþ ¼ hri;
r ¼ hri
ð15:47Þ
470
15
Continuum Damage Model with Subloading Surface Concept
The uniaxial stress–strain relation is written as ee ¼
1 rþ r þ 1 D 1 hD E
ð15:48Þ
which is extended to the three-dimensional state in which the signs of principal stress are same as follows: 1 8 > 1 þ v r v r þ r þ r > i 1 2 3 > > for r1 ; r2 ; r3 0A > > 1 D 1 D E < E 1 eei ¼ > > 1 þ v r v r þ r þ r > i 1 2 3 > > for r1 ; r2 ; r3 \0A > : E 1 hD E 1 hD
ð15:49Þ
noting Eq. (15.9)2 . Now, we define the following arrays of principal stresses and strains. 2
3 r1 r 4 r 2 5 ; r3
2
3 ee1 ee 4 ee2 5 ee3
ð15:50Þ
Here, we split the principal stresses as follows: r ¼ rþ þ r
ð15:51Þ
where rþ and r are the tensile and the compressive component, respectively, of r defined by 2
rþ
3 h r1 i 4 hr2 i 5; hr3 i
2
3 hr1 i r 4 hr2 i 5 hr3 i
ð15:52Þ
Further, incorporate the operators PDþ
2 3 0 0 H ðr1 Þ 1 4 0 H ðr2 Þ ¼ 0 5; 1D 0 0 H ðr3 Þ
PD
2 3 0 0 1 4 H ðr 1 Þ ¼ 0 H ðr 2 Þ 0 5 1 hD 0 0 H ðr 3 Þ
ð15:53Þ where HðsÞ is defined by
HðsÞ ¼
0 for s [ 0 1 for s 0
ð15:54Þ
15.3
Hyperelastic Equations with Damage
471
In addition, incorporate the identity tensor with the unilateral damage:
IDþ
8
D D D D T > e > > < e ¼ E P þ þ P E I þ þ I I r ; " #1 > v D T > 1þv D > D D >r ¼ P þ þ P I þ I I ee : E E þ
ð15:60Þ
Note, here, that it would be difficult to derive the time-derivative of Eq. (15.58). Therefore, the stress calculation by the time-integration of rate form constitutive equation based on the forward-Euler method would be difficult and then the return-mapping has to be adopted in the stress integration process by adding the following equation. etrail* n ðr ÞDk ½E ðr Þ1 r ¼ 0
472
15.4
15
Continuum Damage Model with Subloading Surface Concept
Evolution of Damage Variable
The continuum damage variable D is interpreted as an indirect measure of density of micro voids and microcracks (Leckie and Onate 1981) and its evolution rule was given by Lemaitre (1992) as follows:
D ¼
a H ðep ep Þ D p D Y ¼ g e k f 1D
ð15:61Þ
a Y H ðep epD Þ f 1D
ð15:62Þ
where gD
where f and a are the material constants, and epD is the threshold value of the R accumulation of the plastic strain, i.e. ep e p dt. Y is the strain energy release rate which is given in the following.
15.4.1
Bilateral Damage
It follows from Eq. (15.19) that Y¼
@wD ðee ; DÞ 1 e ¼ ee : E :e @D 2
ð15:63Þ
because of 1 1 e wD ¼ r : ee ¼ ee : ð1 DÞ E ð15:64Þ :e 2 2 noting Eq. (15.20). Therefore, the variable Yð¼ ee : E : ee =2Þ in Eq. (15.61)
means the strain energy function in the undamaged effective state. The strain energy release rate Y is rewritten by the current stress r instead of the elastic strain ee as follows: Y ¼
req2 2ð1 DÞ2 E
Rv
ð15:65Þ
where r 2 2 m Rv ð1 þ vÞ þ 3ð1 2vÞ eq r 3
ð15:66Þ
15.4
Evolution of Damage Variable
473
which is called the stress triaxiality function, noting 1 1 1 1 1 1 1 ee : E : ee ¼ ðE : r Þ : E : ðE : r Þ ¼ E : r : r ¼ ðE 1 : rÞ : r 2 2 2 2ð1 DÞ2 1 ¼ ½ð1 þ mÞr : r mðtrrÞ2 2ð1 DÞ2 E ( #)
1 1 1 2 0 0 ð1 þ mÞ r þ ðtrrÞI : r þ ðtrrÞI mðtrrÞ ¼ 3 3 2ð1 DÞ2 E
1 1 ¼ ð1 þ mÞr0 : r0 þ ð1 2mÞðtrrÞ2 2 3 2ð1 DÞ E " !#2 9 !2 8 rffiffiffi rffiffiffi =
< e p ¼ 6 O for K or > : p e ¼ O for other
15.5.2.2
n :E :e [0
ð15:99Þ
Unilateral Damage
The stress versus strain relations are given for the unilateral damage model in the following.
15.5
Elastoplastic-Damage Model with Subloading Surface Model
479
(1) Stress transformation tensor The elastic strain ee is calculated by subtracting the plastic strain ep obtained by the return-mapping method from the strain e and then the stress r is calculated by substituting ee into Eq. (15.42). On the other hand, the plastic strain rate with the explicit expression of the plastic multiplier will be formulated and the calculation by the forward-Euler method will be shown below.
rþ
r
Let E ffi O and E ffi O be postulated, which holds in the state that the influence of the variations of the signs Hðri Þ and the directions nri of principal stresses are negligible. Then, the elastic strain rate for the unilateral damage is given by substituting Eq. (15.61) with Eq. (15.87) into Eq. (15.44) for the abbreviations
rþ
of the term E
and E
r
as follows: " # Dr 1 e D n: r rþ r e r þg þ hE Þ : e e ¼ E pD ðE M
ð15:100Þ
which is rewritten as
n:E
Dr
n: r : e pD n M
!
¼M
pD
n: r n: r þ gD pD n : ðEr þ þ hEr Þ : ee ð15:101Þ pD M M
from which the positive plastic multiplier in terms of the strain rate is described by ! n: r ¼ K pD M M
n : E Dr : e pD þ n : E Dr : n þ g D n : ðEr þ + hEr Þ : ee
ð15:102Þ
Then, the plastic strain rate is given by
ep ¼
n : EDr : e n Dr M pD þ n : E : n þ gD n : ðEr þ þ hEr Þ : ee
ð15:103Þ
where the loading criterion is given as follows (Hashiguchi, 2017): 8 < p e 6¼ O for K [ 0 or n : EDr : e [ 0 : p e ¼ O for other
ð15:104Þ
The actual stress is given from Eqs. (15.42) and (15.103) by Z r¼E
Dr
: e
n : EDr : e dt M
pD
þ n : EDr þ : n þ gD n : ðEr þ þ hEr Þ : ee
! n
ð15:105Þ
480
15
Continuum Damage Model with Subloading Surface Concept
The strain is given from Eq. (8.1) with Eqs. (15.42) and (15.87) by
e ¼ EDr1 : r þ
Z
n : r dt M
pD
n
ð15:106Þ
The calculation must be performed in the coordinate system with the principal þ stress base nri because the tensor EDr is defined in the base nri . Therefore, all the other tensor variables must be first transformed in the base nri as follows, although variables are expressed usually in the fixed base fei g. Tijr ¼ Qrri Trs Qrsj Tr ¼ QrT TQr
ð15:107Þ
noting T ¼ Trs er es ¼ Trs er nri nri es nrj nrj ¼ Trs er nri es nrj nri nrj ¼ Trs Qrri Qrsj nri nrj ¼ Tijr nri nrj Incidentally, note here that the numerical calculation by this model requires the particular caution in the null stress state because EDr is the discontinuous functions of the signs of the principal stresses for r ¼ O.
It is unnecessary to use the explicit expression of the plastic multiplier K in Eq. (15.102), if it is calculated by the return-mapping method. The stress is calculated from hyperelastic equation with the substitution of the elastic strain ee which is calculated by subtracting the plastic strain ep from the total strain e. (2) Principal stress formulation
It is difficult to formulate the explicit equation of the positive plastic multiplier k or K in the principal stress formulation described in Sect. 15.3.2(2). Therefore, the analysis in the forward-Euler method is impossible and thus the return mapping method must be used as described in the beginning of (1) Stress transformation tensor. In this formulation, note also that the numerical calculation requires the particular caution in the null stress state because PDþ and PD are the discontinuous functions of the signs of the principal stresses for r ¼ O. It should be noticed that the explicit expression of the elatoplastic-damage constitutive relation is of importance for the physical interpretation of the constitutive equation. In this context, the constitutive equation in (1) stress transformation tensor is more beneficial than the one in (2) principal stress formulation. Needless to say, the explicit forward-calculation can be performed only by (1) stress transformation tensor.
15.6
Anisotropic (Orthotropic) Damage Tensor
15.6
481
Anisotropic (Orthotropic) Damage Tensor
The following asymmetric effective stress was proposed by Murakami and Ohno (1981) and Murakami (1988). r ¼ ðI DÞ1 r
ð15:108Þ
extending the relation r ¼ r=ð1 DÞ in the one-dimension with the second-order symmetric damage tensor D. However, an asymmetric stress tensor makes the mathematical formulation and mechanical analysis very complicated. Then, various symmetrized effective stress tensors have been proposed. For instance, r ¼ ½rðI DÞ1 þ ðI DÞ1 r=2
ð15:109Þ
r ¼ ðI DÞ1 rðI DÞ1
ð15:110Þ
and
have been proposed by Murakami and Ohno (1981) and Betton (1986), respectively. On the other hand, Cordebois and Sidoroff (1982a, b) proposed the effective stress tensor r ¼ ðIDÞ1=2 rðI DÞ1=2
ð15:111Þ
Further, extending Eq. (15.19) to the anisotropic damage, Lemaitre et al. (2000) assumed the following Gibbs energy. /D ¼
1þv 3ð1 2vÞ r2m trðHr0 Hr0 Þ þ 1 gDm 2E 2E
ð15:112Þ
where 1 H ðIDÞ1=2 ð¼ HT Þ; Dm trD 3
ð15:113Þ
g is the hydrostatic sensitivity parameter concerning the variation of the Poisson’s ratio with damage, while g ffi 3 is used most often. The particular case chosen as D¼DI with the scalar-valued damage variable D and g ¼ 1 corresponds to the isotropic damage.
482
15
Continuum Damage Model with Subloading Surface Concept
The elastic strain is derived from Eq. (15.112) as follows: ee ¼
@/D 1 þ v 3v ¼ r r @r E E
mI
ð15:114Þ
which is of identical form to the Hooke’s law but the effective stress is related to the actual stress as follows (Lemaitre et al., 2000): 0
r ðHr0 HÞ þ
rm I 1 gDm
ð15:115Þ
by comparing eeij
@/D @ 1þv 3ð1 2vÞ r2m 0 0 Hpq rqr Hrs rsp þ ¼ ¼ @rij @rij 2E 1 gDm 2E
0 0 @r @r 1þv 3ð1 2vÞ 2rm 1 qr sp þ Hpq dij Hrs r0sp þ Hpq r0qr Hrs ¼ @rij @rij 1 gDm 3 2E 2E 0 1 2v r 1þv m ¼ dij Hpi r0sp Hjs þ E E 1 gmm
0 1þv rm v rm ¼ dij 3dij Hpi r0sp Hjs þ 1 gDm E E 1 gDm 2 3
0 0 7 1þv rm v 6 rm 0 Hpi r0sp Hjs þ H ¼ dij 6 r H þ daa 7 p js i sp 5dij E E 4|fflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflaa 1 gDm 1 gD m ffl} 0
derived from Eq. (15.112) with Eq. (15.114). Equation (15.115) is rewritten as follows: r ¼ MðDÞ : r
ð15:116Þ
r ¼ MðDÞ : r0 þ rm MðDÞ : I
ð15:117Þ
i.e.
where ~ MðDÞ ¼ HH
1 1 2 II ð15:118Þ H I þ I H2 þ trH2 I I þ 3 9 3ð1 gDm Þ
15.6
Anisotropic (Orthotropic) Damage Tensor
483
~ ijkl Hik Hij due to Eq. (1.196)4 , noting with the notation ðHHÞ 0 0 ðHr0 HÞij ¼ Hik r0kl Hlj ¼ Hik r0kl Hlj Hrk r0kl Hlr dij =3 ¼ Hik ðrkl rm dkl ÞHlj Hrk rkl rm dkj Hlr dij =3 ¼ Hik Hlj Hir Hrj dkl þ dij Hlr Hrk =3 þ Hrk Hkr dij dkl =9 rkl ~ H2 I þ I H2 =3 þ trH2 I I=9 r ¼ HH ij
1 2 1 II I ~ H I þ I H2 þ trH2 I I þ :I ¼ MðDÞ : I ¼ HH 3 9 3ð1 gDm Þ 1 gDm ~ : I ¼ H2 ; H2 I : I ¼ 3H2 ; I H2 : I ¼ trH2 I HH
Equation (15.116) is inverted as follows (Mengoni and Ponthot 2015): r 0 : H2 0 r ¼ H1 r H1 2 H2 þ ð1 gDm Þr m I trH
ð15:119Þ
r ¼ M1 ðDÞ : r
ð15:120Þ
H2 H2 1 þ ð1 gDm ÞI I 3 trH2
ð15:121Þ
where M1 ðDÞ ¼ H1 H1
The rate of the damage tensor is given by m Y p D¼ e S
ð15:122Þ
where S and m are material parameters and e p is defined by X 3 p p P e eP n nP
ð15:123Þ
n¼1
ePp and nP are the principal values and the normalized principal direction vectors of plastic strain rate. Y is given by " 2 # q2eff 2 peff ð1 þ vÞ þ 3ð1 2vÞ Y ¼ 2E qeff 3
ð15:124Þ
484
15
Continuum Damage Model with Subloading Surface Concept
where qeff
rffiffiffi 3 rm 0 ðH^ r0 HÞ ; peff 1 gðtrDÞ=3 2
ð15:125Þ
The yield condition for the abovementioned anisotropic damage model based on the undamaged stress r in Eq. (15.115) is given by rffiffiffi 3 ðHr0 HÞ0 FðHÞ ¼ 0 2
ð15:126Þ
The plastic strain rate is given by
ep ¼ kn
ð15:127Þ
0 HðHr0 HÞ0 H n ¼ 0 0 HðHr0 HÞ H
ð15:128Þ
where
The derivation of this equation by Prof. Yuki Yamakawa (Tohoku Univ.) is shown in Appendix I. The unilateral formulation for the anisotropic damage has been given by Ladeveze (1983) and Desmorat (1999) as described in Lemaitre and Desmorat (2005).
15.7
Subloading-Overstress Damage Model
The subloading-damage model formulated in the preceding sections will be extended to the rate-dependent equation, i.e. the subloading-overstress damage model by replacing the plastic strain rate e p to the viscoplastic strain rate e vp in Eq. (14.29) in all the equations formulated in Sect. 15.5. The rates of the isotropic hardening variable is given by Eq. (14.36) itself, and the rates of the damage variable in Eq. (15.61), the kinematic hardening in Eq. (15.82) and the elastic-core in Eq. (15.84) are extended as follows:
H ¼ Cf Hn
ð15:129Þ
15.7
Subloading-Overstress Damage Model
485
D ¼ CgD
ð15:130Þ
a ¼ Cf kn
D
ð15:131Þ
c ¼ Cf D cn
ð15:132Þ
The accelerated creep can be described in addition to the steady-state creep after the transient creep by the present subloading-overstress damage model.
15.7.1
Bilateral Damage
Combining the bilateral damage elastic strain rate in (15.22) and the subloading-overstress model in Eq. (14.29), one has 1 E e¼ 1D
15.7.2
1
CgD r þ Cn : rþ 1D
ð15:133Þ
Unilateral Damage
The stress versus strain relations are given below. (a) Stress transformation tensor Combining the unilateral damage elastic strain rate in (15.44) and the subloading-overstress model in Eq. (14.29) with Eq. (15.130), one has 1
e ¼ ðEDr þ Þ fr þ ½CgD ðEr þ þ hEr Þ þ DðEr þ þ hE Þ:ee g þ Cn ð15:134Þ r
(b) Principal stress formulation Combining the unilateral damage elastic strain rate in (15.60) and the subloading-overstress model in Eq. (14.29), one has "
# Z v D 1þv D D D T P þ þ P I þ þ I I r þ Cn dt e ¼ E E
ð15:135Þ
where n is the outward-normal vector represented in the principal direction of the stress tensor. The creep behavior (under r ¼ const:Þ described by the viscoplastic-damage model is illustrated in Fig. 15.2. The transient and the steady-state creep behaviors are described by the viscoplastic model and the accelerating creep behavior is
486
15
Continuum Damage Model with Subloading Surface Concept
Accelerating Steady-state creep creep (D 1, Transient (D 0, ) F 0 F (1 D) F 0 , creep R cm Rs ) (D=0)
t Fig. 15.2 Creep behavior described by viscoplastic-damage model
described by the viscoplastic-damage model. The viscoplastic strain rate is induced infinitely because the scalar variable C increases as the damage variable approaches unity, i.e. D ! 1 leading to F ! ð1 DÞF ! 0 and R ! cm Rs .
15.8
Subloading-Gurson Model for Ductile Damage
Consider the ductile damage caused by the growth of the voids (or porous media). The plastic deformation with the ductile damage is influenced by the hydrostatic stress even if the base material is the Mises material, the plastic deformation behavior of which is independent of hydrostatic stress. The elastoplastic constitutive model taking account of the nucleation and the growth of spherical voids was proposed first by Gurson (1977) and further studied by Needleman and Rice (1978), Tvergaard and Needleman (1984), Needleman and Tvergaard (1985), etc. The yield surface is introduced, which takes account of the void volume fraction and the mean stress with the evolution rule of the void volume fraction. It is often called the Gurson model and its elaboration taken account of the void coalescence is called the GTN (Gurson-Tvergaard-Needleman) model. The subloading-void(Gurson) model will be described in this section. The following yield condition is derived by Gurson (1977) based on the symmetric deformation analysis of the rigid-plastic Mises material containing a spherical cavity. fn ðr; F; nÞ ¼
req F
2
3 rm n2 1 ¼ 0 þ 2n cosh 2F
ð15:136Þ
where n is the void volume fraction. Equation (15.136) is reduced to the von Mises yield condition, i.e. req ¼ F for n ¼ 0. The dependence of the yield function on the pressure in Eq. (15.136) is shown in Fig. 15.3.
15.8
Subloading-Gurson Model for Ductile Damage
487 0.0
1.0
0.01
0.8 1 0.6 eq / F
0.05 0.1
0.4
0.3
0.2
0.0
0
1
2
m / F
3
4
Fig. 15.3 Effect of void volume fraction in Gurson yield surface
The rate of the void volume fraction n is given by sum of the growth rate ngrow
and the nucleation rate of new void nnucl as follows (Needleman and Rice, 1978):
n ¼ ngrow þ nmucl
ð15:137Þ
where 9 p = ngrow ¼ ð1 nÞtre p epq nmucl ¼ a1 F þ rm þ a2 e ;
ð15:138Þ
The coefficients a1 and a2 are given by Chu and Needleman (1980) as follows: " #9 fn 1 F þ rm rn 2 > > > a1 ¼ pffiffiffiffiffiffi exp > > > Sn 2 2pSn = "
fn 1 e en a2 ¼ pffiffiffiffiffiffi exp Sn 2 2pSn eqp
2 #
> > > > > > ;
ð15:139Þ
which is derived postulating that the voids nucleates according to the probability distribution with the stress rn and the strain en as their mean values together with sn as their standard deviation, and fn is the volume fraction of void nucleating particles. The subloading surface for the normal-yield surface in Eq. (15.136) is given by replacing F to RF in Eq. (15.136) as follows: fn ðr; F; nÞ ¼
eq 2 r 3 rm n2 1 ¼ 0 þ 2n cosh RF 2 RF
ð15:140Þ
488
15
Continuum Damage Model with Subloading Surface Concept
The time-differentiation of Eq. (15.140) is given by
eq req r RF req ðR F þ R FÞ RF R2 F 2 3 rm 3 rm RF rm ðR F þ R FÞ þ 2n sinh 2n ¼ 0 2 2 2 RF 2 RF
f n ðr; F; nÞ ¼ 2
ð15:141Þ
from which one has " !# !# " req eq 3 rm R F R eq F r r þ rm rm þ þ 3n sinh 2RF n ¼ 0 2 RF F R F R 2 RF ð15:142Þ Assume the associated flow rule
p
e ¼ k nn
ðk [ 0Þ
ð15:143Þ
where @fn @fn = n @r @r n
n n ¼ 1
ð15:144Þ
It follows by substituting Eqs. (9.9), (11.12), (15.137) and (15.143) into Eq. (15.142) that 2 0 2 0 13 13 0 w 0 w req 4 eq U kA5 3 rm 4 F k h U k eq @F k h A5 2 r r þ þ 3n sinh r m rm @ þ RF F R F R 2 RF " rffiffiffi # 2 w 0 w 2RF ð1 nÞ k trn þ a1 F k h þ rm þ a2 k ¼0 3
resulting in
req eq 3 rm r þ 3n sinh 2a1 RF rm 2 RF 2 RF eq 0
r 3 rm F w U eq þ 3nrm sinh h þ 2r RF F 2 RF R " #) rffiffiffi 2 þ 2RF ð1 nÞtrnw þ a1 F 0 hw þ a2 k ¼ 0 3
ð15:145Þ
15.8
Subloading-Gurson Model for Ductile Damage
489
where
hw H = kð¼
pffiffiffiffiffiffiffiffi 2=3Þ
ð15:146Þ
Noting 1 rm ¼ I : r; 3
r
eq
! rffiffiffi rffiffiffi 3 0 3 r0 : r 3 r0 : r ¼ kr k ¼ 2 2 kr0 k 2 req
¼
ð15:147Þ
one has
req eq 3 rm r þ 3n sinh 2a1 RF rm 2 RF 2 RF
0
r 3 rm 2 þ n sinh a1 RF I : r ¼ 3 RF 2 RF 3
ð15:148Þ
Substituting Eq. (15.148) into Eq. (15.145), the plastic multiplier is derived as follows:
k¼
tSG : r M
SG
ð15:149Þ
where SG
t
M
SG
r0 3 rm 2 þ n sinh a1 RF I 3 RF 2 RF 3
eq 0 r 3 rm F w U þ 3nrm sinh h þ 2req RF F 2 RF R " rffiffiffi # 2 þ 2RF ð1 nÞtrnw þ a1 F 0 hw þ a2 3
ð15:150Þ
ð15:151Þ
The strain rate for Eqs. (15.143) and (15.149) is given by
e ¼ E1 : r þ
tSG : r M
SG
nw ¼
nw tSG :r E1 þ SG M
ð15:152Þ
490
15
Continuum Damage Model with Subloading Surface Concept
from which the plastic multiplier in terms of strain rate is derived as follows:
tSG : E : e
K¼ M
SG
ð15:153Þ
þ tSG : E : nw
The stress rate for Eq. (15.153) is given by
tSG : E : e
r ¼ E : e
M
SG
þt
SG
:E:
E:n ¼ E w
nw
E : nw ðtSG : EÞ
M
SG
þt
SG
:E:
nw
:e
ð15:154Þ The loading criterion is given by (
ep¼ 6 O for K or tSG : E : e [ 0 p e ¼ O for other
ð15:155Þ
The elaboration of the Gurson model was proposed by Tvergaard (1982) (see also Tvergaard and Needleman, 1984) by introducing the void coalescence into the yield condition in Eq. (15.136). It is further extended by the concept of the subloading surface as follows: wðr; F; nÞ ¼
req 3 q2 r m þ 2n q1 cosh q3 n2 1 ¼ 0 RF 2 RF
ð15:156Þ
where n ðnÞ is the extension of the void volume fraction n so as to represent the loss of the load-carrying capacity due to the void coalescence, i.e. 8 for n nc
< ni; j ¼ 1 expðgi ðh; j Þðti; j þ DtÞ 1=ni ðh;j Þ ð16:22Þ 1 1 > : ti; j ¼ ½ð In Þ gi ðh;j Þ 1 ni;j1 for the isothermal phase transformation, where “,j” denotes the value at the time tj and ni;j1 is the known value at the time tj1 . Designating the times as ta and tb when the volume fraction becomes nia and nib , respectively, we have ( gi ðhÞtani ðhÞ ¼ lnð1nia Þ gi ðhÞtbni ðhÞ ¼ lnð1nib Þ from which gi ðhÞ and ni ðhÞ are calculated by 8 1 1 > > > gi ðhÞ ¼ ni ðhÞ ln 1 n > > ia ta > < lnð1nia Þ log > lnð1 nib Þ > > ni ðhÞ ¼ > > ta > : log tb The base phase is the austenite in many cases.
ð16:23Þ
498
16.3
16
Subloading Phase-Transformation Model
Stress Rate Versus Strain Rate Relation
Substituting Eqs. (16.2), (16.17), (16.18) and (16.19) into Eq. (16.1), we have the constitutive equation as follows: @Fi n:r dh ni þ Fi ni n: dr 1 @h F @E i¼1 dh: r þ de ¼ E1 : dr þ n p @h M ð16:24Þ ! 1=3 n n X X 1 0q @q @q 0 dn I þ 3Ki ð1 XÞdni r dh þ 3q q @h @ni i i¼1 i¼2 n P
It follows from Eq. (16.24) that @Fi n:r n: dr i¼1 dh ni þ Fi ni @E1 F @h n E: de ¼ n: dr þ n: E: n: E: n dh: r þ p @h M ! 0 1=3 n n X X 1 q @q @q dh þ dni n: E: I þ 3Ki ð1XÞdni n: E: r0 3q q @h @ni i¼1 i¼2 Pn
from which the magnitude of plastic strain increment, denoted by dK instead of dk, is derived as follows: " n X 1 @Fi n: r dh ni þ Fi ni n: E: de dK ¼ p @h F M þ n: E: n i¼1 ! 1=3 n X @E1 1 0q @q @q dh: r þ dh þ dn n: E: I ð16:25Þ n: E: @h 3q q @h @ni i i¼1 # n X 0 3Ki ð1 XÞdni n: E: r i¼2
It follows from Eqs. (16.1) and (16.2) that dr ¼ E:
@E1 dh: r dep deh deTp de @h
ð16:26Þ
which leads to the following equation by substituting Eqs. (16.12), (16.18), (16.20) and (16.24) into Eq. (16.26).
16.3
Stress Rate Versus Strain Rate Relation
dr ¼ E:
de
499
@E1 dh: r @h "
n X 1 @Fi n: r dh ni þ Fi ni p n: E: de @h F M þ n: E: n i¼1 ! 1=3 n X @E1 1 0q @q @q dh: r þ dh þ dn n: E: I ð16:27Þ n: E: @h 3q q @h @ni i i¼1 # n X 3Ki ð1 XÞdni n: E: r0 n i¼2
1 þ 3q
! ) 0 1=3 n n X X q @q @q 0 dn I 3Ki ð1 XÞdni r dh þ q @h @ni i i¼1 i¼2
The loading criterion for the plastic strain increment is given by dep ¼ 6 O dep ¼ O
for dK [ 0 for other
ð16:28Þ
The subloading phase-transformation model was formulated within the framework of the initial subloading surface model in this chapter. It can be generalized to the extended subloading surface model in Chap. 11 which incorporates the evolution rule of the elastic-core, i.e. the similarity-center of the subloading surface and the normal-yield surface so that the cyclic loading behavior can be also described appropriately. Further, it can be extended to the finite strain theory based on the multiplicative decomposition in Chap. 17.
Chapter 17
Multiplicative Hyperelastic-Based Plasticity with Subloading Surface Concept
Elastoplastic constitutive equation has been formulated first as the infinitesimal elastoplasticity and then developed as the hypoelastic-based plasticity. However, they are incapable of describing the finite deformation/rotation exactly as was described in Sect. 8.2. The multiplicative hyperelastic-based plasticity is capable of describing the elastoplastic deformation and rotation accurately, in which the deformation gradient tensor is decomposed definitely into the purely elastic (hyperelastic) part and the purely plastic part by the multiplicative decomposition. Therefore, it has been studied widely by many workers, e.g., Lion (2000), Wallin et al. (2003), Dettmer and Reese (2004), Wallin and Ristinmaa (2005), Vladimirov et al. (2008, 2010), Hashiguchi and Yamakawa (2012), Hashiguchi (2013, 2016, 2017), etc. after the advocacy of the multiplicative decomposition by Kroner (1960), Lee and Liu (1967), Lee (1969) and Mandel (1965, 1971, 1973, 1974). The multiplicative decomposition of the deformation gradient tensor with the isoclinic concept (Mandel 1971) leads to the inclusion of the rigid-body rotation in the elastic deformation gradient tensor and thus leads to the description of constitutive relation in the intermediate configuration unloaded to the stress-free state along the hyperelasticity. Here, it should be emphasized that the multiplicative decomposition with the definite uniqueness holds in the general elastoplastic materials, although the isoclinic concept is often misunderstood to hold only in the crystalline materials. The multiplicative hyper-elastic-based plastic constitutive equation will be explained comprehensively in this chapter.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_17
501
502
17.1
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
Exact Elastic–Plastic Decomposition of Deformation Measure
The deformation measure must be decomposed into the elastic and the plastic parts definitely for the exact description of elastoplastic deformation as was explained in Sect. 8.1. In addition, any deformation (rate) measures are defined by the deformation gradient tensor as described in Chap. 4: Deformation/rotation tensors. These requirements can be realized by the multiplicative decomposition of the deformation gradient tensor up to the finite deformation. In addition, the multiplicative hyperelastic-based plasticity can be formulated rigorously and concisely in the intermediate configuration based on the isoclinic concept by which the intermediate configuration is not influenced by the rotation of the substructure of material. The physical background and the rigorous formulation of the multiplicative decomposition will be described in this section.
17.1.1
Necessity of Multiplicative Decomposition of Deformation Gradient Tensor
First of all, we must remind (1) the deformation must be decomposed into the purely elastic deformation and the purely plastic deformation as was described in Sect. 8.1.1 and (2) any deformation (strain) measures are defined by the deformation gradient tensor F ¼ @x=@X, where X and x are the position vectors of the material particle in the reference and the current configurations, respectively. Therefore, the decomposition of the deformation gradient tensor into the purely elastic part and the purely plastic part is required to formulate the exact elastoplastic constitutive equation. Consider the deformation gradient tensor F which is the most basic variable by which all the deformation measures are defined. It is defined by the ratio of the current infinitesimal line-element vector to the reference infinitesimal line-element vector and thus it is multiplicatively decomposed exactly as follows: F¼
@x @x @X ¼ @X @X @X
ð17:1Þ
where X is the fictitious position vector of the material particle in the stress-free state unloaded virtually by the hyperelastic constitutive equation, while the plastic deformation is induced in addition to the elastic deformation in the actual unloading process. Then, it follows that F ¼ Fe Fp
ð17:2Þ
17.1
Exact Elastic–Plastic Decomposition of Deformation Measure
503
by setting Fe ¼
@x ; @X
Fp ¼
@X @X
ð17:3Þ
i.e. dx ¼ FdX ¼ Fe dX; dX ¼ Fp dX
ð17:4Þ
Equation (17.2) has been proposed by Eckert (1948), Kroner (1960), Lee and Liu (1967), Lee (1969) and has been studied by Mandel (1971,1972, 1973), Kratochvil (1971) and called the multiplicative decomposition (or Kroner decomposition or Lee decomposition) of the deformation gradient tensor. The unloaded configuration to the stress-free state is called the intermediate (or stress-free or relaxed) configuration. The elastoplastic deformation process based on this notion is illustrated in Fig. 17.1 where the initial (reference), the intermediate and the current configurations are specified by the symbols K0 , K and K, respectively. The tensors in the initial configuration K0 are designated by the uppercase letters as T, the ones in the intermediate configuration K by the uppercase letters with the upper bar as T and the ones in the current configuration K by the lowercase letters as t. The notion of the multiplicative decomposition is illustrated in the uniaxial loading process in Fig. 17.2, replacing the infinitesimal line-element vectors dX, dx and dX by l0 , l and l, respectively. Here, note that a slight plastic deformation in the opposite direction to the first loading process is induced actually in conjunction with the elastic unloading of material particles (Material particles which are elongated to ellipsoidal shape in the first loading process are returned to spherical shape in the unloading process which causes the mutual slips of material particles leading to the slight plastic deformation), while the purely elastic deformation is induced F
x
dx
X dX
K0
Fp
XdXdX X
Fe
K
Fig. 17.1 Multiplicative decomposition of deformation gradient tensor
K
504
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
Stress
Actual unloading process
Elastic strain energy Virtual unloading process along hyperelasticity
0
l0
l˜
l
l
Length
Fig. 17.2 Unloading process in uniaxial deformation of a bar
only at the initiation of unloading process. Therefore, the length of real material reduces not to l but to ~l, where the difference ~l l is induced by the plastic deformation in the opposite direction to the preceding loading. Here, note that the micromechanical structure of plastically deformed material is heterogeneous, possessing the statically-indeterminate structure, and thus different amounts of distressing are required in order that all material points reach stress-free states. Needless to say, it is impossible to separate the material into infinitesimal pieces and then gather them as ever. Consequently, X and l are not actual but merely virtual quantities based on the hyperelasticity. The difference ~l l must be described by the plastic constitutive equation, i.e. the cyclic plasticity model represented by the extended subloading surface model in Chap. 11 by which the closed hysteresis loop can be depicted. In other words, the deviation of the actual unloading curve from the hyperelastic curve is the physical background for the appearance of the closed hysteresis loop in the actual stress–strain curve. It can never be described by the Chaboche model, the two surface model, etc. which describe only the elastic deformation in the unloading process but it can be described by the extended subloading surface model in which the elastic-core moves with the plastic deformation.
17.1.2
Embedded Base Vectors in Intermediate Configuration
It follows from Eq. (3.9) that 8 i i j j i > F ¼ g G ¼ d g G ¼ F g G ; F1 ¼ Gi gi > i j i j i > < 1 i j j Fe ¼ gi G ¼ dij gi G ¼ Feij gi G ; Fe ¼ Gi gi > > > : Fp ¼ G Gi ¼ di G G j ¼ F pi G G j ; Fp1 ¼ G Gi i i i j j i
ð17:5Þ
17.1
Exact Elastic–Plastic Decomposition of Deformation Measure
505
noting i i gi Gi ¼ gi G Gj G j G Gj ¼ dij
ð17:6Þ
with the metric tensors 8 i i T < G Gi G ¼ G Gi ¼ G i i T G Þ : G Gi iG ¼i G Gi ð¼ g gi g ¼ g g i ð ¼ g T Þ
ð17:7Þ
i
where Gi and G are the primary and the reciprocal base vectors in the intermediate configuration, which are given by 8 < Fe1 gi ð¼ e
8 < FeT gi ð¼ e g G Þ i G ¼ : pT i p !G F G ð¼ G Þ
g GÞ ; Gi ¼ : Fp Gi ð¼ p ! Þ GG
ð17:8Þ
noting Eqs. (3.17) and (3.21).
17.2
Deformation Tensors
The multiplicative hyperelastic-based plasticity can be formulated essentially in the intermediate configuration as will be explained in the later sections. Therein, various deformation (rate) measures in the current, the intermediate and the reference configurations will be used in the formulation. They are explained in-brief in this section.
17.2.1
Elastic and Plastic Right Cauchy-Green Deformation Tensor
Let the following elastic and plastic Cauchy-Green deformation tensors be defined noting Eqs. (3.24), (3.28) and (4.16). ð17:9Þ
C ¼ FT F ¼ U2 ¼ Gi gi gj G j ¼ gij Gi G j e
e2
C F F ¼U ¼ eT
e
8
l FF1 > > < e l Fe F e1 1 p 1 > lp Fe Fp Fp1 Fe ¼ Fe L Fe > > : p p p1 L F F
ð17:13Þ
17.2
Deformation Tensors
507
F F1 ¼ ðFe Fp Þ ðFe Fp Þ1 ¼ Fe Fp þ Fe Fp Fp1 Fe1
¼ Fe Fe1 þ Fe Fp Fp1 Fe1 The strain rates and spins and their elastic and the plastic parts in the current configuration are defined based on Eq. (17.12) with Eq. (17.13) as d ¼ de þ dp ð17:14Þ w ¼ we þ wp h i h i 8 < d ¼ sym½l ¼ sym F F1 ¼ R sym U U1 RT h i h i : w ¼ ant½l ¼ ant F F1 ¼ R RT þ R ant U U1 RT h i h i 8 < de ¼ sym½le ¼ sym Fe Fe1 ¼ Re sym Ue Ue1 ReT h i h i : we ¼ ant½le ¼ ant F e Fe1 ¼ R e ReT þ Re ant U e Ue1 ReT
p dp ¼ sym½lp ¼ sym Fe L Fe1 p wp ¼ ant½lp ¼ ant Fe L Fe1
ð17:15Þ
ð17:16Þ
ð17:17Þ
T noting Re TReT ¼ Re TT ReT . It should be noticed that the velocity gradient tensor in the current configuration is not decomposed into the purely elastic part described only by the elastic deformation gradient tensor and the purely plastic part described only by the plastic deformation gradient tensor. (2) Strain rate and spin tensors in intermediate configuration The following velocity gradient tensor based on the contravariant-covariant pull-back of the velocity gradient tensor l by Fe to the intermediate configuration is additively decomposed into the purely elastic part and the purely plastic part, noting Eq. (17.12) with Eq. (17.13). e
where
p
L ¼ L þL
ð17:18Þ
8 e1 e e G > > < L F lF ð¼ l G Þ G e L Fe1 le Fe ð¼ e le G Þ ¼ Fe1 Fe > > : p G L Fe1 lp Fe ð¼ e lp G Þ ¼ Fp Fp1
ð17:19Þ
e
p
Therefore, L with L and L can be pertinently adopted in the formulation of elastoplastic constitutive equation. Let them be decomposed additively into the symmetric and the antisymmetric parts, i.e.
508
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
L ¼ DþW e e e p p p L ¼ D þW ;L ¼ D þW
ð17:20Þ
with e
p
e
D ¼ D þD ; where
W ¼ W þW
p
e p p e D sym½L; D sym L ; D sym L e p e p W ant½L; W ant L ; W ant L
ð17:21Þ
ð17:22Þ
e
The rate of C is given from Eqs. (17.9) and (17.18) with Eq. (17.19)2 as e e e e p C ¼ 2sym C L ¼ 2sym C L L
ð17:23Þ
noting e C ¼ FeT Fe ¼ FeT Fe þ FeT Fe ¼ FeT Fe Fe1 Fe þ FeT FeT FeT Fe e e T e e e e eT e ¼ C L þL C ¼ C L þ C L e The Mandel stress tensor M ¼ C S (Mandel, 1973b) in the intermediate configuration will be used as the stress measure, because it is work-conjugate to the velocity gradient tensor L in the intermediate configuration as will be delineated in Sect. 17.6. (3) Substructure spin The substructure of material is to be its skeleton and thus the rotation of the substructure is induced by the elastic distortion and the rigid body rotation. On the other hand, the plastic deformation is induced by the mutual slips between material particles along the substructure and thus it does not influence the rotation of the substructure. This notion has been suggested by Mandel (1971), Kratochvil (1971), Dafalias (1983, 1984, 1985), Loret (1983), etc. Eventually, the rotation of substructure is independent of the plastic deformation but induced by the elastic distortion and the rigid-body rotation both of which are included in the elastic deformation gradient Fe . Then, the S
spin of substructure W is given by the elastic spin, i.e. S
e
W ¼W ¼WW
p
ð17:24Þ
The continuum spin W is known from the outside appearance of material. In S
contrast, the spin of W consisting of the elastic distortion spin and the rigid-body rotation cannot be captured from the outside appearance of material as far as the plastic spin is unknown, while the plastic spin depends on the plastic property on the occurrence of the mutual slips between material particles. Thus, it is obliged to formulate the plastic spin as the constitutive equation.
17.2
Deformation Tensors
509
Then, the substructure spin can be calculated by subtracting the plastic spin from the continuum spin.
17.3
On Limitation of Hypoelastic-Based Plasticity
It was delineated that the hypoelastic equation is incapable of describing the constitutive equation of the elastic deformation exactly in Sect. 7.7. Further, let the fact the hypoelastic-based plasticity is incapable of describing the constitutive equation of the elastoplastic deformation exactly be examined from the view point of the multiplicative decomposition required for the exact description of the elastoplastic constitutive equation in the following. The plastic strain rate and the plastic spin in the intermediate configuration depend only on the plastic deformation gradient as shown in Eqs. (17.19) and (17.22). Then, let their flow rules be assumed as follows: ( p p p p D ¼ sym½Fp Fp1 ¼ k N N T ¼ N [ 0Þ ð17:25Þ ðk p pT p p W ¼ ant½Fp Fp1 ¼ k X ðX ¼ X Þ
p
p
where k is the positive plastic multiplier, and N and X are the functions of stress and internal variables in the intermediate configuration. Now, we assume that the elastic deformation is infinitesimal so that Ue ffi I; Ve ffi I
ð17:26Þ
8 e e < F ffi R e e1 sym½U U ffi sym½Ue ¼ Ue : 1 ant½Ue Ue ffi ant½Ue ¼ O
ð17:27Þ
The elastic and the plastic strain rates and spins in Eq. (17.16) and (17.17) in the current configuration are reduced for Eq. (17.27) as 8 e eT e e e e eT > > > d ffi R sym½U R ¼ R U R < e we ffi Re ReTþ R ant½Ue ReT ¼ Re ReT p p > > dp ffi Re sym L ReT ¼ Re D ReT > : p p p w ffi Re ant L ReT ¼ Re W ReT
ð17:28Þ
The plastic flow rules for Eq. (17.28) are given from Eq. (17.25) as follows: (
dp ¼ k np ðknp k ¼ 1Þ
wp ¼ k xp
ð17:29Þ
510
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
where
p
np ¼ Re N ReT p xp ¼ Re X ReT
ð17:30Þ
On the other hand, the elastic strain rate in the current configuration is given by
de ¼ E1 : s from Eq. (7.109) as the hypoelasticity. Consequently, the following hypoelastic-based plastic constitutive equations hold from Eqs. (17.28) 2 and (17.29) as follows: (
d ¼ de þ dp ¼ E1 : s þ k np w ¼ we þ wp ¼ Re ReT þ k xp
ð17:31Þ
As shown above, the hypoelastic-based plastic constitutive equation can hold under the infinitesimal elastic deformation. Besides, the cumbersome time-integrations of the cororational rates for the numerical calculation of the stress and internal variables are required as described in Subsect. 3.4.4.
17.4
Further Multiplicative Decomposition of Plastic Deformation Gradient Tensor
The plastic deformation gradient tensor is decomposed into the storage parts and the dissipative parts for the back-stress (kinematic hardening variable) and the elastic-core. The deformation gradient F is multiplicatively decomposed into the elastic deformation gradient Fe and the plastic deformation gradient Fp as described in Sect. 17.1. Further, decompose Fp into the plastic storage part Fpks causing the kinematic hardening and its plastic dissipative part Fpkd multiplicatively (Lion, 2000). Analogously, decompose Fp into the plastic storage part Fpcs causing the translation of elastic-core and its plastic dissipative part Fpcd multiplicatively as follows (see Fig. 17.3): 8 < Fp ¼ Fp Fp ; ks kd :
F ¼ p
Fpcs Fpcd ;
_
Fpks ¼ Gi G i; Fpcs
^i
¼ Gi G ;
_
Fpkd ¼ G Gi i
Fpcd
^
ð17:32Þ
¼ Gi Gi
The following representations in terms of the base vectors hold analogously to _
_i
Eq. (17.5) by introducing the base vectors ðGi ; G Þ in the kinematic-hardening _
^
^i
intermediate configuration K and ðGi ; G Þ in the elastic-core intermediate config^
uration K.
17.4
Further Multiplicative Decomposition of Plastic Deformation Gradient Tensor
511
F
x
dx
(
dX
( X
Fcsp
F
Fe
p cd
K
dX
K
X dX
X
Fp
F
K0
Fksp
) dX ) X
p kd
K
) K
Fig. 17.3 Multiplicative decompositions of deformation gradient for elastoplastic material with translations of back-stress and elastic-core
8 _i _ j _ j _ i p i p 1 < Fp ¼ G G ð¼ dij Gi G ¼ Fks ¼ Gi G i j Gi G Þ; Fks ks _ _ _i p i _ p 1 i i j j : Fp ¼ G i G ð¼ dj Gi G ¼ Fkd j Gi G Þ; Fkd ¼ Gi G kd
ð17:33Þ
8 ^i ^ j ^ j ^ i < Fp ¼ G G p 1 p i ð¼ dij Gi G ¼ Fcs ¼ Gi G i cs j Gi G Þ; Fcs ^ ^ ^i p i ^ p 1 i i j j : Fp ¼ G ¼ Gi G i G ð¼ dj Gi G ¼ Fcd j Gi G Þ; Fcd cd
ð17:34Þ
noting ( F ¼ Gi G ¼ i
p
_i _
Gi G Gj G j ¼ Fpks Fpkd ^i ^
Gi G Gj G j ¼ Fpcs Fpcd
ð17:35Þ
with the metric tensors (
_
_
_i
_i
_
_T
^
^
^i
^i
^
^T
G Gi G ¼ G Gi ð¼G Þ G Gi G ¼ G Gi ð¼G Þ
ð17:36Þ
512
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
where
_
Gi ¼
8 p p > < Fks1 Gi ð¼ ks G _G Þ ! > : Fpkd Gi ð¼ pkd G_G Þ
_i
; G ¼
8 < :
_
i
G
FpksT G ð¼ pks G Þ
ð17:37Þ
!_
FpkdT Gi ð¼ pkd GG Þ
8 8 ^ > i G p1 p < ^ ^ Fks Gi ð¼ ks GG Þ ^ i < FpksT G ð¼ pks G Þ Gi ¼ ; G ¼ ! ! > p T i : p : Fkd Gi ð¼ pkd G^G Þ Fkd G ð¼ pkd G G^ Þ
ð17:38Þ
Based on the right and left Cauchy-Green deformation tensors C FT F, the _p
^p
_p
^p
following tensors of the storage parts Cks ; Ccs and the dissipative parts Ckd ; Ccd are defined. 9 _p _i _ j _i _ j _ pT = _p 2 pT p Cks Fks Fks ð¼ Uks Þ ¼ ðG Gi ÞðGj G Þ ¼ Gij G G ð¼ Cks Þ _ ; Cp FpT Fp ¼ G Gi G j ð¼ CpT Þ kd
kd
ij
kd
kd
ð17:39Þ ^p
^p
2
^i
^
^i
j
^
^ pT
j
Ccs FpcsT Fpcs ð¼ Ucs Þ ¼ ðG Gi ÞðGj G Þ ¼ Gij G G ð¼ Ccs Þ
) ð17:40Þ
^
Cpcd FpcdT Fpcd ¼ Gij Gi G j ð¼ CpT cd Þ while one has p Cks p
Ccs
!p _ p p _ GG ¼ FpksT Cks Fpks1 ksCks ! _ p ^ p p GG ¼ FpcsT Ccs Fpcs1 csCcs
¼
FpksT Fpks T Fpks Fpks1
9 > =
> p1 ;
i
j
¼ Gij G G ¼ G
¼ FpcsT Fpcs T Fpcs Fcs
ð17:41Þ p
Further, the plastic velocity gradient L is additively decomposed for the kinematic hardening as follows: p
p
p
L ¼ Lks þ Lkd where
8 p p > > Lks Fks Fpks1 > > ! < p p p p p p 1 p 1 p _ p1 p _ L F F F F ¼ F L F ¼ L kd ks kd kd ks ks kd ks ks kd > > > p _p > p 1 p1 p p p p G :L kd Fkd Fkd ¼ Fks Lkd Fks ¼ksLkd G
ð17:42Þ
G
G
;
ð17:43Þ
17.4
Further Multiplicative Decomposition of Plastic Deformation Gradient Tensor
513
noting p
p
L ¼ ðFpks Fpkd Þ ðFpks Fpkd Þ1 ¼ ðFks Fpkd þ Fpks Fkd ÞFpkd ;1 Fpks1 p
p
p
¼ Fks Fpks1 þ Fpks Fkd Fpkd1 Fpks1 Here, it follows that p
p
p
p
p
p
Lks ¼ Dks þ Wks ; Lkd ¼ Dkd þ Wkd p
p
p
D ¼ Dks þ Dkd ;
p
p
ð17:44Þ
p
W ¼ Wks þ Wkd
ð17:45Þ
8 p p p p > < D sym L ; W sym L p p p p Dks sym Lks ; Dkd sym Lkd > p p p p : W ant L ; W ant L ks ks kd kd
ð17:46Þ
where
_p
Further, the plastic velocity gradient tensor L in the kinematic-hardening intermediate configuration can be additively decomposed into the purely storage _p
_p
part Lks and the purely dissipative part Lkd as follows: _p
_p
_p
L ¼ Lks þ Lkd
ð17:47Þ
where 8 _ _p > pG p1 p p p > > L ¼ F L F ¼ L _ < ks ks ks G
p
_p
ð17:48Þ
Lks Fpks1 Fks ; > > > p : _p Lkd Fkd Fpkd1 noting p
p
Fpks1 L Fpks ¼ Fpks1 ðLks þ Lkd ÞFpks ¼ Fpks1 ðFks Fpks1 þ Fpks Fkd Fpkd1 Fpks1 ÞFpks p
p
p
Here, it follows that _
_
_
_
_
_
_
_
_
Lp ¼ Dp þ Wp ; Lpks ¼ Dpks þ Wpks ; Lpkd ¼ Dpkd þ Wpkd _
_
_
_
_
_
Dp ¼ Dpks þ Dpkd ; Wp ¼ Wpks þ Wpkd
ð17:49Þ ð17:50Þ
514
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
where 8_ _ _ _ p p p p > > < D sym½L ; W ant½L _ _ _ _ Dpks sym½Lpks ; Dpkd sym½Lpkd > > _ _ _ : _p Wks ant½Lpks ; Wpkd ant½Lpkd
ð17:51Þ
Analogously, the following additive decomposition of the velocity gradient holds for the elastic-core. p
p
p
L ¼ Lcs þ Lcd
ð17:52Þ
where 8 p p > > Lcs Fcs Fp1 > cs < ! p ^p ^ p G p Lcd Fcs Fcd Fp1 Fp1 ¼ Fpcs Lcd Fp1 ¼pcs L pcd G cs cs cd > _ > p ^p > p G :L F Fp1 ¼ Fp1 L Fp ¼p L p _ cd
cs
cd
cs
cd cs
ð17:53Þ
cd G
noting p 1 p p p1 L ¼ Fpcs Fpcd Fpcs Fpcd ¼ Fcs Fpcd þ Fpcs Fcd Fp1 cd Fcs p
p
p1 p1 p ¼ Fcs Fp1 cs þ Fcs Fcd Fcd Fcs
Here, it follows that p
p
p
p
p
p
p
p
ð17:54Þ
W ¼ Wcs þ Wcd
p
p
p
ð17:55Þ
p p p p Dcs sym Lcs ; Dcd sym Lcd p p p p Wcs ant Lcs ; Wcd ant Lcd
ð17:56Þ
Lcs ¼ Dcs þ Wcs ; p
D ¼ Dcs þ Dcd ;
Lcd ¼ Dcd þ Wcd
where
^
Further, the plastic velocity gradient tensor Lp in the elastic-core intermediate ^
configuration can be additively decomposed into the purely storage part Lpcs and the ^
purely dissipative part Lpcs as follows: ^
^
^
Lp ¼ Lpcs þ Lpcd
ð17:57Þ
17.4
Further Multiplicative Decomposition of Plastic Deformation Gradient Tensor
515
where 8 _ > pG > > ^p p > < L ¼ Fp1 L Fp ¼p L_ ; cs G cs cs
ð17:58Þ
p
^p
> Lcs Fp1 > cs Fcs ; > > p : ^p Lcd Fcd Fp1 cd noting
p p p p p p p1 p1 p1 p Fp1 Lcs þ Lcd Fpcs ¼ Fp1 Fcs Fp1 Fpcs cs L Fcs ¼ Fcs cs cs þ Fcs Fcd Fcd Fcs Here, it follows that ^
^
^
^
^
^
^
^
^
Lp ¼ Dp þ Wp ; Lpcs ¼ Dpcs þ Wpcs ; Lpcs ¼ Dpcs þ Wpcs ^
^
^
Dp ¼ Dpcs þ Dpcs ;
^
^
ð17:59Þ
^
Wp ¼ Wpcs þ Wpcs
ð17:60Þ
h^ i h^ i 8^p p p p > D sym L ant L ; W > >
h^ i h^ i > ^ p > ^p p p : Wcs ant Lcs ; Wcd ant Lcd
ð17:61Þ
where
_p
The time-derivative of Cks in Eq. (17.39) is given by
__
_p
p GG
Cks ¼ 2pks Dks
p p p p pT p ¼ 2FpT ks Dks Fks ¼ 2Fks D Dkd Fks
ð17:62Þ
noting
p
_p
C
ks
p
p
p
¼ ðFpks T Fpks Þ ¼ Fpks T Fks þ Fks T Fpks ¼ Fpks T Fks Fpks 1 Fpks þ Fpks T Fpks T Fks T Fpks p
p
¼ Fpks T Fks Fpks 1 Fpks þ Fpks T ðFks Fpks 1 ÞT Fpks ¼ Fpks T Lks Fpks þ Fpks T Lks T Fpks ¼ 2Fpks T Dks Fpks p
p
p
Analogously from Eq. (17.40), one has
^p
^^
p GG
Ccs ¼ 2pcs Dcs
p p p p p pT ¼ 2FpT cs Dcs Fcs ¼ 2Fcs D Dcd Fcs
ð17:63Þ
516
17.5
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
Formulation and Calculation in Intermediate Configuration: Isoclinic Concept
The velocity gradient tensor cannot be decomposed muliplicatively into the purely elastic and the purely plastic parts in the current configuration as shown in Eq. (17.12) with Eq. (17.13) but the velocity gradient tensor in the intermediate configuration can be done exactly as shown in Eq. (17.18) with Eq. (17.19). Therefore, the elastoplastic constitutive equation must be formulated in the intermediate configuration. In addition, the Mandel stress (Mandel 1973b) and the Mandel-like internal variables (back-stress and elastic-core) based in the intermediate configuration must be adopted, which are work-conjugate to the velocity gradient tensors in the intermediate configuration as shown in Sect. 5.9. The deformation analysis may be performed by the following procedure. p
(1) The plastic velocity gradient tensor L composed of the plastic strain rate and the plastic spin is calculated by the plastic constitutive equation. Then, the
p
p
plastic deformation gradient tensor Fp is updated by the relation F ¼ L Fp . (2) The elastic deformation gradient Fe is calculated by subtracting the plastic deformation gradient Fp from the deformation gradient F. (3) The stress in the intermediate configuration is calculated by substituting the elastic deformation gradient into the hyperelastic equation. The rigid-body rotation Rr is included in the elastic deformation gradient tensor F which is given by excluding the plastic deformation gradient tensor Fp from the total deformation gradient tensor F in which the rigid-body rotation is included. The inclusion of the rigid-body rotation in addition of the elastic substructure rotation into the elastic deformation gradient tensor and thus the fact that the substructure does not rotate in the intermediate configuration is called the isoclinic concept (Mandel 1971, 1973, 1974). Thus, the intermediate configuration is independent of the rigid-body rotation by virtue of the isoclinic concept. Then, the constitutive relation can be formulated rigorously in the intermediate configuration as will be executed in the subsequent sections. The elastic and the plastic deformation gradient tensors are described in the polar decomposition as follows: e
Fe ¼ Re Ue ¼ Ve Re ; Fp ¼ Rp Up ¼ Vp Rp
ð17:64Þ
The physically meaningful rotation of material is to be the rotation of the substructure designated by Rs which is composed of the rigid-body rotation Rr and the elastic distortional rotation Red and thus Rs ¼ Re holds, while the plastic rotation Rp induces merely the mutual slips of material particles. The rigid-body rotation Rr is involved in the rotation part Re of the elastic deformation gradient Fe , leading to
17.5
Formulation and Calculation in Intermediate Configuration: Isoclinic Concept
RS ¼ Re ¼ Rr Red
517
ð17:65Þ
Then, the plastic deformation gradient Fp is calculated independent of the rigid-body rotation and the elastic deformation gradient Fe is calculated by e Fe ¼ FFp1 . Further, the elastic Cauchy-Green deformation tensor C FeT Fe is calculated, from which the second Piola–Kirchhoff stress tensor S in the intermediate configuration is calculated through the hyperelastic relation as will be shown in the next section. The calculations of the internal variables, i.e. the kinematic hardening variable and the elastic-core are also calculated by the similar processes. Eventually, the stress is determined independent of the rigid-body rotation by virtue of the isoclinic concept. The extensive debates as to which an elastic deformation gradient or a plastic deformation gradient must include the rigid-body rotation have been repeated for a long time after the proposition of the multiplicative decomposition. The inclusion of the rigid-body rotation in the plastic deformation gradient has been insisted by Lee (1969), Green and Naghdi (1965), Fardshisheh and Onat (1974), Casey and Naghdi (1980), Lubarda and Lee (1981), Dafalias (1985a), Boyce et al. (1988), Lubarda (2002, 2004), Khan and Huang (1995), Han et al. (2003), Wu (2004), Simo and Ortiz (1985), Asaro and Lubarda (2006), Harrysson and Ristinmaa (2007), etc. This situation would be caused by worrying the fact that the elastic distortion is known from the current stress but the rigid-body rotation is unknown and thus it is possible to exclude only the elastic distortion but it is impossible to exclude both of the rigid-body rotation and the elastic distortion from the current configuration in order to get to the intermediate configuration. On the other hand, the inclusion of the rigid-body rotation in the elastic deformation gradient has been insisted by Holsapple (1973), White (1975), Van der Giessen (1989), Haupt (2002), Wallin et al. (2003), Dettmer and Reese (2004), Wallin and Ristinmaa (2005), Vladimirov et. al. (2008, 2010), etc. The debates on this issue have been commented repeatedly without a definite conclusion by various authors (Cleja-Tigoiu and Soos 1990; Clifton 1972; Dashner 1986; Lubliner 1990; Simo 1992, 1998, Simo and Hughes 1998, etc.). This long and voluminous debate is now a relic of the past, concluding the rationality for the inclusion of the rigid-body rotation Rr into the elastic deformation gradient tensor Fe based on the isoclinic concept.
17.6
Stress Measures
Introduce the second Piola–Kirchhoff stress tensor in the intermediate configuration, which is the contravariant pulled-back (Eq. (3.24)) of the Kirchhoff stress tensor, i.e. GG
S ¼ Fe1 sFeT ð¼ s
! T Þ ¼ Fe1 FSFT FeT ¼ Fp SFpT ¼ p S G G ð¼ S Þ ð17:66Þ
518
Multiplicative Hyperelastic-Based Plasticity with Subloading …
17
noting Eq. (5.17) S ¼ F1 sFT . The Mandel stress (Mandel 1973) is given by e T M C S ¼ FeT Fe Fe1 sFeT ¼ FeT sFeT ¼ e s G ð6¼ M Þ G
ð17:67Þ
Here, note that the work-conjugate stress measure with the velocity gradient G
L ð¼ e l G Þ in the intermediate configuration is the Mandel stress M (Mandel 1973) as known from T T x0 ¼ s : l ¼ tr½ Fe SFeT Fe LFe1 ¼ trðFe SFeT FeT L FeT Þ e
T
e
s : lp ¼ M:L
p
T
¼ trðFeT Fe S L Þ ¼ trðC S L Þ ¼ C S: L ¼ M: L referring to Eq. (5.58), from which one has e
s : le ¼ M:L ;
The 2nd Piola–Kirchhoff stress push-forwarded to the intermediate configura e tion, S, is given by the following equation with the strain energy function w C , e noting Eq. (7.3), where C in Eq. (17.10) stands for the purely elastic deformation. S¼2
e @we C e
@C
T
ð¼ S Þ
ð17:68Þ
and the Mandel stress is given by e
M C S ¼ 2C
e
e @we C @C
e
T
ð6¼ M Þ
ð17:69Þ
T
while M ¼ M holds in the elastic isotropy for which we is the function of the e invariants of only C . The rate of the Mandel stress is given noting Eq. (17.23) as
e
e
e
e
e
p
M ¼ ðC SÞ ¼ L : C ¼ 2L : sym½C ðL L Þ
ð17:70Þ
e
where L is the fourth-order hyperelastic tangent modulus tensor given by e
L
@M 1 e e e ¼ S þ C :C 2 @C
ð17:71Þ
17.6
Stress Measures
519
with e @ 2 we C @S C 2 e¼4 e e @C @C @C e
17.7
ð17:72Þ
Internal Variables
Let the 2nd Piola–Kirchhoff stress-like variables for the kinematic hardening variable _
_
^
^
Sk based in K and for the elastic-core Sc based in K be formulated by the hyperelastic _p
^p
equations with the potential energy functions wk ðCksÞ and wc ðCcsÞ as follows: _p
_
Sk ¼ 2
@wk ðCksÞ _
p
@Cks
_
^p
^
T
ð¼ Sk Þ; Sc ¼ 2
@wc ðCcsÞ ^
p
@Ccs
^T
ð¼ Sc Þ
ð17:73Þ _p
^
p
Here, we adopt the postulate that wk and wc are the functions of only Cks and Ccs, respectively, because there would not exist the reason why the other tensor variable(s) _
^
must be involved in these functions, so that the Mandel-like variables Mk and Mc _
^
p
p
becomes the symmetric tensors depending only on the variable Cks and Ccs , respectively. Then, it follows noting Eq. (17.73) that _p _
_
_p
_p
Mk ¼ Cks Sk ¼ 2Cks ^p ^
^
^p
@wk ðCks Þ _p
@Cks
^p
Mc ¼ Ccs Sc ¼ 2Ccs
@wc ðCcs Þ ^p
@Ccs
9 > > > ð¼ Mk Þ; > > = _
T
> > > ð¼ Mc Þ > > ; ^
_
ð17:74Þ
T
^
Here, noting that the kinematic hardening variable Sk and the elastic-core Sc are the symmetric tensors, let their contravariant push-forward to the intermediate configuration be incorporated so as to keep the symmetry similarly to S ¼ Fp SFpT in Eq. (17.66) as follows: 9 ! _ _ T p_ p _ pT p p1 p T T > = Sk ksSk G G ¼ Fks Sk Fks ð ¼ Sk Þ; Sk ksSk __ ¼ Fks Sk Fks ð¼ Sk Þ > ! ^
Sc pcs Sc G G ¼
GG
^
Fpcs Sc FpcsT ð¼
T Sc Þ;
^ > p T ; Sc pcs Sc ^^ ¼ Fp1 ð¼ ScT Þ > cs Sc Fcs ^
GG
ð17:75Þ
520
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
The Mandel-like variables Mk and Mc are given noting Eq. (17.41) as follows: 9 ! _ p pT p _G T > Mk ¼ Cks Sk ¼ G Sk ¼ Sk ¼ FpT M F ¼ M ¼ M k ks k = ks ks kG _
_p _
_
pT Mk ¼ Cks Sk ¼ FpT ¼ pksM ks Mk Fks
Mc ¼ ^
p Ccs Sc ^p
¼ GSc ¼ Sc ¼
^
G
_ p1 p > T ; _ ¼ Mk ¼ F ks Mk Fks
k G
!
^ G ¼ pcsMcG ð¼
^ pT FpcsT Mc Fcs ^
pT
pT M c ¼ CcsSc ¼ Fcs Mc Fcs
ð17:76Þ
G ¼ pcsM ^ cG
^
McT
¼
¼
9
T = Mc Þ >
p Fp1 cs Mc Fcs
> ;
ð17:77Þ
noting Mk ¼ Sk ¼
_ Fpks Sk FpT ks
Mc ¼ Sc ¼
Fpcs Sc FpT cs
¼
^
¼
_ p1 _ Fpks Cks Mk FpT ks ^ p1 ^ p pT Fcs Ccs Mc Fcs
¼
_ pT FpT ks Mk Fks ^
pT ¼ FpT cs Mc Fcs
9 = ;
Here, the following work-conjugate equalities hold by virtue of Eqs. (17.43), (17.53), (17.76) and (17.77). (
p
p
p
^p
_
_p
_
_
^
^p
^
^
ða:dpkd ¼ÞMk : Lkd ð¼ Mk : Dkd Þ ¼ Mk : Lkd ð¼ Mk : Dkd Þ ^
ðc:dpkd ¼ÞMc : Lcd ð¼ Mc : Dcd Þ ¼ Mc : Lcd ð¼ Mc : Dkc Þ
ð17:78Þ
noting 8
T > > a ! Mk ð¼ Mk Þ; > > > > ^ T Þ; ^ ¼ MM ð 6¼ M > > b r!M k > > > T > > c ! Mc ð ¼ Mc Þ; > > > > ^ ¼ M M ð ¼ M ^ T Þ; > bc ¼ ca ! M > c c k c > > _ _ T > > _ > r ¼ rc ! M ¼ MMc ð 6¼ M Þ; > > > > ^ ð ¼ MT Þ; > > a ¼ cR^c ðca ¼ RðcaÞÞ ! Mk ¼ Mc RM c > k < _ T _ ^ r ¼ ra ¼ r þ R^ c ! M ¼ MM ¼ M þ R M ð ¼ 6 M Þ; k c > > v v T > > rv ¼ r þ a ! Mv ¼ M þ Mk ð 6¼ Mv Þ; > > R > > R > @f ðMÞ > @f ðrÞ T @f ðrÞ @f ðMÞ > > !N¼ > n¼ = = ð 6¼ N Þ > > @r @r @M @M > > " # " # > > > T > @f ðMÞ @f ðMÞ > > = sym N ¼ sym ð ¼ N 6¼ NÞ > > > @M @M > > > > ^ Þ @f ðM ^ Þ > > @f ð^ c Þ @f ð^ c Þ @f ð M c c ^TÞ ^ ¼ >n !N > = = ð ¼ N c : ^c ¼ c @c @c @Mc @Mc
ð17:88Þ
17.8
Normal-Yield, Subloading and Elastic-Core Surfaces
523
N ( N) My
Normal-yield surface M
M
RM
ˆ ) = F (H ) f (M
ˆ N c
M M
Subloading surface f (M ) = RF ( H )
Mc ˆ M c
Mk
Elastic-core surface ˆ ) = R F (H ) f (M c c
Mk
Limit elastic-core surface ˆ ) = F (H ) f (M c
0
M ij M y M k (M M k ) / R M / R ˆ ( / R )M M Mk M y
(a) General material
M3
N (= M' / ||M' || ) N) Normal-yield surface
My
M
R
Mk
M
M
Mk
ˆ M c
ˆ ' || = F ( H ) 3 / 2|| M
M M
Subloading surface ˆ ' || RF ( H ) ' RM 3 / 2 || M c
ˆ N c
Mc
Elastic-core surface ˆ 3 / 2 || M'c|| = R c F ( H )
Limit elastic-core surface ˆ 3 / 2 || M'c|| = F ( H )
M1
M2 (b) Mises metals
Fig. 17.4 Normal-yield, subloading, elastic-core and limit elastic-core surfaces in the intermediate configuration.
524
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
The subloading surface is given from Eq. (17.87) with Eq. (17.88)7 as follows: _
^ Þ ¼ RFðHÞ f ðM þ RM c
ð17:89Þ
from which the normal-yield ratio R is calculated. The material-time derivative of the conjugate kinematic hardening variable Mk is given by ^ Mk ¼ R Mk þ ð1 RÞ Mc R M ð17:90Þ c leading to
^ M ¼ M Mk ¼ M R Mk ð1 RÞ Mc þ R M c
ð17:91Þ
The elastic-core surface which passes through the elastic-core Mc and is similar to the normal-yield surface with respect to the back-stress Mk in the hyperelastic-based plasticity is given noting Eq. (11.14) as follows:
ˆ ) ℜ F ( H ), i.e. ˆ ) / F (H ) f (M ℜ c = f (M c = c c
17.9
ð17:92Þ
Plastic Flow Rules
The purely-plastic strain rate is given in the following associated flow rule proposed by Hashiguchi (2017, 2020). p
pT
D ¼ k Nð¼ D Þðk 0Þ
ð17:93Þ
where k is the positive plastic multiplier and " # " # T @f ðMÞ @f ðMÞ = sym N sym ð¼ N ÞðkNk ¼ 1Þ @M @M
ð17:94Þ
which is the normalized and symmetrized tensor. If the strain energy functions we is e given by the invariants of only C leading to the elastic isotropy, the symmetry of the
Mandel
stress
M¼M
T
holds
resulting
in
the
symmetry
T
@f ðMÞ=@M ¼ ð@f ðMÞ=@MÞ . The purely-dissipative part of the plastic velocity gradient for the kinematic hardening variable is given by virtue of Eq. (11.13) as follows: _p
Dkd ¼
_ pT 1 _ k Mk ð¼ Dkd Þ bk F
ð17:95Þ
17.9
Plastic Flow Rules
525
which is push-forwarded to the intermediate configuration by virtue of the common _p
_
transformation rule of Lkd and Mk in Eq. (17.80) as follows: p
Dkd ¼
1 pT k Mk ¼ Dkd bk F
ð17:96Þ
Analogously, the dissipative parts of the plastic velocity gradient for the elasticcore is given by referring to Eq. (11.35) as follows:
^ ^ ^ 1 v ^ 1 ^ F 0 fHn ^^ ð17:97Þ ¼ kN k M Mc þ ck ðN k k Mk Þ þ k Mc H cs R bk F
^p
Dcd
^
where
2 3 2 3 ^ ^ ^T ^ @f ðM Þ @f ðM Þ ð¼ N ÞðjjNjj ¼ 1Þ 4 5 4 5 = sym N ¼ sym ^ ^ @M @M ^
leading to p Dcd
F 0 fHn ^ 1 v 1 ^ ¼ kN k M Mc þ ck N Mk þ Mc F cs R bk F
¼ k Mcd ð¼
ð17:98Þ
pT Dcd Þ
where 9 > M N Mcdc ; > =
0 1 v ^ > ^ þ c N 1 M þ F fHn M > Mcdc MM c k k c ; F cs R bk F ^p
ð17:99Þ
^
by virtue of the common transformation rule of Lcd and Mc in Eq. (17.80). The variables R c and Cn in the material parameter u in Eq. (11.52) for the evolution rule of the normal-yield ratio R are given based on Eqs. (11.14) and (11.51) for the infinitesimal strain theory. They are given for the multiplicative hyperelastic-based plasticity as follows (Hashiguchi 2016, 2020a, b): ˆ )/F R c ≡ f (M c
ð17:100Þ
Cn N^c :Nð1 Cn 1Þ
ð17:101Þ
526
17
where
Multiplicative Hyperelastic-Based Plasticity with Subloading …
^ Þ @f ðM ^ Þ @f ðM c c ^ ^ T ÞðkN ^ k ¼ 1Þ Nc = ð¼ N c c @Mc @Mc
ð17:102Þ
which is the normalized outward-normal to the elastic-core surface in Eq. (17.92). p p Let the plastic spin W , the kinematic hardening spin Wkd and the elastic-core ^
p
p
spin Wcd be given by adopting the work-conjugate pairs ðM; D Þ; p p ðMk ; Dkd Þ and ðMc ; Dcd Þ in Eq. (17.78) with Eqs. (17.93), (17.96) and (17.98) p instead of ðs; d Þ in Eq. (8.94) as follows: p
p
W ¼ gp ant½M D ¼ gp ant½M Nk p
p
ð17:103Þ
Wkd ¼ gpk ant½Mk Dkd ¼ ½gpk =ðbk FÞ ant ½Mk Mk k ¼ O p
p
Wcd ¼ gpc ant½Mc Dcd ¼ gpc ant ½Mc Mcd k p
where gpk and gpc are the material constants. The plastic spin tensor W diminishes if T
the symmetry of the Mandel stress, i.e. M ¼ M due to the elastic isotropy and the p plastic isotropy due to Mk ¼ Mc ¼ O hold. Further, the spin tensor Wcd diminishes for Mc ¼ O. The substructure spin is given by substituting Eqs. (17.22) and (17.103) into Eq. 17.22 as follows: p
W ¼ ant½L gp ant½M D ¼ ant½L gp k ant½M N
ð17:104Þ
The velocity gradients are given by substituting Eqs. (17.93), (17.96), (17.98) and (17.103) into Eqs. (17.20)3, (17.44)2 and (17.54)2 as follows: p
L ¼ k ðN þ gp ant½M NÞ p
Lkd ¼ k Mk =ðbk FÞ p
ð17:105Þ
Lcd ¼ k ðMcd þ gp ant½Mc Mcd Þ The substitution of Eq. (17.105) into Eqs. (17.70) and (17.84) yields: 9 e e > > M ¼ 2L :sym½C fL kðN þ gp ant½M NÞg > > > > > p pT > p k pT > Mk ¼ kfFcs C :Fcs ½N ½1=ðbk FÞMk Fks Fks > = p ð17:106Þ þ 2sym½ðN þ g ant ½M N ½1=ðbk FÞMk ÞMk g > > > > > p pT > > Mc ¼ kfFpcs Ck :FpT > cs Mcdc Fks Fks > > ; p p þ 2sym½ðMcdc þ g ant½M N gc ant½Mc Mcd ÞMc g
17.10
Plastic Strain Rate
17.10
527
Plastic Strain Rate
The elastic constitutive equation is given by Eqs. (17.68) and (17.69). The plastic strain rate will be formulated in this section. The formulation of the plastic modulus given in this section is not necessary in the implicit numerical calculation by the return-mapping based on the closet-point projection. The time-differentiation of Eq. (17.87) leads to the consistency condition of the subloading surface as follows:
@f ðMÞ :MRF RF ¼ 0 @M
ð17:107Þ
It holds from Eq. (17.87) that @f ðMÞ @M
: Mð¼ f ðMÞÞ ¼ RF
ð17:108Þ
by the Euler’s theorem for the homogenous function f ðMÞ of M in degree-one, and then it follows that N: M ¼
@f ðMÞ M
@f ðMÞ @f ðMÞ @f ðMÞ :M ¼ f ðMÞ ¼ RF M M M
which leads to @f ðMÞ N: N ¼ 1 RF @M
ð17:109Þ
@f ðMÞ T N 6¼ N ; N ¼ 1 M @M
ð17:110Þ
where @f ðMÞ
The substitution of Eq. (17.109) into Eq. (17.107) leads to ! F R þ N: M N:M ¼ 0 F R The further substitution of Eq. (17.91) into Eq. (17.111) leads to " # F R ^ N: M N: M þ M RMc þ R Mk þ ð1 RÞ Mc ¼ 0 F R
ð17:111Þ
ð17:112Þ
528
Multiplicative Hyperelastic-Based Plasticity with Subloading …
17
Furthermore, substituting the relation _
^ ¼ M M ðM M Þ ¼ M M RM c k c k
ð17:113Þ
Equation (17.112) is rewritten as
F0 H R_ N: M N: M þ M þ R Mk þ ð1 RÞ Mc ¼ 0 F R
ð17:114Þ
where p p p H ¼ fHd M; H; D = D D ¼ fHn ðM; H; NÞ k
R = U ( R,
c , Cn ) || D
p
|| = U ( R,
for D p
c , Cn )
ð17:115Þ
O
ð17:116Þ
based on Eqs. (11.12) and (11.53) with Eq. (17.93). The substitutions of Eqs. (17.106), (17.115) and (17.116) into Eq. (17.114) lead to the consistency condition:
p
N : M M k ¼ 0
ð17:117Þ
from which it follows that
N: M k¼ p ; M
N: M D ¼ p N M p
ð17:118Þ
where
ð17:119Þ The substitution of Eq. (17.106)1 into Eq. (17.117) leads to the consistency condition: h e i o e n e e p 2N: L : sym C L 2N : L : sym C N þ gp ant ½M N þ M K ¼ 0 ð17:120Þ
17.10
Plastic Strain Rate
529
using the symbol K for the plastic multiplier in terms of the strain rate instead of k in terms of the stress rate. The plastic multiplier is given from Eq. (17.120) as follows: e e 2N:L :sym C L h e i K¼ p e M þ 2N:L :sym C N þ gp ant ½M N
ð17:121Þ
where the following holds from Eq. (17.19). e
C L ¼ FeT lFe
ð17:122Þ
The loading criterion is given by ( p
e
e
D ¼ 6 O for K or N:L : sym½C L [ 0 p D ¼ O for other
ð17:123Þ
The normal-yield ratio is calculated by Eq. (11.50) (Eq. (12.32) for metals) by replacing r; a and c to M; Mk and Mc ; respectively. The formulation of the multiplicative hyperelastic-based plasticity shown in this section would possess the generality, while the past formulations, e.g. Wallin et al. (2003), Dettmer and Reese (2004), Wallin and Ristinmaa (2005), Vladimirov et al. (2008, 2010), Hashiguchi and Yamakawa (2012) possess the limitations, e.g. the elimination of the plastic spin leading to the overall formulation in the reference configuration.
17.11
Material Functions for Metals and Soils
Metals exhibit the pressure-independence of the plastic deformation leading to the plastic incompressibility and thus its yield surface has the cylindrical shape in the stress space. On the other hand, soils exhibit the pressure-dependence of plastic deformation leading to the plastic compressibility and thus it possesses the yield surface with ellipsoidal shape, conical shape, hexagonal pyramid shape, etc. Material functions contained in the constitutive equations formulated in the preceding sections are given for metals and soils in this section.
17.11.1
Metals
The hyperelastic equation and the yield function for metals are shown below in the intermediate configuration.
530
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
(1) Hyperelastic equation The following strain energy function of elastic deformation in Eq. (7.51) for the Modified Neo-Hookean elasticity (3) may be adopted (Vladimirov et al. 2008, 2010) in the multiplicative hyperelastic-based plastic constitutive equation. pffiffiffiffiffiffiffiffiffiffiffiffie i 1 e pffiffiffiffiffiffiffiffiffiffiffiffie e 1 h e det C þ l trC 3 2 ln det C we C ¼ K det C 1 2 ln 4 2 ð17:124Þ where l and K are the material constants. The substitution of Eq. (17.124) into Eq. (17.68) reads: e1 1 e e1 S ¼ K det C 1 C þ l G C 2
ð17:125Þ
noting Eq. (4.63). It follows from Eq. (17.125) that e 1 e e M ¼ C S ¼ K det C 1 G þ l C G 2
ð17:126Þ
(2) Hyperelastic equation for kinematic hardening variable Assume the following strain energy function for the kinematic hardening, which possesses the identical form to the shear part in Eq. (17.124). qffiffiffiffiffiffiffiffiffiffiffiffiffiffi _ p _ 1 _p p trCks 3 ln det Cks wk Cks ¼ Ck 2
ð17:127Þ
where Ck is material constant. It is derived from Eq. (17.73) that _
Sk ¼ 2
_ p @wk Cks _
p
@Cks
_ _ p1 ¼ Ck G Cks
ð17:128Þ
Then, the Mandel-like kinematic hardening variable is given from Eqs. (17.75) and (17.128) as follows: Mk ¼ Ck Fpks FpT ks G
ð17:129Þ
noting Mk ¼
_ Fpks Sk FpT ks
¼
Fpks Ck
1 _ _ p1 _ pT p pT p FpT G Cks Fks ¼ Fks Ck G Fks Fks ks
17.11
Material Functions for Metals and Soils
531
(3) Hyperelastic equation for elastic-core Assume the following strain energy function for elastic-core analogously to the kinematic hardening variable in Eq. (17.127). qffiffiffiffiffiffiffiffiffiffiffiffiffi ^ p ^ 1 ^p p trCcs 3 ln det Ccs w Ccs ¼ Cc 2 c
ð17:130Þ
where Cc is material constant. The following relations hold. ^ p @wc Ccs
^ ^ p1 ¼ Cc G Ccs
ð17:131Þ
p pT ^ Mc ¼ Fpcs Sc FpT cs ¼ Cc Fcs Fcs G
ð17:132Þ
^
Sc ¼ 2
^
p
@Ccs
(4) Yield function Assume the von Mises yield function and the plastic equivalent hardening, i.e. rffiffi3ffi ^ 0 ^0 ¼ f M M 2
ð17:133Þ
Z rffiffiffi 2 Dp dt H¼ 3
ð17:134Þ
rffiffiffi rffiffiffi 2 p D ¼ 2 k H¼ 3 3
ð17:135Þ
FðHÞ ¼ F0 f1 þ Sr ½1 expðcH H Þg; for which we have
0
F ¼ F H;
0
F F0 Sr cH expðcH H Þ;
F ! ð1 þ h1 ÞF0 holds for H ! 1 in Eq. (17.134).hIt follows for Eq. (17.133) that i 0 ^ 0 ^ sym M ^ ¼ ^ ¼ M ; N h i N ð17:136Þ ^ 0 ^0 sym M M
17.11.2
Soils
The hyperelastic equation and the yield functions for soils are shown below in the intermediate configuration.
532
Multiplicative Hyperelastic-Based Plasticity with Subloading …
17
(1) Hyperelastic equation The hyperelastic equation of soils was delineated in Sect. 13.9. Its extension to the multiplicative-hyperelastic equation will be described in the following. The isotropic hyperelastic equation is given by introducing the function of the e e variables lnJ e and trC (ð1=2ÞðtrC 3Þ in detailed expression) which stand for the volumetric strain and deviatoric strain in the infinitesimal strain theory, respectively, as follows: S¼2
e @w ln J e ; trC @C
e
e e e @w ln J e ; trC @ ln J e @w ln J e ; trC @trC ¼2 þ 2 e e e @ ln J e @C @C @trC ð17:137Þ
where 8 e e e J ¼ det Fe ; C ¼ FeT Fe ; det C ¼ J e2 > > > e e e e e e1=3 > F ¼ Fvol F ; Fvol J g; F J e1=3 Fe < e e e e C FeT Fe ¼ J e2=3 C ; trC ¼ J e2=3 trC
> e e > > det F ¼ det C ¼ 1 > : e e e0 C ¼ G; trC ¼ 3; C ¼ O for Fe ¼ Fevol
ð17:138Þ
Fevol is the elastic volumetric part and Fe is the so-called unimodular tensor designating the isochoric (constant volume, i.e. deviatoric) part of Fe . In addition, e trC ¼ 3 Fe ¼ Fevol and ln J e ¼ 0ðFe ¼ Fe Þ are required in the purely volumetric deformation and the purely deviatoric deformation, respectively. The necessity of e ð1=2Þ in ð1=2ÞtrC for the deviatoric infinitesimal strain was suggested by Prof. Yuki Yamakawa, Tohoku university. The following partial derivatives hold noting Eq. (4.63). 8 @ ln J e 1 1 e1 e e1 > pffiffiffiffiffiffiffiffiffiffiffiffie det C C ¼ C > e ¼ < 2 @C 2J e det C
e > @trC 1 e e1 > e2=3 : G trC C e ¼ J 3 @C
ð17:139Þ
The substitution of Eq. (17.139) into Eq. (17.137) reads:
e e @w ln J e ; trC @w ln J e ; trC e2=3 1 e e1 e1 ð17:140Þ S¼ C þ J G C trC e 3 @ ln J e @trC The strain energy function in Eq. (13.182) is rewritten in the intermediate configuration as follows:
17.11
Material Functions for Metals and Soils
533
e e ~ PM0 þ pe J e1=~j þ ð1=2ÞG0 J en=~j trC 3 w ln J e ; trC ¼ pe ln J e þ j ð17:141Þ noting
n J en=~j ¼ exp ln J e ~ j
where PM0 is the initial value of the pressure defined in terms of the Mandel stress M, i.e. PM ð1=3ÞtrM. The following partial derivatives hold for Eq. (17.141). 8 e e e @w ln J ; trC n > > ¼ pe PM0 þ pe J e1=~j G0 J n=~j trC 3 < ~ j @ ln J e @w ln J e ; trC n=~ j > > : ¼ G0 J e e @trC
ð17:142Þ
n=~ j
noting @J en=~j =@ ln J e ¼ ðn=~ jÞJ e . Equation (17.140) with Eq. (17.142) reads:
h i e1 n 1 e e1 n=~ j e S ¼ pe PM0 þ pe J e1=~j G0 J e trC 3 C þ G0 J en=^j J e2=3 G trC C ~ j 3
ð17:143Þ from which the Mandel stress is given as follows: h e i n e 0 M ¼ C S ¼ pe PM0 þ pe J e1=~j G0 J en=~j trC 3 G þ G0 J en=~j Ce ~ j ð17:144Þ e0
e0
noting J e2=3 C ¼ C by virtue of Eq. (17.138)7 . It is follows from Eq. (17.144) that
P M þ pe 1 ¼ J e1=~j ¼ exp ln J e for Fe ¼ Fevol ~ j PM0 þ pe
ð17:145Þ
PM þ pe for Fe ¼ Fevol PM0 þ pe
ð17:146Þ
i.e. j ln eev ¼ ln J e ¼ ~ and 0
M ¼ G0 J
en=~ j
e0
C ¼ G0
P M þ pe PM0 þ pe
n
e0
C
ð17:147Þ
534
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
which would describe appropriately the basic characteristics in the volumetric and the deviatoric deformations. Equation (17.144) is rewritten in terms of the Kirchhoff stress s and the elastic unimodular left Cauchy-Green deformation tensor be in the current configuration as follows: h i n s ¼ pe ðps0 þ pe ÞJ e1=~j G0 J en=~j ðtrbe 3Þ g þ G0 J en=~j be0 ~ j
ð17:148Þ
where s ¼ FeT MFeT ¼ Fe SFeT
ð17:149Þ
be Fe FeT
ð17:150Þ
and ps0 is the initial value of ps ð1=3Þtrs, noting
e
FeT C FeT ¼ be ; e trC ¼ trbe
e0
FeT C FeT ¼ be0
ð17:151Þ
It follows from Eq. (17.148) that j ln ln J e ¼ e
ps þ pe for Fe ¼ Fevol ps 0 þ pe
ð17:152Þ
and s0 ¼ G0
ps þ pe ps0 þ pe
n be0
ð17:153Þ
The above-mentioned elastic equation is reduced to the infinitesimal strain theory by adopting the strain energy function
e e e e ~ðp0 þ pe Þ exp v þ G0 exp n v ee2 w eev ; eed ¼ pe eev þ j d ~ ~ j j
ð17:154Þ
as follows: r¼
e e e ev n ev e e2 pe ðp0 þ pe Þ exp G0 exp n ed g þ G0 exp n v ee0 ~ ~ ~ ~ j j j j
ð17:155Þ
17.11
Material Functions for Metals and Soils
535
resulting in
e p þ pe e ¼ exp v for ee0 ¼ O eed ¼ 0 ~ j p0 þ pe i.e.
eev ¼ ~ j ln
p þ pe p0 þ pe
for ee0 ¼ O eed ¼ 0
ð17:156Þ
and
e e p þ pe n e0 e r0 ¼ 2G0 exp n v ee0 ¼ 2G0 ~ j p0 þ pe
ð17:157Þ
where p is the pressure, i.e. p ðtr rÞ=3 and its initial value is denoted by p0 . ee is the infinitesimal elastic strain tensor and eev ¼ tree ; eed ¼ kee0 k. The elastic balk modulus K and the elastic shear modulus G adopted for the above-mentioned elastic equation in the infinitesimal strain theory are given as follows: 0
r p p þ pe p þ pe n ; G ¼ e0 ¼ G0 ð17:158Þ K ¼ e ¼ ~ j p0 þ pe ev 2 e The elastic constitutive equations of soils formulated above possess the following physical validities. (1) It is applicable up to the finite deformation/rotation. (2) It is applicable up to the negative pressure range which depends on the preconsolidation pressure. (3) The shear modulus increases depending on the pressure. (4) They are consistent for the multiplicative hyperelasticity and the infinitesimal hyperelasticity. (2) Yield function The yield surface in Eq. (13.74) is extended to the multiplicative strain theory as follows:
PM ½ð1=2Þ nh F F=2
2 þ
9 8 0 2 < M = :MF=2;
¼1
ð17:159Þ
536
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
where pffiffiffi 7 2 6 sin /c 7 M¼ ¼ Mc 8 þ cos 3hM 3 sin /c 8 þ cos 3hM cos 3hM
pffiffiffi 03 6tr tM ;
ð17:160Þ
0
M t0M 0 M
ð17:161Þ
Equation (17.159) is expressed in the separated form of the function f PM ; q of the stress and the hardening function F, i.e. f PM ; q ¼ F
ð17:162Þ
where the yield stress function is given as (Hashiguchi and Mase 2008; Hashiguchi 2017): 8 h 2 i < PM 1 þ q=PM f PM ; q ¼ 1 : ~ PM q nPM n
~ n 2ð1 nÞn;
n 1 2n; PMq q
for nh ¼ 0 for nh 6¼ 0 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 PM þ 2~nq2
0 M M
ð17:163Þ
ð17:164Þ ð17:165Þ
and the isotropic hardening function is given by (Hashiguchi 2017). FðHÞ ¼ F0 exp
ln J p ~j ~ k
ð17:166Þ
where J p ¼ det Fp .
17.12
Calculation Procedures
The calculation procedure for the above-mentioned formulations is described in this section. The deformation gradient tensor is updated by Fn þ 1 ¼ f ½n;n þ 1 Fn
ð17:167Þ
17.12
Calculation Procedures
537
where f ½n;n þ 1 I þ Du rxn
ð17:168Þ
with the displacement vector u, designating rxn ð Þ @ð Þ=@xn and noting f ½n;n þ 1 Fn þ 1 F1 n ¼
@xn þ 1 @X @xn þ 1 @ ðxn þ DuÞ ¼ ¼ ¼ I þ Du rxn @X @xn @xn @xn ð17:169Þ
The plastic multiplier K is calculated by the input of the velocity gradient L into Eq. (17.121), while L is calculated from the current velocity gradient l by Eq. (17.19). Then, substituting it into Eq. (17.105), the plastic and the dissipative p
p
p
parts L , Lkd and Lcd are calculated. On the other hand, K is calculated directly from the plastic flow rule in Eq. (17.93) under L ¼ O in the plastic corrector step in the return-mapping method. Thereafter, the stress and the tensor-valued internal variables are calculated by the procedures described below. The rates of the plastic gradient and its dissipative parts are given from Eqs. (17.19)3, (17.43) 3 and (17.58) 3 as follows: 8 p p p > > < F p ¼ L_ pF p p p FKd ¼ Lkd Fpkd ¼ Fp1 ks Lkd Fks Fkd > > ^p p p p :p Fcd ¼ Lcd Fpcd ¼ Fp1 cs Lcd Fcs Fcd p
p
p
ð17:170Þ
where L , Lkd and Lcd are given by Eq. (17.105), while k or K is calculated from
Eq. (17.118) or (17.121) by the inputs of M or L which is calculated by substituting l and Fe into Eq. (17.19)1 . The storage parts Fe ; Fpks and Fpcs of the deformation gradient are given by substituting the results of the time-integrations of Eq. (17.170) into Eq. (17.32) as follows: Fpks ¼ Fp Fp1 kd ;
Fe ¼ FFp1 ; e
_p
Fpcs ¼ Fp Fp1 cd
ð17:171Þ
^p
Further, C ; Cks and Ccs are calculated by substituting Eq. (17.171) into Eqs. (17.10), (17.39) and (17.40). Further, the stress S, the kinematic hardening _
^
e
_p
^p
variable Sk and the elastic-core Sc are calculated by substituting C ; Cks and Ccs into Eqs. (17.68) and (17.73). The isotropic hardening variable and the normal-yield ratio are calculated by the time-integration of Eqs. (17.115) and (17.116). The stress tensors in the current configuration are calculated by Eq. (17.66)1 , i.e. s ¼ Fe SFeT and r ¼ ðdet FÞs.
538
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
The time-integrations of Eq. (17.170) for the tensors Fp , Fpkd and Fpcd can be executed in high efficiency by the tensor exponential method (Miehe 1996; Weber and Anand 1990; Hashiguchi and Yamakawa, 2012) which is delineated below. The tensor Z in the incremental relation in terms of the tensor exponential p
_p
^p
function described in Eq. (A.54) in Appendix J is given by L , Lkd and Lcd to obtain the updated values of Fp , Fpkd and Fpcd , respectively, as shown in Eq. (17.170), so that they are updated by substituting Eq. (17.105) as follows:
n h i o 8 p p p p > F ¼ exp L Dt F ¼ exp N þ g ant M N DK Fpn n þ 1 n þ 1 n þ 1 n þ 1 > n n þ 1 n þ 1 > > > p > p p > Fpkd n þ 1 ¼ exp Fp1 > ks n þ 1 Lkd n þ 1 DtFks n þ 1 Fkd n > > > p1 p p > > ¼ expfF > > ks n þ 1 DKn þ 1 ½Mk n þ 1 =ðbk Fk þ 1 Þ Fks n þ 1 gFk dn < p p p p Fcd n þ 1 ¼ exp Fp1 cs n þ 1 Lcd n þ 1 DtFcs n þ 1 Fc dn n o > > p1 p p > p > ¼ exp F DK ðM þ g ant ½M M ÞF nþ1 cd n þ 1 cn þ 1 cd n þ 1 > c cs n þ 1 cs n þ 1 Fc dn > > > > Hn þ 1 ¼ Hn þ fHn n þ 1 DKn þ 1
> >
> > 2 p hRn Re i p DKn þ 1 > > : Rn þ 1 ¼ ð1 Re Þ cos1 cos exp u þ Re p 2 1 Re 2 1 Re ð17:172Þ in which the plastic multiplier DKn þ 1 is calculated by Eq. (17.121) or by the return-mapping projection. Here, it is required to input their initial values at the start of the numerical calculation. One may input Fp0 ¼ Fpkd0 ¼ Fpcd0 ¼ G for initial isotropic materials under a null initial stress state. Fe , Fpks and Fpcs are calculated uniquely by substituting Fp , Fpkd and Fpcd calculated from Eq. (17.172) into p p p1 Eq. (17.32), i.e. Fe ¼ FFp1 ; Fpks ¼ Fp Fp1 kd ; Fcs ¼ F Fcd . Further, the Mandel stress M and the internal state variables Mk and Mc are calculated by substituting Fe , Fpks and Fpcs into the hyperplastic Eqs. (17.69) and (17.74) with Eqs. (17.76) and (17.77). Therein, the rigid-body rotation is not included in the plastic deformation gradient but included in the elastic deformation gradient tensor, obeying the isoclinic concept, i.e. the independence of the intermediate configuration on the direction of the rotation of the substructure. The detailed implicit calculation procedure by the return-mapping projection is described in Hashiguchi (2020).
17.13
Isotropic Hardening Stagnation
The isotropic hardening stagnation behavior is formulated within the framework of the infinitesimal strain in Sect. 12.2. It will be extended to the framework of the multiplicative hperelastic-based plasticity in this section.
17.13
Isotropic Hardening Stagnation
539
Firstly, the plastic deformation gradient tensor Fp is multiplicatively decomposed into the storage part Fpss and the dissipative part Fpsd for the isotropic hardening stagnation as follows: Fp ¼ Fpss Fpsd ð17:173Þ from which the storage part is expressed as Fpss ¼ Fp Fp1 sd
ð17:174Þ
We define the following right Cauchy-Green tensor for the storage part. p
p Css ¼ FpT ss Fss
ð17:175Þ
which is based in the isotropic hardening stagnation intermediate configuration K . Then let the following normal-isotropic hardening stagnation surface. p
GðC ss Þ ¼ K
ð17:176Þ
where K is the size of the surface. Further, let the following subloading-isotropic hardening stagnation surface be incorporated. p
GðC ss Þ ¼ R K
ð17:177Þ
which passes through the current plastic deformation gradient Fp and is similar to
the normal-isotropic hardening stagnation surface, where R is the ratio of the size of the subloading-isotropic hardening surface to that of the normal one. The consistency condition of the subloading-isotropic hardening stagnation surface is given by p
p
@GðCss Þ @Css p : p : Fss ¼ R K þ R K p @Fss @C
ð17:178Þ
ss
in which one has p
@Css ¼ I Fpss þ FpT ss I @Fpss p @C ss @Fpss
ijkl
¼
@FpT Fp ss
@Fpss
ss
ijkl
p ¼ dil F psskj þ Fsski dlj
ð17:179Þ
540
Multiplicative Hyperelastic-Based Plasticity with Subloading …
17
noting Eq. (1.196). Then, Eq. (17.178) leads to p
@GðCss Þ p
@C ss p @GðC Þ p
ss
@C ssij i.e.
p
@GðC ss Þ p
@C ssij
: ðI Fpss
þ FpT ss
p
IÞ : Fss ¼ R K þ R K
p
p p ðdil Fsskj þ Fssik dlj ÞF sskl ¼ R K þ R K
p p ðFsskj F sski
p p þ Fssik F sskj Þ
ð17:180Þ
¼ R K þ R K
ð17:181Þ
Further, one has p
p
p
p1 p p1 F Fsd ¼ L Fpss F Fp1 sd ¼ F F p
p
p1
þ Fp Fsd Fss ¼ F Fp1 sd
p
ð17:182Þ
p1
¼ L Fpss þ Fp Fsd
ð17:183Þ
noting Eq. (17.174). The substitution of Eqs. (17.182) and (17.183) into Eq. (17.180) leads to p
@GðCss Þ p
@Css
p1
p
p p : ðI Fpss þ FpT ss IÞ : ðL Fss þ F Fsd Þ ¼ R K þ R K
ð17:184Þ
which is reduced to
N : ðI Fpss
þ FpT ss
p IÞ : ðL Fpss
p1 þ Fp Fsd Þ
p . @GðC ss Þ ¼K p @Css
for R ¼ 1 ð17:185Þ
in the normal-isotropic stagnation state, where p p @GðCss Þ . @GðCss Þ N¼ p p @Css @C ss
Now, we assume
ð17:186Þ
17.13
Isotropic Hardening Stagnation
1
K ¼ CR
h N: ðI Fpss
1
¼ 2CR
þ FpT ss
h N: ðI Fpss
541
p IÞ:L Fpss
þ FpT ss
p @GðC ss Þ i p @Css
p p @GðC ss Þ NÞFss i p
IÞ: kðN þ g ant½M p
@Css
ð17:187Þ and 1
p1
p
1
p p pT Fp FSd ¼ ðC 1ÞR h N: ðI FpSS þ FpT SS IÞ:L FSS iðI FSS þ FSS IÞ N
1
1
p p pT p ¼ ðC 1ÞR h N: ðI FpSS þ FpT SS IÞ: kðN þ g ant½M NÞFSS iðI FSS þ FSS IÞ N
ð17:188Þ so as to fulfill Eq. (17.185), noting Eq. (17.105). It follows by substituting Eqs. (17.188) into Eq. (17.183) that p
p p1
p p1
Fss ¼ F Fsd þ F Fsd
1
p
1
p
p pT p ¼ L Fpss ðC 1ÞR ðI Fpss þ FpT ss IÞ N h N : ðI Fss þ Fss IÞ : L Fss i
1
p
1
p
p pT p ¼ fL þ ðC 1ÞR ðI Fpss þ FpT ss IÞ N ½N : ðI Fss þ Fss IÞ:L gFss
1
1
¼ fkðN þ gp ant½M NÞ þ ðC 1ÞR ðI FpSS þ FpT ss IÞ N
p p ½N:ðI Fpss þ FpT ss IÞ: kðN þ g ant½M NÞgFss
ð17:189Þ p
The time-integration of Fss in the form p
p
Fss ¼ ZFss
ð17:190Þ
Fpssn þ 1 ¼ exp ðZn þ h DtÞFpssn ð0 h 1Þ
ð17:191Þ
can be performed to
where p
1
1
p
p pT Z ¼ L þ ðC 1ÞR ðI Fpss þ FpT ss IÞ N ½N : ðI Fss þ Fss IÞ:L
ð17:192Þ noting Eq. (17.189).
542
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
Let the rate of the isotropic hardening variable be given by 8rffiffiffi > 2m > p pT p p > R h N :ðI F þ F IÞ : k ðN þ g ant½M NÞF i ¼ k fH n > ss ss ss > 3 > > < p p for N : ðI Fpss þ FpT ss IÞ : L Fss H¼ > > > > ¼ N: ðI Fpss þ FpT IÞ : k ðN þ gp ant½M N ÞFpss [ 0 > ss > > : 0 for others ð17:193Þ where
f
Hn
rffiffiffi 2m p p R h N: ðI Fpss þ FpT ¼ ss IÞ : ðN þ g ant½M NÞFss i 3
ð17:194Þ
p
The function GðCss Þ for Eq. (12.28) is given simply as follows: p
p
GðC ss Þ ¼ jjCss jj;
17.14
p
R ¼ jjCss jj = K
ð17:195Þ
Subloading-Overstress Model
The subloading-overstress model described in Sect. 14.4 will be extended to the multiplicative hyperelastic-based viscoplasticity in this section.
17.14.1
Constitutive Equation
The deformation gradient F is multiplicatively decomposed into the elastic deformation gradient Fe and the viscoplastic deformation gradient Fvp instead of the plastic deformation gradient Fp in the multiplicative elastoplasticity described in the preceding sections. Then, we first adopt the following equation instead of Eq. (17.2). F ¼ Fe Fvp
ð17:196Þ
Further, the viscoplastic deformation gradient Fvp is multiplicatively decomposed into the viscoplastic storage part Fvp ks causing the kinematic hardening and its and into the viscoplastic storage part Fvp dissipative part Fvp cs causing the elastic-core kd vp and its dissipative part Fcd as follows:
17.14
Subloading-Overstress Model
543
Fvp ¼
vp Fvp ks Fkd vp Fvp cs Fcd
ð17:197Þ
Then, the following right Cauchy-Green deformation tensors for the viscoplastic deformation are introduced analogously to Eqs. (17.39) and (17.40). ( _ vp vp vp vpT vp Cks FvpT ks Fks ; Ckd Fkd Fkd ^ vp
vp vpT vp vp Ccs FvpT cs Fcs ; Ccd Fcd Fcd
ð17:198Þ
The velocity gradient l in the current configuration is additively decomposed into the elastic and the viscoplastic parts: l ¼ le þ lvp
ð17:199Þ
where 8 > < l FFe 1 vp vp le F Fe1 ; lp Fe F Fvp1 Fe1 ¼ Fe L Fe1 > vp : vp L F Fvp1
ð17:200Þ
Further, the velocity gradient L in the intermediate configuration is additively decomposed into the elastic and the viscoplastic parts as follows: e
vp
L ¼ L þL
ð17:201Þ
where
L Fe1 lFe vp e vp L Fe1 le Fe ¼ Fe1 Fe ; L Fe1 lvp Fe ¼ F Fvp1
ð17:202Þ
from which it follows that
L ¼ DþW e e e vp vp vp L ¼ D þW ;L ¼ D þW e
vp
D ¼ D þD ; where
e
W ¼ W þW
vp
8 < De¼ sym½L; e W ¼e ant½L e D ¼ sym L ; W ¼ ant L vp vp : vp vp D ¼ sym L ; W ¼ ant L
ð17:203Þ ð17:204Þ
ð17:205Þ
544
Multiplicative Hyperelastic-Based Plasticity with Subloading …
17
vp
The viscoplastic velocity gradient L is additively decomposed for the kinematic hardening and the elastic-core into the storage and the dissipative parts as follows: vp
vp
vp
L ¼ Lks þ Lkd ;
vp
vp
vp
L ¼ Lcs þ Lcd
ð17:206Þ
where 8 vp vp vp vp vp1 > < Lks Fks Fks ¼ Dks þ Wks
! _ vp vp vp vp vp1 vp _ vpG > : Lkd Fvp ¼ Dkd þ Wkd L F ¼ L kd G ks kd ks ks
ð17:207Þ
vp vp vp vp Dks sym Lks ; Wks ant Lks vp vp vp vp Dkd sym Lkd ; Wkd ant Lkd
_ vp
Lkd ¼
ð17:208Þ _
vp
Fkd Fvp1 kd
vp vp Fvp1 ks Lkd Fks
¼vp ks
vpG
!
Lkd G_
ð17:209Þ
vp
ð17:210Þ
8 vp vp vp vp > vp1 > < Lcs Fcs ¼ Fcs ¼ Dcs þ W! cs ; ! ^
^ vp
vpG
vp vp vp1 > ¼vp > cs Lcd G : Lcd Fcs Lcd Fcs
^ vp
vp
¼ Dcd þ Wcd
vp vp vp vp Dcs sym Lcs ; Wcs ant Lcs vp vp vp vp Dcd sym Lcd ; Wcd ant Lcd
Lcd ¼
vp
Fcd Fvp1 cd
vp vp Fvp1 cs Lcd Fcs
¼vp cs
vpG
Lcd G
ð17:211Þ ! ð17:212Þ
The limit subloading, dynamic loading, normal-yield, limit elastic-core, elastic-core and static-subloading surfaces are shown in Fig. 17.5. The viscoplastic strain rate is given by extending Eq. (14.29) as follows: vp
D ¼ CN
ð17:213Þ
The static subloading surface (Fig. 17.5) over which the viscoplastic strain rate is induced is given by f ðMs Þ ¼ Rs FðHÞ
ð17:214Þ
17.14
Subloading-Overstress Model
545 ˆ) ˆ( N N
Limit subloading surface M
M M M R
Mk
My
Ms
M
f (M) = cm Rs F ( H ) ˆ N c
Subloading surface
f (M) = RF ( H ) Normal-yield surface ˆ ) F (H ) f (M = Static subloading surface f (M s ) = RsF ( H )
Ms
Mc ˆ M c
M ks
Mk
0
M ij Elastic-core surface ˆ ) f (M c = c F (H )
Limit elastic-core surface ˆ ) f (M c = F (H )
ˆ M y
My
Ms
M s M ks
Mk
M /R Rs R M
Fig. 17.5 Limit loading, subloading, normal-yield, limit elastic-core, elastic-core and static-subloading surfaces in subloading-overstress model
analogously to Eq. (14.24), where Ms ¼ Ms Mks ¼
Rs M R
ð17:215Þ
noting Eq. (14.25). The evolution rule of the static normal-yield ratio RS is given based on Eq. (14.26) as
ð17:216Þ
with Eq. (14.27), and the time-integration for R c Cn= const. is given by Eq. (14.28) R vp with the replacement evp ¼ jjD jjdt.
546
17
Multiplicative Hyperelastic-Based Plasticity with Subloading … vp
vp
The dissipative parts Dkd and Dcd of the viscoplastic stain rate tensors for the kinematic hardening varialbe and the elastic-core are given by replacing the plastic
multiplier k to the viscoplastic multiplier C in Eqs. (17.96) and (17.98) as follows: vp
D ¼ CN 1 vp vpT Dkd ¼ Mk ð¼ Lkd Þ bk F vp Dcd ¼ CMcd
ð17:217Þ
The plastic spin and the spins for the kinematic hardening and the elastic-core
are given by replacing the plastic multiplier k to the viscoplastic multiplier C in Eq. (17.103) as follows: vp
vp
vp
vp
W ¼ gvp ant½M D ¼ gvp ant½M NC vp Wcd ¼ gvp k ant½Mk Dkd ¼ ½gk =ðbk FÞant½MMC ¼ O vp Wcd
¼
vp gvp k ant½Mk Dkd
ð17:218Þ
¼ gvp c ant½Mc Mcd C
vp where gvp ; gvp k and gc are the material constants. Then, the viscoplastic velocity gradient tensor and those for the kinematic hardening variable and the elastic-core are given from Eq. (17.105) with the
replacement of the plastic multiplier k to the viscoplastic multiplier C; respectively, as follows: vp
L ¼ CðN þ gvp ant½M NÞ vp
vp
Lkd ¼ Dkd ¼ CMk =ðbk FÞ vp Lcd
ð17:219Þ
¼ CðMcd þ g ant½Mc Mcd Þ p
The rates of the viscoplastic gradient tensors are given from Eq. (17.200)3 , (17.209) and (17.212) as follows: 8 vp vp vp > F > < F vp ¼ L _ vp vp1 vp vp vp Fkd ¼ Lkd Fvp ¼ F L F kd ks Fkd kd ks > > : vp ^ vp vp vp1 vp vp vp Fcd ¼ Lcd Fcd ¼ Fcs Lcd Fcs Fcd
ð17:220Þ
The numerical time-integration for these variables can be performed effectively by the tensor exponential method which are shown for the plastic deformation in Eq. (17.172).
17.14
Subloading-Overstress Model
547
Fig. 17.6 Stress–strain curve predicted by subloading-overstress model vp The storage parts Fe ; Fvp ks and Fcs of the deformation gradient are given by substituting the results of the time-integrations of Eq. (17.220) into Eqs. (17.196) and (17.197) as follows:
Fe ¼ FFvp1 ; e
_ vp
vp vp1 Fvp ks ¼ F Fkd ;
vp vp1 Fvp cs ¼ F Fcd
ð17:221Þ
^ vp
Further, C ; Cks and Ccs are calculated by substituting Eq. (17.221) into Eqs. (17.9) _
and (17.198). Furthermore, the stress S, the kinematic hardening variable Sk and the ^
e
_ vp
^ vp
elastic-core Sc are calculated by substituting C ; Cks and Ccs into the hyperelastic _p
^p
_ vp
^ vp
Eqs. (17.68) and (17.73) with the replacements of Cks and Ccs into Cks and Ccs . Furthermore, M, Mk and Mc are calculated by Eqs. (17.69), (17.76) and (17.77) with the replacement of the plastic strain rate to the viscoplastic one. The rate of isotropic hardening variable is generally described as follows:
H ¼ fHn ðM; H; NÞC
ð17:222Þ
The stagnation of isotropic hardening is incorporated by replacing the plastic
strain rate k to the viscoplastic strain rate C in the formulation described in Sect. 17.13. The stress–strain behavior predicted by the subloading-overstress model is illustrated in Fig. 17.6. As described in Sect. 14.5.3, the subloading-overstress model is no more than the generalization of the subloading surface model to the description of the viscoplastic deformation in the general rate and thus the original subloading surface model for the rate-independent elastoplastic deformation can be disused by using the subloading-overstress model.
548
17.14.2
17
Multiplicative Hyperelastic-Based Plasticity with Subloading …
Calculation Procedure
The numerical calculation is performed by the procedure described in the following. vp Fvp , Fvp kd , Fcd , H and Rs are updated analogously to Eq. (17.172) as follows:
h i n o 8 vp p vp vp > C F ¼ exp L Dt F ¼ exp N þ g ant M N Dt Fvp nþ1 nþ1 nþ1 nþ1 > nþ1 n n > nþ1 > > > > vp1 vp vp vp > Fvp > kd nþ1 ¼ exp Fks nþ1 Lkdnþ1 DtFks n þ 1 Fkdn > > > > vp1 vp > > ¼ expfFks > n þ 1 Cn þ 1 Dt½Mkdn þ 1 =ðbk F n þ 1 ÞgFkdn > > < vp vp vp1 vp vp Fcdnþ1 ¼ exp Fcs nþ1 Lcd nþ1 Dt Fcs nþ1 Fcdn > n o > > vp1 vp p > > Fvp ¼ exp F C Dt ðM þ g ant½M M Þ F n þ 1 cdn þ 1 cn þ 1 cdn þ 1 > c cs n þ 1 cs n þ 1 cdn > > > > > Hnþ1 ¼ Hn þ fHn nþ1 Cnþ1 Dt > > >
> > > 2 p hRs n Re i p Cnþ1 Dt 1 > : Rs nþ1 ¼ ð1 Re Þ cos cos Þ þ Re exp us p 2 1 Re 2 1 Re ð17:223Þ in which the viscoplastic multiplier Cn þ 1 is calculated by Eq. (14.34) or (14.35) or by the return-mapping projection. The stress, the kinematic hardening variable and the elastic-core are calculated by the procedure described at the end of the preceding subsection. It should be noticed that the calculation by the elastoplastic constitutive equation with the quite complex plastic modulus in Eq. (17.119) is not necessary even for the time-independent elastoplastic deformation analysis with the forward-Euler method by calculating it as a quasi-static deformation using the subloadingoverstress model.
Chapter 18
Viscoelastic-Viscoplastic Model of Polymers
Various polymers are used widely as the industrial materials, which exhibit not only the plastic deformation but also the large elastic deformation contrastively to metals. In addition, these deformations are of the time-dependence. The rigorous viscoelastic constitutive equation for the polymers has been formulated by Simo (1987a), Holzapfel and Simo (1996), Holzapfel (1996), etc. The concise explanation of the viscoelastic model for the polymers will be given without resorting to the thermodynamic interpretation in this chapter, referring to the well-known book by Holzapfel (2000). The viscoelastic model for the polymers will be extended to the multiplicative viscoplastic-plastic subloading surface model by simply adding the viscoplastic strain rate based on the subloading-overstress model in Chap. 17.
18.1
Viscoelastic Rheological Model
Polymers is the assembly of long polymer chains. Therefore, the deformation mechanism of polymers is quite different from that of hard solids which are the assembly of atoms such as metals, glasses and ceramics, etc. described in the preceding chapters. The former exhibits the entropic elasticity and the latter the energetic elasticity. Their main differences are shown in Table 18.1. The viscoelastic rheological model for the deformation of polymers is shown in Fig. 18.1, where the Prony series is adopted for the viscoelastic deformation. It is composed of parallel series of the one purely-elastic part represented by the spring E1 and the arbitrary number m of Maxwell models. The spring E1 ; E2 ; . . .; Em denotes the purely-elastic response part induced in the equilibrium state after the elapse of infinite time t ! 1, which is represented by the spring E1 . The Maxwell models describe the purely elastic deformation and the viscous deformation induced in the non-equilibrium state. Here, it is postulated that the former causes both the
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_18
549
550
18
Viscoelastic-Viscoplastic Model of Polymers
Table 18.1 Comparison of elastic properties of polymers and hard solids Microstructures
Polymers Assembly of molecular chains
Hard solids as metals Assembly of atoms
Cause of elastic deformation
Entropic elasticity Molecular chains move by micro-brown movements toward the maximum number of molecular shapes (most random) Stress is induced by entropy reduction, meanwhile internal energy does not change Large Large Low
Energetic elasticity Interatomic forces apply between atoms Stress is induced by increase of internal energy, meanwhile entropy does not change
Deformation
Total Elastic Elastic modulus Compressibility (Plastic) Influence of temperature increase under constant stress
Incompressible (Almost zero) Shrink by Brownian movements of molecules
Small Quite small High Compressible (Large) Stretch
Fig. 18.1 Viscoelastic rheological model
volumetric and the isochoric (volume-preserving) deformation but the latter causes only the isochoric deformation based on the experimental observation. The volumetric part and the isochoric part are designated by ð Þvol and ð Þiso , respectively.
18.2
Viscoelastic Deformation with Elastic Strain Energy Function
18.2
551
Viscoelastic Deformation with Elastic Strain Energy Function
The viscoelastic constitutive equation for the polymer will be described in this section.
18.2.1
Elastic Strain Free-Energy Function
Then, the free energy function in the isothermal viscoplastic process based on the rheological model shown in Fig. 18.1 is given by WðC ; C1 ; C2 ; Cm Þ ve
ve
¼W1 ðJ ve ; C Þ þ
m P a¼1
ve
ve
1 ve isoa ðC ; Ca Þ¼W1 vol ðJ Þ þ Wiso ðC Þ þ
m P
ve
isoa ðC ; Ca Þ
a¼1
ð18:1Þ where ð18:2Þ
F ¼ Fve Fvp Fve vol
J
ve1=3
ve Fve ¼ Fve vol F ; ve ve1=3 ve g; F J F ; J ve detFve
ð18:3Þ
ve
C FT F ¼ FvpT C Fvp ve
ð18:4Þ ve
veT
ve
C FveT Fve ¼ ðFFvp1 ÞT FFvp1 ¼ FvpT CFvp1 ; C ¼ Cvol C (
ve
ð18:5Þ
ve
1=3 ve Cvol FveT G vol Fvol ¼ ðdet C Þ ve ve ve C FveT Fve ¼ ðdet C Þ1=3 C
ð18:6Þ
and Ca ða ¼ 1; 2; . . .; mÞ are the internal variables denoting the viscoelastic deformation histories. The variables in the intermediate configuration K is designated by adding the over-bar ðÞ and the isochoric part by the under-bar ð Þ. g is the metric tensor in the current configuration. C is the right Cauchy-Green deforve mation tensor in the reference configuration. C is the viscoelastic right CauchyGreen deformation tensor and G is the metric tensor in the intermediate configuration.
552
18
Viscoelastic-Viscoplastic Model of Polymers
Here, note that the time-independent deformation represented by the spring proceeds in the equilibrium state ðt ! 1Þ, which is observed in the deformation under the infinitesimal strain rate, is decomposed into the volumetric and the isochoric deformations, so that the strain-energy function is composed of the voluve 1 ve metric part W1 vol ðJ Þ and the isochoric part Wiso ðC Þ. On the other hand, based on the experimental observation, it is postulated that the rate-dependent deformation behaviors represented by the Maxwell models are limited to the isochoric deformation (Holzapfel 1996), so that the free energy functions isoa ð a ¼ 1; 2; . . .; mÞ are the functions of the isochoric right Cauchy-Green viscoelastic deformation ve tensor C and the strain-like internal variables Ca ða ¼ 1; 2; . . .; mÞ, i.e. ve isoa ðC ; Ca Þ by which the viscoelastic deformation histories are described. The following normalization conditions are satisfied. W1 vol ð1Þ ¼ 0;
18.2.2
W1 iso ðGÞ ¼ 0;
m X
isoa ðG; GÞ ¼ 0
ð18:7Þ
a¼1
Second Piola–Kirchhoff Stress Tensor
The second Piola–Kirchhoff stress tensor S which is the work-conjugate to the right ve Cauchy-Green viscoelastic deformation tensor C is given from Eq. (18.1) by ve
S ¼ 2 @WðC
; C1 ; C2 ; ...; Cm Þ ve @C
1
¼S þ
m P a¼1
1
1
Qa ¼ Svol þ Siso þ
m P a¼1
Qa
ð18:8Þ
where 1
Svol ¼ 2
∞ =2 S iso
Qα = 2
ev ev @W1 dW1 vol ðJ Þ vol ðJ Þ ev1 ¼ J ev C ev dJ ev @C
∞ ∂Ψiso (Cve )
∂Cve
= J ve −2/3
ve
: [2
∞ ∂Ψ iso (Cve )
∂Cve
ð18:9Þ
]
ve ∂ϒ isoα (Cve , Γα ) ve −2/3 ve : [2 ∂ϒ isoα (C , Γα ) ] == J ∂Cve ∂Cve
ð18:10Þ
ð18:11Þ
18.2
Viscoelastic Deformation with Elastic Strain Energy Function
553
noting ffi pffiffiffiffiffiffiffiffiffiffiffiffi ev 1 ev ev @W1 ðJ Þ @W ðJ Þ @ detC vol vol ¼2 ¼2 ev ev @J ev @C @C 1 ev ev pffiffiffiffiffiffiffiffiffiffiffiffiffi @Wvol ðJ Þ 1 1 @W1 ev ev1 ev ev1 vol ðJ Þ ffiffiffiffiffiffiffiffiffiffiffiffi ffi p ðdetC ¼2 ÞC ¼ ð18:12Þ detC C ev ev ev @J 2 detC @J
1 Svol
∞ Siso =2
∞ ∂Ψiso (Cve )
∂Cve
=2
Qα = 2
=2
∞ ∂Ψiso (Cve ) ∂Cve : ve ∂Cve ∂C
∞ ∂Ψiso (Cve )
∂Cve
ve −1/3
: (detC )
( − 13 Cve ⊗ Cve −1)
∂ϒ isoα (Cve , Γα ) ∂ϒ isoα (Cve, Γα ) ∂Cve =2 : ve = Qα : J ve −2/3 ve T ve ∂C ∂Cve ∂C
ð18:13Þ
ð18:14Þ
where ve
ve
@ isoa ðC ; Ca Þ Qa ¼ 2 ve @C
ð18:15Þ
∂C = (det Cve )−1/3 veT = J ve −2/3 veT ∂Cve
ð18:16Þ
ve ≡ − 1 Cve ⊗ Cve
ð18:17Þ
−1
3
noting Eq. (4.63)2. Qa ða¼ 1; 2; . . .; mÞ in Eq. (18.8) are the stresses applied to the elements composed of the springs and the dash-pots, which are induced in the non-equilibrium state in general. Qa are called the non-equilibrium stresses, while they are the deviatoric stresses. is the fourth-order symmetrizing tensor, i.e. ijkl ≡ (δ ik δ jl + δ ilδ jk ) / 2 in Eq. (1.193). The Mandel stress M, which will appear later in the viscoplastic constitutive equation, is given by ve
M¼C S
ð18:18Þ
Here, in order to capture the simple interpretation for the variation of the relation between the involved variables with a time, suppose the one-dimensional
554
18
Viscoelastic-Viscoplastic Model of Polymers
generalized Maxwell model (see Fig. 18.1). Equation (18.8) is described in the one-dimensional deformation as follows: r ¼ r1 þ
m X
qa ¼ E 1 eve þ
a1
m X
ð18:19Þ
qa
a¼1
where qa which are the stresses applying to each Maxwell models satisfy the ve
following relations, designating the total viscoelastic strain rate by e .
qa qa þ e ¼ Ea ga
ð18:20Þ
qa ¼ Ea eve ¼ Ea eve sa
ð18:21Þ
ga Ea
ð18:22Þ
ve
i.e.
qa þ where
sa
which designates the relaxation time. The scalar quantity Ea eve can be regarded as the fictitious stress rate sa applied to the a-th Maxwell model calculated supposing that only the spring activates but the dashpot does not activate, i.e. sa Ea eve
ð18:23Þ
while the actual stress applied to each Maxwell model is qa as already mentioned above. Equation (18.21) is also derived by the different way as follows: ( qa ¼
Ea ðeev ca Þ
ga c a
:
ð18:24Þ
The upper equation yields
ev
qa ¼ Ea ðe ca Þ
ð18:25Þ
Substituting the lower equation in Eq. (18.24) to this equation, one obtains Eq. (18.21). (Note) Physical interpretation of the relaxation time: Consider the relaxation by applying the strain eve 0 at t ¼ 0 and keeping it for t 0, ¼ const: Then, Eq. (18.21) leads to resulting in Ea eve 0
18.2
Viscoelastic Deformation with Elastic Strain Energy Function
0 ¼ qa þ
555
qa dqa 1 ! ¼ dt sa qa sa
ðaÞ
The time-integration of this equation reads: lnð
qa t qa t t Þ¼ ! ¼ expð Þ ! qa ¼ qa0 expð Þ qa0 qa0 sa sa sa
ðbÞ
Eventually, sa is the time required that the stress is reduced to qa0 =e. The substitution of Eq. (18.23) into Eq. (18.21) leads to
qa þ
qa ¼ sa sa
ð18:26Þ
The following relation holds from Eq. (18.26). t t qa t qa þ exp sa ¼ exp exp sa sa sa sa
ð18:27Þ
which leads to ! # t qa ¼ exp sa
" exp
! t sa sa
ð18:28Þ
noting ðÞ ¼ dðÞ=dt. Further, this equation is time-integrated as exp
! T qa ðTÞ exp sa
0 sa
!
Z qa ð0Þ ¼
T
exp 0
! t sa dt sa
ð18:29Þ
leading to qa ðTÞ ¼ exp
! Z t T exp qa ð0Þ þ sa 0
ðT tÞ sa
!
sa dt
ð18:30Þ
Now, let the strain energy function in the quadratic form be assumed for the spring element, i.e. 1 2 w1 ðeve Þ ¼ E1 eve 2
ð18:31Þ
556
18
Viscoelastic-Viscoplastic Model of Polymers
for which one has r1 ¼
@w1 ðeve Þ ¼ E 1 eve @eve
ð18:32Þ
It follows from Eqs. (18.23) and (18.32) that sa ¼ ba E 1 eve ¼ ba r1 ¼ ba
@w1 ðeve Þ @eve
ð18:33Þ
where ba Ea =E 1
ð18:34Þ
Equation (18.26) is extended to the three-dimensional and nonlinear deformation regime as follows (Valanis 1972; Simo 1987a; Govindjee and Simo 1992; Holzapfel 1996; Holzapfel and Simo 1996; Holzapfel and Gasser 2001):
Qa þ
Qa ¼ Sisoa sa
ð18:35Þ
where Sisoa is the fictitious isochoric second Piola–Kirchhoff stress tensor extended from sa in the one-dimensional state, which is formulated explicitly by extending Eq. (18.33) to the three-dimensional state in terms of the second Piola–Kirchhoff stress tensor noting Eq. (18.10) as follows: ∞
Sisoα = β α S iso = 2 βα
∞ ∂Ψiso (C ve )
∂Cve
= β α J ve −2/3
ve
: [2
∞ ∂Ψ iso (Cve )
∂Cve
]
ð18:36Þ
which leads to 1
ve
Sisoa ¼ ba Siso ðC Þ
ð18:37Þ
Bearing Eq. (18.30) in mind and noting Eq. (18.37), we get the convolution integrals of Eq. (18.35) for some open-interval s 2 ð0; tÞ under the initial condition Qa ¼ Qa0 for T ¼ 0 as follows: Qa ¼ expð
T ÞQa0 þ sa
Z
1
T
expð 0
T t Þba Siso ðtÞdt sa
ð18:38Þ
18.2
Viscoelastic Deformation with Elastic Strain Energy Function
557
where Qα 0 = J 0ve −2/3 0ve: 2
∂ϒ isoα (C ve 0 , Γα 0 )
ð18:39Þ ∂C ve 0 noting Eq. (18.11), while the proof for Eq. (18.38) is omitted in Holzapfel (2000). Consequently, substituting Eqs. (18.9), (18.10) and (18.38) into Eq. (18.8), the stress is given as follows: S = J ev
∞ dΨ vol ( J ev )
−1
C ev + J ve −2 /3
dJ ev m
+∑
[
α =1
exp(− τT
α
ve
:[2 T
)Qα 0+ ∫0
∞ ∂Ψ iso (Cve )
∂Cve
]
•∞ − exp(− Tτ t )βα S iso(t )dt] α
ð18:40Þ
ve
1 ve The free energy functions W1 iso ðJ Þ and Wiso ðC Þ are given referring to Holzapfel (1996) as follows: ve W1 iso ðJ Þ ¼
ve
k 1 ðb ln J ve þ veb 1Þ J b2
W1 iso ðC Þ ¼
3 X N X lp A¼1
a p¼1 p
veap
ðkA
1Þ
ð18:41Þ
ð18:42Þ
where kve A is the isochoric part of the principal stretch of the viscoeastic deformation, noting Eq. (4.67), and k (bulk modulus), b, ap and lp (shear moduli) are the material constants. The incompressible Mooney model in Subsection 7.3.3 was adopted for the explicit strain energy function in the long-term response part and the two element Maxwell models were adopted for the viscoplastic parts. Then, increasing the amplitude as A ¼ 0:01 ðmÞ, A ¼ 0:02 ðmÞ, A ¼ 0:03 ðmÞ and so on by each cycle to the cube of 0.1 (m) and the calculated result of the Cauchy stress was shown by Holzapfel (1996) using the material constants l1 ¼ 6:30 105 (N/m2 Þ, l2 ¼ 0:012 105 (N/m2 Þ, l3 ¼ 0:10 105 (N/m2 Þ and a1 ¼ 1:3, a2 ¼ 5:0, a3 ¼ 2:0. The similar result was calculated using the Mooney model trough the subroutine of the Abaqus FEM program by Ishikawa (2015).
18.3
Viscoelastic-Damage Model: Subloading-Mullins Effect
Let the viscoelastic-damage model be formulated by combining the viscoelastic model and the damage model in the following, while the damage of polymer materials is induced by the rupture of polymer chains by tension and called the Mullins effect since it was found by Mullins (1947, 1969).
558
18
Viscoelastic-Viscoplastic Model of Polymers
Fig. 18.2 Evolution of damage variable
The elastic strain energy function in Eq. (18.1) is extended to the viscoelasticdamage model as follows: ve
ev
ve 1 Wf ðC ; f;C1 ; C2 ; . . .; Cm Þ ¼ W1 vol ðJ Þ þ ð1 fÞ½Wiso ðC Þ m X ve þ isoa ðC ; Ca Þ
ð18:43Þ
a¼1
where f is the damage variable given by Miehe (1995) simplifying the formulation of Simo (1987a) as follows (Fig. 18.2): fðcd Þ ¼ f1 ½1 expðcd =tÞ 0
f ðcd Þ ¼
@fðcd Þ c ¼ f1 d exp t @cd
cd t
ð18:44Þ ! ð18:45Þ
where f1 denotes the maximum value of the Mullins-damage variable f, and t is the material constant relating to evolution speed of the Mullins-damage variable. ve The variable cd evolves when C lies on the Mullins-damage surface defined by ve
UðC Þ cd ¼ 0
ð18:46Þ
and the surface expands so that the evolution rule is given by ve
@UðC Þ cd ¼ UðC Þ ¼ :C ve @C
ve
ve
[0
ð18:47Þ
The varaiable cd is induced discontinuosly when it reaches the Mullins-damage surface and in addition the input-incremental step must be taken to be infinitesimal in the numerical calculation such that cd does not go out finitely from the Mullins-damage surface. In what follows, the revised evolution rule of the variable cd in which these shortcomings in the past formulation are excluded will be formulated by incorporating the subloading concept.
18.3
Viscoelastic-Damage Model: Subloading-Mullins Effect
(Cve ) Cve ve C
C ve
0
559
Normal - damage evolution surface (Cve ) d Subloading - damage evolution surface (Cve ) R d d
C ijve
Fig. 18.3 Normal- and subloading-damage surfaces
The surface defined in Eq. (18.46) is renamed as the normal-Mullins damage surface and the subloading-Mullins damage surface ve
ve
UðC Þ Rd cd ¼ 0 ! Rd ¼ UðC Þ=cd
ð18:48Þ
are incorporated (Fig. 18.3), where Rd ð0 Rd 1Þ is called the normal-Mullins ve damage ratio designating the approaching degree of C to the normal-Mullins damage surface in Eq. (18.46). The consistency condition for the subloadingMullins surface is given by ve
@UðC Þ ve C Rd cd cd Rd ¼ 0 @C ve
ð18:49Þ
Then, assume the following evolution rule of the variable cd . ve • • γ d = R dn 〈Φ (Cve )〉 = R dn 〈 ∂Φ (Cve ) : C• ve 〉 ∂C
ð18:50Þ
where nð 1Þ is the material constant. The rate of Rd is given from Eq. (18.49) with Eq. (18.50) as follow: ve • Φ (Cve ) • ve n +1 ∂Φ (C ) • ve 1 R d = γ1 [ ∂ ve : C − R d 〈 ve : C 〉 = γ d d ∂C ∂C
ve
〈 ∂Φ (Cve ∂C
)
•
: Cve 〉 (1 − R dn+1)
Φ (Cve ) • ve for ∂ ve : C ≥ 0 ∂C ð18:51Þ
560
18
Viscoelastic-Viscoplastic Model of Polymers
which is the monotonically-increase function of Rd fulfilling
⎧ 1 ∂Φ (Cve ) • ve : C 〉 for Rd = 0 ⎪= γ 〈 ∂Cve ⎪ d • ⎪ ⎪ (Cve ) • R d ⎨< 1 〈 ∂Φ ve : Cve 〉 (> 0) for Rd < 1 γ ∂C ⎪ d ⎪= 0 for Rd = 1 ⎪ ⎪⎩< 0 for Rd > 1
ð18:52Þ
Therefore, the normal-Mullins damage surface evolves automatically so as to ve envelope the variable C always. The smooth transition from the non-evolution state to the normal-evolution state of damage can be described by incorporation of the above-mentioned formulation based on the subloading surface concept as shown in Fig. 18.4.
Stress
Strain Normal - damage evolution state Non - damage evolution Smooth transion from non - damage evolution to normal damage evolution by subloading surface concept Fig. 18.4 Stress-strain curve with Mullins effect
18.3
Viscoelastic-Damage Model: Subloading-Mullins Effect
561
ve
Besides, the Mullins-damage function UðC Þ was given by Holzapfel (2000) as follows: ve
ve
ve
UðC Þ ¼ C1 ½Ic ðC Þ 3 þ C2 ½IIc ðC Þ 3 ve
ve
¼ C1 ðtrC 3Þ þ C2 ½ð1=2Þðtr2 C trC
ve2
Þ 3Þ
ð18:53Þ
in the Mooney-Rivlin type in Eq. (7.60), where C1 and C2 are the material constants, noting Eq. (4.70). The stress S for Eq. (18.43) is given as follows: m X @WD ðC ; f; C1 ; C2 ; . . .; Cm Þ 1 ve 1 ve ¼ S ðJ Þ þ ð1 fÞ½S ðC Þ þ Qa ve iso vol @C a¼1 ve
S¼2
ð18:54Þ nothing Eq. (18.8), which is rewritten by Eq. (18.40) as follows: S = J ev
∞ ( J ev ) dΨ vol
dJ ev m
{
−1
C ev + (1 − ζ ) J ve −2/3
[
+ ∑ exp(− τT α =1
18.4
α
ve
:[2
∞ (Cve ) ∂Ψ iso
∂Cve
]
•
∞ (t )dt]} )Qα 0+ ∫0 exp(− Tτα− t ) βα S iso T
(18.55)
Viscoplastic Constitutive Equation in Glassy State
The viscoplastic deformation is induced in addition to the viscoelastic deformation in the temperature lower than the glass transition temperature, leading to the glassy state. The viscoplastic deformation can be described by the subloading-overstress model in Sect. 17.14. The rheological model for the viscoelastic-viscoplastic model is shown in Fig. 18.5. The deformation in the glassy state can be calculated by the following procedure. (1) The voscoplastic deformation gradient Fvp is calculated by the method described in Sect. 17.14.2 and then the viscoelastic gradient Fve ¼ FFvp1 is calculated. (2) The second-Piola Kirchhoff stress tensor S is calculated by substituting J ve ¼ ve det Fve and C ¼ Fve FveT ðFve J ve1=3 Fve Þ into Eq. (18.40). The stress versus stain curves for polymer at different temperatures are schematically shown in Fig. 18.6, exhibiting the facts: (1) The deformation becomes brittle, (2) the plastic deformation is induced, (3) the volumetric strain is induced and (4) the upper yield point appears exhibiting the glassy behavior with the temperature lowering. The constitutive equation in the glassy state has been
562
18
Viscoelastic-Viscoplastic Model of Polymers
Fig. 18.5 Viscoelastic-viscoplastic rheological model
Fig. 18.6 Variation of stress versus strain curve with temperature decrease in polymer
studied by Boyce et al. (1988), Anand and Gurtin (2003), Richeton et al. (2005), Ward and Sweeney (2013), Bergstrom (2015), Matsubara et al. (2015), etc. In the glassy polymer, firstly prior to the initial yield, the intermolecular resistance to segment rotation exhibiting the peak stress followed by the strain softening must be exceeded. The intermolecular deformation is regarded as the rate-independent behavior. Then, once the material begins to flow, the molecular alignment occurs, altering the configurational entropy and exhibiting the inherent deformation behavior of rubbers. The viscoplastic constitutive equation of the polymer in the glassy state would be given by the modification of the viscoplastic multiplier in Eq. (14.34), referring
18.4
Viscoplastic Constitutive Equation in Glassy State
563
to Richeton et al. (2005) based on the Eyring equation (Eyring 1935, 1936), as follows (see Appendix K): 〈 R − Rs 〉 n ) T Γ ≡μ v exp(u cR cCn ) [1 − [ R / (c m R s )] sinh(
ð18:56Þ
by which a viscoplastic relaxation is induced at accelerated rate as the overstress increases, since sinh x increases with the increase of x.
Chapter 19
Corotational Rate Tensors
It was studied in Chap. 6 that the material-time derivatives of state variables, e.g. stress and internal variables in elastoplasticity, do not possess the objectivity and thus, instead of them, we must adopt their objective time-derivatives. The responses of simple constitutive equations introducing corotational rates with various spins including the plastic spin will be examined in this chapter.
19.1
Hypoelasticity
Consider the hypoelastic constitutive equation in Eq. (7.110), i.e.
r ¼ K ðtrdÞI þ 2Gd0
ð19:1Þ
Equation (19.1) is described by the following equation for Eq. (3.43), noting r12 ¼ r21 ; wS12 ¼ wS21 and using Eq. (5.83)1 with Eq. (5.79), for the simple shear deformation described in Sect. 5.10.2. "
r11 2r12 ws12 Sym:
r12 þ ðr11 r22 Þws12 r22 þ 2r12 ws12
#
0 ¼G 1
2 0 1 1 h c¼G 0 cos2 h 1 0 ð19:2Þ
where the spin of material is represented by the symbol ws as shown in Eq. (3.43).
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_19
565
566
19.1.1
19
Corotational Rate Tensors
Zaremba-Jaumann Rate
When the Zaremba-Jaumann rate in Eq. (3.40) is adopted for the corotational rate, Eq. (19.2) leads to the following equation by setting ws ¼ w with Eq. (5.83)2. 2
4 r11 c r12 sym:
3 c r12 þ ðr11 r22 Þ 5 ¼ G 0 2 1 r22 þ c r12
from which we have
9 > > > > =
r11 c r12 ¼ 0
c r12 þ ðr11 r22 Þ ¼ G c > > 2 > > ; r22 þ c r12 ¼ 0
1 c 0
ð19:3Þ
ð19:4Þ
Substituting
r22 ¼ r11 ; c ¼
r11 r12
ð19:5Þ
obtained from the first and the third equations into the second equation in Eq. (19.4), yields
r12
r11 r11 þ r11 ¼ G r12 r12
ð19:6Þ
the time-integration of which is given as r12 ¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2Gr11 r211
ð19:7Þ
Substituting this equation into the second equation of Eq. (19.4), we have
r11 G sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2ffi ¼ c r11 1 1 G the integration of which is given by r11 cos1 1 ¼c G
19.1
Hypoelasticity
567
i.e. r11 ¼ r22 ¼ Gð1 coscÞ
ð19:8Þ
The substitution of Eq. (19.8) into Eq. (19.7) leads to r12 ¼ G sin c
ð19:9Þ
The continuum spin w designates the instantaneous rate of rotation of the principal directions of strain rate, i.e. the instantaneous rate of rotation of a cross depicted momentarily on the material surface. Therefore, if it is used in the simple shear
deformation with the constant shear strain rate, i.e. c ¼ const: leading to w ¼ const:, the material is regarded to rotate in a constant angular velocity, while the strain rate d is also kept constant. Then, the oscillatory shear stress is predicted by the hypoelastic constitutive equation using the Zaremba-Jaumann rate with the continuum spin as shown in Eq. (19.9) and depicted in Fig. 19.1 (cf. e.g. Dienes 1979).
19.1.2
Green-Naghdi Rate
Consider the Green-Naghdi rate for the corotational rate with the relative spin wS ¼ XR , i.e. Equation (3.39). It follows from Eq. (5.89) that 2 0 1 0 c¼ w ¼ X ¼ RR ¼ 1 4 þ c2 1 0 S
R
T
1 h 0
ð19:10Þ
3.0
12 /G 2.0
Green-Naghdi rate by ΩR
1.0
0.0
Jaumann rate by w
/2
– 1.0
Fig. 19.1 Description of simple shear deformation of hypoelastic material by Jaumann rate and Green-Naghdi rate (Dienes 1979).
568
19
Corotational Rate Tensors
The substitution of Eq. (19.10) into Eq. (19.2) reads: 2
4 r11 2r12 h
3
r12 þ ðr11 r22 Þ h 5
r22 þ 2r12 h
sym: from which we have
2 0 ¼G cos2 h 1
9 > > > > > = 2 þ ðr11 r22 Þ h ¼ G h cos2 h > > > > > ; þ 2r12 h ¼ 0
1 h 0
ð19:11Þ
r11 2r12 h ¼ 0
r12
r22 It is obtained that
r22 ¼ r11
1 r11 h¼ 2 r12
ð19:12Þ
! ð19:13Þ
from the first and the third equations, and dr11 dr12 1 d 2 r11 ¼ 2r12 ! ¼ 2 dh2 dh dh
ð19:14Þ
from the first equation in Eq. (19.12). Substituting Eqs. (19.13) and (19.14) into the second equation in Eq. (19.12), i.e. dr12 =dh þ 2r11 ¼ 2G= cos2 h, we have the ordinary differential equation d 2 r11 dh
2
þ 4r11 ¼
4G cos2 h
ð19:15Þ
The roots of the characteristic equation of the second-order homogeneous linear differential equation for Eq. (19.15) are given by pffiffiffiffiffiffiffiffiffi 16 ¼ 2i m þ 4m ¼ 0 ! m ¼ 2 2
Thus, the complementary function of Eq. (19.15) is given by the following equation. pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r11 ¼ Acos2h þ Bsin2h ¼ A2 þ B2 sin½2h þ arctanðA=BÞ ð19:16Þ where A and B are the integral constants. Further, adding the particular solution for Eq. (19.15) itself, the general solution of Eq. (19.15) is obtained as follows:
19.1
Hypoelasticity
r11 ¼
569
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi A2 þ B2 sinð2h þ arctanðA=BÞÞ þ 4Gðcos2h ln cosh þ hsin2h sin2 hÞ ð19:17Þ
Assuming that the initial stress is zero, Eq. (19.17) becomes r11 ¼ 4Gðcos2h ln cosh þ hsin2h sin2 hÞ
ð19:18Þ
Furthermore, substituting Eq. (19.18) into Eq. (19.14), we have r12 ¼
1 dr11 ¼ 2Gcos2hð2h 2tan2 h ln cosh tanhÞ 2 dh
ð19:19Þ
These equations have been derived by Dienes (1979). The relative spin XR designates the mean rate of rotation of the cross depicted on the material surface at the beginning of deformation. Therefore, it coincides with the continuum spin w at the initial state but it decreases gradually with the shear deformation. Then, the oscillation of shear stress observed in Jaumann rate is not predicted if the Green-Naghdi rate is adopted as the corotational rate as seen in Fig. 19.1 calculated by Eq. (19.19).
19.2
Kinematic Hardening Material
For the sake of simplicity, consider the response of a rigid plastic material fulfilling d ¼ dp and assume the Mises material with linear kinematic hardening in Eq. (8.86), i.e. 2 2 ^kdp0 k¼ ha dp0 a ¼ ha n 3 3
ð19:20Þ
setting ck ¼ ð2=3Þha , which was analyzed by Dafalias (1983). Then, it holds for the simple shear that " a11
a21
a12 a22
#
" 0 1 ¼ ha 3 1
1 0
#
ð19:21Þ
c
by substituting Eq. (5.83)1 into Eq. (19.20). Equation (19.21) is described for Eq. (3.43) as follows: " a11 2a12 ws12 Sym:
a12 þ ða11 a22 Þws12 a22 þ 2a12 ws12
#
1 0 ¼ ha 1 3
1 c 0
ð19:22Þ
570
19
Corotational Rate Tensors
where ws ¼ ( zðcÞ ¼
0 1 zðcÞ c 1 0
ð19:23Þ
for ws ¼ w
1=2
ð19:24Þ
2=ð4 þ c2 Þ for ws ¼ XR (
z ðcÞ ¼
for ws ¼ w
0
0
ð4 þ4cc2 Þ2
ð19:25Þ
for ws ¼ XR
noting Eqs. (5.83)2 and (19.10). The substitution of Eq. (19.23) into Eq. (19.22) leads to "
a11 2a12 zðcÞ c Sym:
a12 þ ða11 a22 ÞzðcÞ c a22 þ 2a12 zðcÞ c
#
1 0 ¼ ha 1 3
1 c 0
ð19:26Þ
from which we have 9 > a11 2a12 zðcÞ c ¼ 0 = 1 a12 þ ða11 a22 ÞzðcÞ c ¼ 3 ha c > ; a22 þ 2a12 zðcÞ c ¼ 0
ð19:27Þ
In addition, noting a11 ¼ a22 , we have 0 0 a11 ¼ a22 ¼ 2zðcÞa12
)
a012 þ 2zðcÞa11 ¼ 13 ha
ð19:28Þ
where ð Þ0 ¼ dð Þ=dc. Differentiating Eq. (19.28), we have 9 00 0 2za12 2a12 z 0 ¼ 0 = a11 00 0 þ 2a11 z þ 2a11 z 0 ¼ 0 ; a12
which, noting Eq. (19.28), becomes 9 0 a11 > 0 z ¼ 0> 2z ha 2za11 2 > = 2z 3 > > 1 1 00 0 ; a12 þ 4za12 z þ 2 ha a12 z0 ¼ 0 > 2z 3 00 a11
1
ð19:29Þ
19.2
Kinematic Hardening Material
571
Then, it is obtained that
9 z0 0 2 > 2 a11 þ 4z a11 ha z ¼ 0 > = z 3 > z0 0 1 z0 00 ; a12 þ 4z2 a12 þ ha ¼ 0 > a12 z 3 z 00 a11
19.2.1
ð19:30Þ
Zaremba-Jaumann Rate
The substitution of Eqs. (19.24)1 and (19.25)1 into Eq. (19.30) leads to 9 1 00 a11 þ a11 ha ¼ 0 = 3 00 ; a12 þ a12 ¼ 0
ð19:31Þ
from which, noting the initial condition a11 ¼ 0; a12 ¼ 0 for c ¼ 0, we have 9 1 = a11 ¼ a22 ¼ ha ð1 coscÞ > 3 1 > ; a12 ¼ ha sinc 3
ð19:32Þ
It is obtained from Eq. (19.32) that 9 1 > r11 ¼ a11 ¼ r22 ¼ a22 ¼ ha ð1 cos cÞ > = 3 1 1 1 > > r12 ¼ pffiffiffi F þ a12 ¼ pffiffiffi F þ ha sin c ; 3 3 3 noting
ð19:33Þ
ffi pffiffiffi pffiffiffiffiffiffiffiffiqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3=2 ðr11 a11 Þ2 þ ðr22 a22 Þ2 þ 2ðr12 a12 Þ2 ¼ 3ðr12 a12 Þ ¼ F
^k ¼ 0 ðno sum; i ¼ 1; 2Þ. Both r11 and r12 with dii ¼ diip ¼ k ^nii ¼ kðrii aii Þ= kr oscillates in sine curves as shown in Fig. 19.2. The above-mentioned fact that the kinematic hardening model with the Zaremba-Jaumann rate exhibits the oscillation was indicated first by Nagtegaal and de Jong (1982).
19.2.2
Green-Naghdi Rate
Substituting Eqs. (19.24)2 and (19.25)2 into Eq. (19.30) for the relative spin adopted in the Green-Naghdi rate, we have
572
19
12
Green-Naghdi rate by Ω R
Jaumann rate by w
F/ 3 ha /3
F/ 3 F/ 3 ha /3 0
Corotational Rate Tensors
11 2ha /3
(a)
2
0
(b)
2
12 ha / 3
0
ha / 3
2
(c)
Fig. 19.2 Description of simple shear deformation of kinematic hardening material by Zaremba-Jaumann and Green-Naghdi rates (Dafalias 1983)
9 > > > 2 Þ2 > 4 2 2 ð4 þ c > 00 0 > > h a11 a11 þ 4 a ¼ 0 11 a > 2 2 > 2 2 3 4þc ð4 þ c Þ > > = 2 4þc 4c
00 a12
4c
4c
ð4 þ c2 Þ2
ð4 þ c2 Þ2
2 4 þ c2
0 a12 þ4
4 ð4 þ c2 Þ2
a12 þ
1 ha 3
2 4 þ c2
> > > > > > > ¼ 0> > > > ;
i.e. 9 2c 0 16 4 > > ¼ 0 a þ a h > 11 a = 4 þ c2 11 ð4 þ c2 Þ2 3ð4 þ c2 Þ 2c 16 2c > 00 > a12 ha ¼ 0 > þ a0 þ a12 ha ; 4 þ c2 12 ð4 þ c2 Þ2 3ð4 þ c2 Þ 00 a11 þ
ð19:34Þ
19.2
Kinematic Hardening Material
573
the general solution of which is derived as the following equation by the method of variable coefficients (Dafalias 1983). 9 " # ! > 1 1 c 2 > > > 4cf4 tan1 a11 ¼ ha cg 4ðc2 4Þ ln pffiffiffiffiffiffiffiffiffiffiffiffi > 2 > 2 3 4þc 2 = 4þc ! " !# > > > 1 1 c 2 > 3 2 1 > ffiffiffiffiffiffiffiffiffiffiffiffi p 4ðc 4Þ4 tan 4c 1 þ 4 ln c a12 ¼ ha > ; 2 2 3 4þc 2 4þc ð19:35Þ The relation of r11 ; r12 to a11 ; a12 is given by Eq. (19.33) also in this case. An oscillation is not predicted in the simple shear deformation as shown in Fig. 19.2.
19.3
Plastic Spin
The above-mentioned Zaremba-Jaumann rate and the Green-Naghdi rate do not reflect the substructure of material but they are uniquely determined only by the change of external appearance of the material. However, the mechanically meaningful rotation would be the spin of substructure, as known presuming the crystals of metals or the annual ring of woods, which would be the rotation of the principal direction of anisotropy (Kratochvil 1971; Mandel 1971). The concept of the plastic spin is proposed in order to incorporate such rotation into elastoplastic constitutive equations (Dafalias 1983, 1985a, b; Loret 1983). In what follows, in order to interpret the mechanical meaning of the plastic spin, assume the rigid plasticity and the simplest anisotropy, i.e. the traverse anisotropic material (Fig. 19.3) with the parallel line-elements of substructure having the direction ^e1 inclined p=4 from the fixed base e1 in the initial state of deformation and rotates by the angle u in the clockwise direction with the increase of shear strain (Dafalias 1984). Then, it holds that ( ) ( ) cosðp=4 uÞ sinðp=4 uÞ _ _ ð19:36Þ e1 ¼ ; e1 ¼ u sinðp=4 uÞ cosðp=4 uÞ Here, referring to Fig. 19.3, one has tanðp=4 uÞ ¼
1 1þc
ð19:37Þ
from which one has
u c ¼ ¼ c tan2 ðp=4 uÞ 2 2 cos ðp=4 uÞ ð1 þ cÞ
ð19:38Þ
574
19
Corotational Rate Tensors
d1
d2
d2
d1
w3
(= w12 )
a
w 3s
(= ws12 )
tan 1
a
e2
e2
/4
e1
e1
a
w12s
w12 /2
2
w12p (d11, w12s ( w12 w12p ))
d2 = /2
d 1 = /2
0
d11
Fig. 19.3 Substructure spin in traverse isotropic material
Then it holds that
c c u ¼ c sin ðp=4 uÞð¼ ½1 cosf2ðp=4 uÞgÞ ¼ ð1 sin 2uÞ 2 2
2
ð19:39Þ
Using Eq. (19.39) along with Eq. (19.36), it is obtained that
_
e1 ¼
(
sinðp=4 uÞ cosðp=4 uÞ
)
ð1 sin 2uÞ
c 2
ð19:40Þ
19.3
Plastic Spin
575
which is rewritten as (
sinðp=4 uÞ
)
c ð1 sin 2uÞ ¼ 2 cosðp=4 uÞ " 0 ð c =2Þ sin 2u
"
c =2
0
c =2
#
0
ð c =2Þ sin 2u 0
#!(
cosðp=4 uÞ
)
sinðp=4 uÞ
i.e.
_
_
e1 ¼ ws e1
Then, it can be confirmed from Eqs. (5.82)2 that Eq. (3.44) for the spin of substructure holds as follows: " 0 w ¼w ¼ww ¼ c =2 sin 2u 1 0 p c w ¼2 sin 2u 0 s
e
p
# " 0 c =2 ðc =2Þ sin 2u 0
#9 ðc =2Þ sin 2u > > > = 0 > > > ; ð19:41Þ
as illustrated in Fig. 19.3. Kuroda (1997) applied the above-mentioned formulation for the traverse isotropic material to the orthotropic material described in Eq. (12.54) in Sect. 12.7 and showed the numerical calculation results for the rotation of the principal axes of orthotropic yield surface. Next, consider the same problem by the deformation of metal crystals. If the substructure does not rotate, it holds for the slip system in Fig. 19.4 that
v ¼ cðx nÞs ¼ cðs xÞn; vi ¼ cðxr nr Þsi lp ¼
@v @vi ¼ c djr nr si ¼ c si nj ¼ c s n; lpij ¼ @xj @x
1 1 d ¼ cðs n þ n sÞ; dijp ¼ 2 2
@vi @vj þ @xj @xi
1 1 wp ¼ cðs n n sÞ; wpij ¼ 2 2
@vi @vj @xj @xi
p
ð19:42Þ ð19:43Þ
! 1 ¼ cðsi nj þ ni sj Þ 2
ð19:44Þ
1 ¼ cðsi nj ni sj Þ 2
ð19:45Þ
!
Equations (19.42)–(19.45) are extended for multi slip systems of number n as follows:
576
19
v
.
Corotational Rate Tensors
x n
s
Fig. 19.4 Slip system and slip deformation
lp ¼
n X cðaÞ sðaÞ nðaÞ
ð19:46Þ
a¼1
"
n X d ¼ cðaÞ p
a¼1
wp ¼
# n X 1 ðaÞ ðaÞ ðaÞ ðaÞ cðaÞ pðaÞ ðs n þ n s Þ ¼ 2 a¼1
ð19:47Þ
1 pðaÞ ðsðaÞ nðaÞ þ nðaÞ sðaÞ Þ 2
ð19:48Þ
n n X X 1 ðaÞ ðaÞ c ðs nðaÞ nðaÞ sðaÞ Þ ¼ cðaÞ qðaÞ 2 a¼1 a¼1
1 qðaÞ ðsðaÞ nðaÞ nðaÞ sðaÞ Þ 2
ð19:49Þ
ð19:50Þ
The simple example of the plastic spin is shown above. Dafalias (1985a, b) provided the general mechanical interpretation of the plastic spin based on Eq. (3.44) as follows. The substitution of Eq. (3.44) into Eq. (3.43) reads:
w
t ¼ t ðw wp Þt þ tðw wp Þ ¼ t þ wp t twp
ð19:51Þ
The relation of the corotational rate and the Zaremba-Jaumann rate of Cauchy stress is given from Eq. (8.94) and (19.51) as
r ¼ rw þ w p r rw p
^rÞr rðr^ ^rÞg ¼ rw þ g p kfðr^ nn nn
19.3
Plastic Spin
577
i.e.
r ¼ rw þ k rn
ð19:52Þ
^ r2 r 2 n ^Þ nr n rn gp ð2r^
ð19:53Þ
where
Now, we derive the elastoplastic constitutive equation. Substituting Eq. (19.52) into Eq. (8.81), it follows that
^ :r ^ :ðrw þ k rn Þ n n ¼ k¼ Mp Mp
ð19:54Þ
from which it is obtained that
k¼
^ : rw p ^ : rw n n ^ ; d ¼ n ~p ~p M M
ð19:55Þ
~ p ¼ M p ^ M n :rn
ð19:56Þ
where
The substitution of Eq. (19.55) into Eq. (19.52), one has
^ : rw n rn r¼r þ ~p M
w
ð19:57Þ
Then, the strain rate is given by
^ : rw ^ : rw n n ^ r Þ þ n d ¼ E : r þ d ¼ E :ðr þ n ~p ~p M M 1
p
1
w
leading to
^ : rw n ð^ n þ E1 :rn Þ d ¼ E :r þ ~p M w 1
ð19:58Þ
The plastic multiplier K in terms of strain rate is given by Eq. (8.83) as it is since Eq. (8.81), i.e. (19.54) holds even in the present formulation. Then, the
Zaremba-Jaumann rate of Cauchy stress is given from Eq. (19.52) with r ¼ E : de by
578
19
Corotational Rate Tensors
rw ¼ E : ðd dp Þ K rn ^ :E :d ^ :E :d n n ^: p rn ¼ E :d E :n ^:E:n ^ Mp þ n ^:E:n ^ M þn i.e.
rw ¼ ½E
^ þ rn Þ ð^ ðE : n n : EÞ : d p ^:E:n ^ M þn
ð19:59Þ
which is related by the non-symmetric tangent modulus tensor. The stress and the kinematic hardening variable are updated by the time-integrations of
r ¼ rw þ wr rw
ð19:60Þ
a ¼ jjdp jjf kn þ ðw wp Þa aðw wp Þ
ð19:61Þ
noting Eqs. (19.51), with a ¼ k f k^n ¼ jjdp jjf k^n base on Eq. (8.89) Hereinafter, limit to the Mises yield condition with the kinematic hardening.
Then, substituting Eq. (8.75) ðep ! dp Þ with Eq. (12.2) into Eq. (8.94), the plastic spin is reduced to the following equation.
^rÞ nn wp ¼ gp kðr^ 0 ^ ^0 ^0 ^0 r r r r r þ aÞg r þ aÞ 0 0 ð^ ¼ gp kðr 0 0 rÞ ¼ gp kfð^ jj^ r jj jj^ r jj jj^ r jj jj^ r jj
ð19:62Þ
¼ gp kðar 0 r 0 aÞ=jj^ r 0 jj which was first proposed by Dafalias (1985a, b), where Eq. (8.94) is regarded as the generalized form of Eq. (19.62) which is limited to the Mises material with the kinematic hardening. Then, considering the simple shear deformation and assuming the rigid plasticity, the initial isotropy and the linear kinematic hardening as in Sect. 19.2, Eq. (19.62) leads to w ¼g p
p
¼ gp
a11 a12 a12 a11
! 1 c 0 1 c a11 a12 1 0 2 a12 a11 a11 1 0 2 a11 a11 a12 a11 c p 0 ¼g c 2 a12 a11 a12 a11 0 a12
0
Substituting Eqs. (5.83), (19.21) and (19.63) into Eq. (19.61), we have
ð19:63Þ
19.3
Plastic Spin
" a11
a12
a12
a11
#
579
0 1 c 2 ¼ ha 3 1 0 2 ! 0 1 0 1 c a11 a12 p c g a11 þ a12 a11 1 0 1 0 2 ! 0 1 c 0 1 a11 a12 p g a11 c 1 0 2 1 0 a12 a11 " # " # a12 a11 c a12 a11 0 1 1 p cþ ¼ ha g a11 c 3 1 0 a11 a12 2 a11 a12 " # " # a12 a11 a12 a11 c p þ g a11 c a11 a12 2 a11 a12 " # " # a12 a11 a a 0 1 12 11 1 cþ ¼ ha c 2gp a11 c 3 1 0 a11 a12 a11 a12 2 3 1 p p h a Þa ð1 2g a Þa ð1 2g 11 12 a 11 11 7 6 3 7 c ¼6 41 5 ha ð1 2gp a11 Þa11 ð1 2gp a11 Þa12 3
from which we obtain 9 da11 > > ¼ ð1 2gp a11 Þa12 = dc da12 1 > ; ¼ ha ð1 2gp a11 Þa11 > dc 3
ð19:64Þ
The nonlinear differential Eq. (19.64) is solved numerically by Dafalias (1985a). The calculation result is shown in Fig. 19.5. As seen in this figure, the non-oscillation curve is obtained by choosing the material parameter gp appropriately. When choosing gp [ 0:25, r11 ¼ r22 and r12 increase monotonically with the increase of shear strain c. On the other hand, the Zaremba-Jaumann and the Green-Naghdi rates are independent of material property and thus they would lack the physical exactness. Oscillatory shear stress is predicted in the simple shear by the hypoelasto-plastic material with the kinematic hardening if the Jaumann rates are adopted for the stress and the back stress as described above. In the plastically-isotropic material, however, it should be noted that the plastic strain rate is independent of the types of stress rate, i.e. material-time derivative, Green-Naghdi rate, Zaremba-Jaumann rate, etc. as known from Eq. (8.81) because of
580
19
Corotational Rate Tensors
4
12
p = 0.5
3
3
11
p = 0.25 p = 0.21
1
p =0 0 0
2
p = 0.25 p = 0.21 p = 0.5
2
2
1
p = 0
p = 0.15 4
6
8
(a)
10
0
0
4
2
p
= 0.15
6 (b)
8
10
3
12 2
p = 0.5 p =0.25
p = 0.21
1
0
1
0
p =0
p = 0.15
2
1
(c)
3
Fig. 19.5 Description of simple shear deformation of kinematic hardening material by the corotational rate with a plastic spin (Dafalias 1985b).
nðrÞ: r ¼ nðrÞ: r
ð19:65Þ
by Eq. (6.63) since only one variable r is involved in Eq. (19.65) exhibiting the isotropy, while the elastic response is influenced by the types of stress rate. In the plastically-anisotropic material, on the other hand, the response depends on the difference of the time-derivative as follows: (1) The plastic strain rate is influenced by the difference of stress rate because of ^ðr; aÞ: r by Eq. (6.64), which is included in the constitutive ^ðr; aÞ: r 6¼ n n relation of plastic strain rate in Eq. (8.81). (2) The evolution of the anisotropic hardening variable, e.g. the back stress is influenced by the difference of the time-derivative, so that the yield surface on
19.3
Plastic Spin
581
which the stress lies is also influenced by it. Therefore, the stress vs. strain relation is influenced by the difference of the time-derivative. It should be recognized that the accurate description of elastoplastic deformation can be attained up to around one hundred percent shear strain by the Jaumann rate without the plastic spin even for plastically-anisotropic materials, since identical deformation behavior is described by using any corotational rate up to that percent as known from Figs. 19.2 and 19.5.
Chapter 20
Localization of Deformation
Even if material is subjected to a homogeneous stress, the deformation concentrates in a quite narrow strip zone as the deformation becomes large and finally the material results in failure. Such a concentration of deformation is called the localization of deformation and the strip zone is called the shear band. The shear band thickness is the order of several microns in metals and ten and several times of particle diameter in soils. Therefore, a large shear deformation inside the shear band is hardly reflected in the change of external appearance of the whole body, although the stress is estimated by the external load and the outer appearance of material. Therefore, a special care is required for the interpretation of element test data and the analysis taken account of the inception of shear band is indispensable when a large deformation is induced. The localization phenomenon of deformation and its pertinent analysis are described in this chapter.
20.1
Element Test
The purpose of the element test of material is to find the constitutive property of material and thus the element test is premised that a homogeneous deformation proceeds reflecting the constitutive property. However, the deformation becomes heterogeneous when a large deformation accompanying with the shear band is induced. Then, the strain (rate) inside the shear band is far larger than the one based on the variation in the outer appearance of the whole material element. Here, the stress (rate) applied to the material element is relevant to the strain (rate) inside the shear band. In other words, the strain (rate) relevant to the constitutive property is not a strain (rate) based on the variation of the outer appearance of material. This fact is illustratively depicted in Fig. 20.1 for the softening material. The strain (rate) observed in the element test is designated by de which is the average strain rate in the whole of the element, while the strain rate inside the shear band is designated by © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_20
583
584
20
Localization of Deformation
Fig. 20.1 Stress–strain curve in the constitutive property and the element test with a localization in a softening state
de. Here, note that the actual stress increment dr is not relevant to de but relevant to de, where dr and de are related by constitutive property. This fact must be considered when we determine the material parameters from element test data.
20.2
Gradient Theory
The stress is determined locally by the deformation at each material point when the deformation is small moderately. However, it is influenced also by the deformations in surrounding material particles, i.e. by the gradient of deformation when the deformation develops to be inhomogeneous as observed inside the shear band, in the necking part, etc. The former can be described by the local theory taking account of only the deformation in each material point but the latter can be described by incorporating the nonlocal theory taking account of the deformations not only in each material point but also in the surrounding material particles. Here, there exists the limitation in the gradient of deformation so that the shear band thickness does not decrease less than a certain limitation. The limitation is regulated by the material constant, called the characteristic length. In the finite element method ignoring this fact, the deformation concentrates in the narrow band zone corresponding to the size of one element. Then, if the finite elements are downsized with finer meshes aiming at obtaining an accurate solution, the shear band thickness is reduced infinitely, resulting in the mesh-size dependence losing the reliability of solution, i.e. the ill-posedness. The non-local theory, called the gradient theory, for the elastoplastic deformation, in which the gradients of strain rate, stress rate and internal variables are incorporated, was proposed by Aifantis (1984). It will be described briefly in this section.
20.2
Gradient Theory
585
Adopting the yield condition with the isotropic hardening for the sake of simplicity and introducing the spatial gradient of the isotropic hardening variable, let the yield condition (8.16) be extended as follows: f ðrÞ ¼ hhFðHÞii
ð20:1Þ
where hh ii designates the second-order gradient, i.e hh ii 1 þ c22 r2 ; r2
@2 @2 @2 þ þ @x21 @x22 @x23
ð20:2Þ
c2 is the material constant reflecting the effect of the gradient of the mechanical state. Here, for the sake of simplicity, the higher order gradient is not incorporated. The first-order gradient is not incorporated because the odd-order gradients in opposite directions are canceled each other. The material-time derivative of Eq. (20.1) leads to @f ðrÞ 0 : r ¼ F ð1 þ c22 r2 Þ½H @r
ð20:3Þ
Substituting Eq. (8.23) with Eq. (8.24) into Eq. (20.3), one has @f ðrÞ 0 : r ¼ F fHn ðr; H; nÞð1 þ c22 r2 Þ½k @r
ð20:4Þ
It follows from Eq. (20.4), noting ð1 þ c22 r2 Þð1 c22 r2 Þ ¼ 1 c42 r4 ffi 1, that
k¼
n :ð1 c22 r2 Þ½r ; Mp
d ¼ E1 : r þ
dp ¼
n :ð1 c22 r2 Þ½r n Mp
n :ð1 c22 r2 Þ½r n Mp
ð20:6Þ
n : ð1 c22 r2 Þ½r Mp n :ð1 c22 r2 Þ½r p 2 2 ffi ½M ð1 þ c2 r Þ þ n : E : n Mp Mp n :ð1 c22 r2 Þ½r r2 ¼ ðM p þ n : E : nÞ 1 þ c22 p M þn : E : n Mp 2 ~ n :ð1 c2 rÞ½r ¼ ðM p þ n : E : nÞð1 þ rÞ Mp
ð20:5Þ
n : E : d ¼ n :r þn : E : n
ð20:7Þ
586
20
Localization of Deformation
where ~ c2 #r2 ; # r 2
Mp
Mp þn : E : n
ð20:8Þ
The plastic modulus M p is given by Eq. (8.27) as it is. On the other hand, it is obtained from Eq. (20.7) that
~ K ¼ ð1 rÞ
~ n:E:d n : E :ð1 rÞd ffi Mp þ n : E : n Mp þ n : E : n
r¼E:dE:n
~ n : E :ð1 rÞd p M þn : E : n
ð20:9Þ
ð20:10Þ
In what follows, the above-mentioned equations are extended for the subloading surface model with the isotropic and the anisotropic hardening. Incorporating the gradient into the internal variables, the subloading surface with the kinematic and the rotational hardening in addition to the isotropic hardening is given as (Hashiguchi and Tsutsumi 2006): f ðr hha ii; hhb iiÞ ¼ hhRFðHÞ ii
ð20:11Þ
Considering Eqs. (20.2) and (20.11) leads to f r ð1 þ c22 r2 Þ½a; ð1 þ c22 r2 Þ½b ¼ ð1 þ c22 r2 Þ½RFðHÞ
ð20:12Þ
The material-time derivative of Eq. (20.12) leads to @f ðr ð1 þ c22 r2 Þ½a; ð1 þ c22 r2 Þ½bÞ :r @r @f ðrð1 þ c22 r2 Þ½a; ð1 þ c22 r2 Þ½bÞ :ð1 þ c22 r2 Þ½a @r @f ðr ð1 þ c22 r2 Þ½a; ð1 þ c22 r2 Þ½bÞ :ð1 þ c22 r2 Þ½b þ 2 2 @ð1 þ c2 r Þ½b
ð20:13Þ
¼ ð1 þ c22 r2 Þ½R F þ R F The gradients of internal variables can be ignored since they are small compared to the gradient of their rates and thus Eq. (20.13) is reduced approximately to @f ð^ r; bÞ @f ð^ r; bÞ @f ð^ r; bÞ :r :ð1 þ c22 r2 Þ½a þ :ð1 þ c22 r2 Þ ½b @r @r @b
¼
Fð1 þ c22 r2 Þ½R þ
RF
0
ð1 þ c22 r2 Þ½H
ð20:14Þ
20.2
Gradient Theory
587
Substituting Eq. (9.9) for the evolution rule of normal-yield ratio, the consistency condition is derived from Eq. (20.14) as follows: @f ð^ r; bÞ @f ð^ r; bÞ @f ð^ r; bÞ :r : 1 þ c22 r2 ½a þ : 1 þ c22 r2 ½b @r @r @b p 2 2 ¼ UF 1 þ c2 r ½kd k þ RF 0 1 þ c22 r2 ½H
ð20:15Þ
Further, substituting the associated flow rule (9.21) into Eq. (20.15), it is obtained that i h @f ð^ r; bÞ @f ð^ r; bÞ f kn ðr; a; F; b nÞ :r : 1 þ c22 r2 k @r @r i h ð^ r; bÞ 2 2 þ @f f bn ðMr ; b; F; n0 Þ : 1 þ c2 r @b k ¼ UF 1 þ c22 r2 k þ RF 0 1 þ c22 r2 k fHn ðr; H; nÞ
ð20:16Þ
which can be approximately given by @f ð^ r; bÞ @f ð^ r; bÞ n Þ 1 þ c22 r2 ½k :r : f kn ðr; a; F; b @r @r @f ð^ r; bÞ : f bn ðMr ; b; F; n0 Þ 1 þ c22 r2 ½k þ @b ¼ ðUF þ RF 0 fHn ðr; H; nÞÞ 1 þ c22 r2 ½k
ð20:17Þ
The plastic multiplier is derived from Eq. (20.17) as " # n :ð1 c22 r2 Þ½r 2 2 n :r k ¼ ð1 c2 r Þ ffi p p M M
ð20:18Þ
0 F U fHn ðr; H; nÞ þ F R 1 @f ð^ r; bÞ 0 :f bn ðMr ; r; F; n Þ r þ f kn ðr; a; F; b nÞ RF @b
ð20:19Þ
where p
M n:
Consequently, the plastic strain rate is given as
dp ¼
n :ð1 c22 r2 Þ½r n p M
ð20:20Þ
The shear band thickness of softening soil was predicted adopting Eq. (20.20) by Hashiguchi and Tsutsumi (2006).
588
20
Localization of Deformation
Here, it is noteworthy that we must use quite small elements with the size of several tens of shear band thickness to take the effect of the gradient into account correctly in the finite element analysis. Therefore, it is nearly impossible to apply the gradient theory to the finite element analysis of boundary value problems in engineering practice at least at present. The gradient theory is used widely for prediction of shear band thickness, size effects, etc. using fine meshes for very small specimens.
20.3
Shear-Band Embedded Model: Smeared Crack Model
Although the gradient theory is not applicable to the analysis of practical engineering problems at present, the practical model for the finite element analysis for softening materials has been proposed as described below. As the deformation becomes large and the shear band is formed, the plastic deformation concentrates inside the shear band and thus the softening is accelerated leading to the rapid reduction of stress. As the result, inversely, the unloading leading to the elastic state occurs outside the shear band. Consequently, the elastoplastic constitutive equation holds only inside the shear band. Then, denoting the strain rate and the plastic strain rate calculated from the external appearance by p d and d , respectively, called the apparent strain rate and the apparent plastic strain rate, respectively, the elastoplastic constitutive equation in terms of the apparent strain rate is proposed. It is called the shear-band embedded model or smeared crack model (Pietrueszczak and Mroz 1981; Bazant and Cedolin 1991). Denoting the ratio of the area of a shear band to the area of a two-dimensional finite element by Sð 1Þ, the following relations are postulated. p
d ¼ Sdp
ð20:21Þ
n :r d ¼ d þ d ¼ d þ Sd ¼ E : r þ S p n M e
p
e
p
1
ð20:22Þ
Tanaka and Kawamoto (1988) proposed the simple equation of S for the plane strain condition as follows: pffiffiffiffiffi S ¼ ðw lÞ=ðl lÞ ¼ w= Fe
ð20:23Þ
supposing simply the square finite element with the side-length l and the shear band having the thickness w, where Fe ð¼ l lÞ is the area of the finite element. The plastic multiplier is expressed in terms of the apparent strain rate from Eq. (20.22) as follows:
20.3
Shear-Band Embedded Model: Smeared Crack Model
! n :r n:E:d K ¼ S p ¼ Mp M S þn : E : n
589
ð20:24Þ
Then, the stress rate is given by n:E:d E:n þn : E : n
ð20:25Þ
Mp n:E:d M p þ Sn : E : n
ð20:26Þ
r ¼ E : d Mp S
Then, we have
n :r ¼
from which it is known that the stress reduction is larger for smaller value of S, i.e. smaller thickness of shear band, noting M p \0, n : E : d [ 0 and S\1. It is desirable to choose material parameters such that Eq. (20.22) or (20.25) fits to a measured stress–strain curve, using the value of S predicted by a pertinent method, if we determine them from element test data.
20.4
Necessary Condition for Shear Band Inception
Discontinuity of the velocity gradient is induced at the shear band boundary. Here, incorporate the coordinate system in which the coordinate axes x 1 and x 2 are taken to be normal and parallel, respectively, to the shear band as shown in Fig. 20.2. The discontinuity of velocity gradient can be induced only in the x 1 -direction. Therefore, only the following quantities are not zero, designating the discontinuous quantity by Dð Þ.
g11
@v1 D ; @x 1
g21
@v2 D @x 1
ð20:27Þ
Then, the discontinuity of strain rate is given by
1 @vi @vj 1 @vi @xr @vj @xr Ddij ¼ þD D D þD ¼ @xj @xi @x r @xj @x r @xi 2 2
1 @vi @x1 @vj @x1 1 þD D ¼ ¼ fgi1 ðn ej Þ þ gj1 ðn ei Þg
@x1 @xj @x1 @xi 2 2 1 i ¼ ðg1 nj þ gj1 ni Þ 2 ð20:28Þ
590
20
Localization of Deformation
Fig. 20.2 Discontinuity of velocity gradient induced in the direction normal to the shear band
where n is the unit vector in the direction normal to the shear band, i.e. the x 1 -direction. On the other hand, the discontinuity in the rate of traction vector tn applying to the discontinuity surface of velocity gradient, i.e. shear band having the direction vector n, is described by 9 ep ep 1 = Ddkl nj ¼ Cj1kl ðgk1 nl þ gl1 nk Þnj > Dtn1 ¼ Drj1 nj ¼ Cj1kl 2 > ep ep 1 ðgk nl þ gl1 nk Þnj ; Dtn2 ¼ Drj2 nj ¼ Cj2kl Ddkl nj ¼ Cj2kl 2 1 Noting 9 ep 1 ep k = Cj1kl ðgk1 nl þ gl1 nk Þnj ¼ Ci1kj g1 ni nj > 2 > ep 1 ep k Cj2kl g1 ni nj ; ðgk nl þ gl1 nk Þnj ¼ Ci2kj 2 1 Equation (20.29) is expressed as (
Dtn1
Dtn2
)
" ¼
ep Ci11j ni nj
ep Ci12j ni nj
ep Ci21j ni nj
ep Ci22j ni nj
#(
g11
g21
)
ð20:29Þ
20.4
Necessary Condition for Shear Band Inception
591
That is to say, (
Dtn1
Dtn2
)
" ¼
A11 A12
#(
g11
)
ðDtni ¼ Aij gj1 ;
g21
A21 A22
Dtn ¼ Ag1 Þ
ð20:30Þ
where Aij is given by the following equation and called the acoustic tensor. ep Aij Crijs nr ns
A nCep n;
ð20:31Þ
Here, noting that the traction rate vector must be continuous, i.e. Dtn ¼ 0 by the equilibrium and thus it must hold from Eq. (20.30) that "
A11 A12 A21 A22
#(
g11
g21
) ¼
( ) 0 0
ðAij gj1 ¼ 0;
Ag1 ¼ 0Þ
ð20:32Þ
In order that Eq. (20.32) has a solution other than the trivial solution g1 ¼ 0, i.e. that the discontinuity of velocity gradient is induced, the following equation must hold, noting g1 ¼ A1 0 with Eq. (1.169).
A A
11 12 det A ¼
¼0
A21 A22
ð20:33Þ
At least one of eigenvalue of the acoustic tensor A is zero when (20.33) holds. The search for the occurrence of n fulfilling Eq. (20.33), i.e. the inception of the shear band and its direction, can be performed by the eigenvalue analysis. Equation (20.33) is given explicitly as (Hashiguchi and Protasov 2004) ep ep ep ep n1 n1 þ C1ij2 n1 n2 þ C2ij1 n2 n1 þ C2ij2 n2 n2 Þ detðnCep nÞ ¼ detðC1ij1
ep
C n n þ C ep n n þ C ep n n þ C ep n n 1 1
1112 1 2 2111 2 1 2112 2 2 ¼ 1111 ep ep ep ep
C1211 n1 n1 þ C1212 n1 n2 þ C2211 n2 n1 þ C2212 n2 n2
ep
C n2 þ ðC ep þ C ep Þn1 n2 þ C ep n2
1111 1 1112 2111 2112 2 ¼ ep 2 ep ep ep
C1211 n1 þ ðC1212 þ C2211 Þn1 n2 þ C2212 n22
ep ep ep ep C1121 n1 n1 þ C1122 n1 n2 þ C2121 n2 n1 þ C2122 n2 n2
ep ep ep ep C1221 n1 n1 þ C1222 n1 n2 þ C2221 n2 n1 þ C2222 n2 n2
ep ep ep ep C1121 n21 þ ðC1122 þ C2121 Þn1 n2 þ C2122 n22
ep ep ep ep C1221 n21 þ ðC1222 þ C2221 Þn1 n2 þ C2222 n22
ep ep ep ep ep ep ep ep ¼ fC1111 n21 þ ðC1112 þ C2111 Þn1 n2 þ C2112 n22 gfC1221 n21 þ ðC1222 þ C2221 Þn1 n2 þ C2222 n22 g ep ep ep ep ep ep ep ep fC1121 n21 þ ðC1122 þ C2121 Þn1 n2 þ C2122 n22 gfC1211 n21 þ ðC1212 þ C2211 Þn1 n2 þ C2212 n22 g
592
20
Localization of Deformation
ep ep ep ep ¼ ðC1111 C1221 C1121 C1211 Þn41 ep ep ep ep ep ep ep ep þ ðC1111 C1222 þ C1111 C2221 þ C1112 C1221 þ C2111 C1221 ep ep ep ep ep ep ep ep C1121 C1212 C1121 C2211 C1122 C1211 C2121 C1211 Þn31 n2 ep ep ep ep ep ep ep ep þ ðC1111 C2222 þ C1221 C2112 C1121 C1211 C1211 C2122 Þn21 n22 ep ep ep ep ep ep ep ep þ ðC1122 C2222 þ C2111 C2222 þ C1222 C2112 þ C2221 C2112 ep ep ep ep ep ep ep ep C1122 C2212 C2121 C2212 C2122 C1212 C2122 C2211 Þn1 n32 ep ep ep ep þ ðC2112 C2222 C2122 C2212 Þn42 ep ep ep2 ¼ ðC1111 C1212 C1112 Þn41 ep ep ep ep ep ep ep ep þ ðC1111 C1222 þ C1111 C2212 C1112 C1122 C1122 C1211 Þn31 n2 ep ep ep ep ep ep ep ep þ ðC1111 C2222 þ C1212 C1212 C1112 C1211 C1211 C1222 Þn21 n22 ep ep ep ep ep ep ep ep þ ðC1122 C2222 þ C12111 C2222 C1122 C1222 C1222 C1122 Þn1 n32 ep ep ep 4 þ ðC1212 C2222 C1222 2 Þn2
which is reduced to det A ¼ a1 n41 þ a2 n31 n2 þ a3 n21 n22 þ a4 n1 n32 þ a5 n42 ¼ 0
ð20:34Þ
9 > > > > > ep ep ep ep ep ep ep ep > C1111 C1222 þ C1111 C2221 C1121 C2211 C1122 C1211 ; > > = ep ep ep ep ep ep ep ep C1111 C2222 þ C1221 C2112 C1121 C1211 C1211 C2122 ; > > > ep ep ep ep ep ep ep ep > C1122 C2222 þ C2111 C2222 C1122 C2212 C2122 C2211 ;> > > > ; ep ep ep2 C2112 C2222 C2122 ;
ð20:35Þ
where ep ep ep2 a1 C1111 C1221 C1121 ;
a2 a3 a4 a5 Setting
n1 ¼ cos h;
n2 ¼ sin h
ð20:36Þ
Equation (20.34) is rewritten as gðhÞ ¼ a5 tan4 h þ a4 tan3 h þ a3 tan2 h þ a2 tan h þ a0 ¼ 0
ð20:37Þ
which, noting the symmetry gðhÞ ¼ gðhÞ, leads to gðhÞ ¼ a5 tan4 h þ a3 tan2 h þ a0 ¼ 0
ð20:38Þ
20.4
Necessary Condition for Shear Band Inception
593
There exists the possibility that a shear band occurs in the direction h fulfilling Eq. (20.38). Here, note that Eq. (20.33) is the only necessary condition for the inception of the shear band. We searched above the discontinuity of the velocity gradient in the direction normal to the shear band, while the traction rate vector must be continuous in that direction. Inversely, on the other hand, the search for the discontinuity in the normal stress rate component applied to the surface normal to the shear band, i.e. Dr 22 6¼ 0, while the discontinuity in the normal strain rate component in the direction parallel
to the shear band must be zero, i.e. De 22 ¼ 0, is called the compliance method (cf. Mandel 1964). Here, note that the normal strain rate component in the direction
normal to the shear band can be discontinuous, i.e. De 11 6¼ 0 in dilative materials. It was applied to the prediction of the direction of shear band formation in soils (Vermeer 1982).
Chapter 21
Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
The theory aiming at the formulation of the deformation phenomenon in metals by the analysis on the microscopic level of crystal slips is called the crystal plasticity. However, it is still now limited to the formulation within the framework of the hypoelastic-based plasticity as represented in the models proposed by Peirce et al. (1982, 1983), Asaro and Lubarda (2006), etc. which is still widely used today. However, the elastoplastic deformation cannot be formulated exactly by the hypoelastic-based model in which the elastic deformation is limited to be infinitesimal and the complicated task for the time-integration of corotational stress rate is required. The basic framework of the elastoplasticity which can describe the elastoplastic deformation exactly is the multiplicative decomposition of the deformation gradient tensor described in Chap. 17, while needless to say it does not possess the defects involved in the hypoelastic-based plasticity. In particular, it should be noted that the isoclinic concept which is inevitable for the exact formulation of the finite elastoplastic deformation within the framework of the multiplicative decomposition was found from the crystal plasticity analysis by Mandel (1971). Unfortunately, however, the existing crystal plasticity model goes back to the hypoelastic-based formulation (cf. Hutchinson 1976; Peirce et al. 1982, 1983; Huang 1991; Asaro and Lubarda 2006; de Souza Neto et al. 2008; Belytschko et al. 2014; Feather et al. 2021), disusing the multiplicative decomposition for the explicit formulation of the stress rate vs. strain rate relation and incorporating the irrational viscoplastic model, i.e. creep model as described in Chap. 14. The exact crystal plasticity model based on the multiplicative decomposition in addition to the hypoelastic-based crystal plasticity will be formulated in this chapter, while it will be extended to describe the cyclic loading behavior by incorporating the concept of the subloading surface.
© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_21
595
596
21.1
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
Description of Strain Rate and Spin by Crystal Lattice Vectors
The intermediate configuration is obtained by excluding the rigid-body rotation in addition to the elastic deformation from the current configuration as was explained in Chap. 17. Therefore, the intermediate configuration is independent of the rigid-body rotation. Consequently, the deformation gradient F is multiplicatively decomposed into the elastic deformation gradient Fe composed of the substructure rotation tensor Re and the right elastic stretch tensor Ue and the plastic deformation gradient Fp as follows (see Fig. 21.1): 8 ¼ FdX x ¼ Fe d X dx ¼ Re de > > < e1 e de x ¼ R dx ¼ U d X ¼ Ue Fp dX ¼ Fe1 dx ¼ Ue1 dx ¼ Fp dX > dX > : x ¼ Fp1 d X dX ¼ F1 dx ¼ Fp1 Ue1 de
ð21:1Þ
are the position vectors of material particles in the initial where X, x and X (reference), the current and the intermediate configuration, respectively. Further, e x is the position vector of material particle in the configuration pulled-back by the substructure rotation Re from the current configuration. Substructure rotation Re
s e w p (w = w )
w nα
sα
nα
ws = w − w p
sα
w =wp (ws = w − w p = O)
Fe F Total deformation/rotation
w =wp
nα0
sα0
(ws = w − w p = O)
nα0
sα0
0
Fig. 21.1 Multiplicative decomposition in crystal plasticity based on isoclinic concept (Mandel 1973, 1974)
21.1
Description of Strain Rate and Spin by Crystal Lattice Vectors
597
Then, it follows that F ¼ Fe Fp ¼ Re Ue Fp
ð21:2Þ
Fe ¼ Re Ue ¼ Ve Re
ð21:3Þ
Fp ¼ Rp Up ¼ Vp Rp
ð21:4Þ
The initial and the current configurations are designated by K0 and K, respectively, and the configurations pulled-back from the current configuration by the substructure rotation Re and by the elastic deformation gradient Fe are designated as e and K, respectively, as shown in Fig. 21.1. K The basic equations for the velocity gradient tensors shown in Sect. 17.2.2 will be described again below for the reader’s convenience. For the current configuration: l ¼ le þ lp l
ð21:5Þ 9 > > > > =
@v 1 ¼ FF @x
p Fe1 > le Fe Fe1 ; lp Fe L > > > ; p p p1 L F F 9 l ¼ dþw = le ¼ de þ we ; lp ¼ d p þ w p d ¼ de þ dp w ¼ we þ wp
ð21:6Þ
ð21:7Þ
h i h i 9 > d ¼ sym½l ¼ sym F F1 ; w ¼ ant½l ¼ ant F F1 > = h i h i e e e e e1 e e1 e ; w ¼ ant½l ¼ ant F F d ¼ sym½l ¼ sym F F > > ; p p dp ¼ sym½lp ¼ sym Fe L Fe1 ; wp ¼ ant½lp ¼ ant Fe L Fe1
ð21:8Þ
ð21:9Þ
For the intermediate configuration: e
p
L ¼ L þL
ð21:10Þ
9 1 L Fe lFe = e L Fe1 le Fe ¼ Fe1 Fe 1 ; p L Fe1 lp Fe ¼ Fp Fp
ð21:11Þ
598
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
L ¼ DþW e e e p p p L ¼ D þW ;L ¼ D þW e
p
D ¼ D þD ;
e
W ¼ W þW
ð21:12Þ p
ð21:13Þ
e p p e D sym L ; D sym L e p e p W ant L ; W ant L
D sym½L; W ant½L;
ð21:14Þ
Now, adopt the embedded base ðg1 ; g2 ; g3 Þ described in Sect. 3.2. Here, the initial primary base vector g1 is chosen parallel to the crystal lattice and denoted by sa and the secondary reciprocal base vector g2 by na in the slip system a, satisfying sa na ¼ 0 by Eq. (3.2)2 , while they are called the lattice vectors, director frame, director triad, isoclinic triad, etc. Limiting to the two-dimensional deformation, g3 ð¼ g3 Þ is the unit vector ðkg3 k ¼ 1Þ and chosen perpendicular to the base vectors sa and na . Further, the base vectors sa and na in the initial configuration arechosen to be the unit vectors and denoted by the symbols sa0 and na0 sa0 ¼ na0 ¼ 1 , respectively, which correspond to the reference base vectors G1 and G2 , respectively, in Eq. (3.2)1. Reminding the aforementioned assumption that the intermediate configuation is independent of the rigid-body rotation and noting the simple shear deformation along the crystalline lattice under the plastic incomressibility, the base vectors are kept unchaged as sa0 and na0 in the process from the initial to the intermediate configurations (see Fig. 21.1). This physical consequence is referred to as the isoclinic concept by Mandel (1973, 1974), while “isoclinic” is the Greek word meaning “same (or constant) direction”. The current primary base vector sa and its reciprocal base vector na are related to the initial base vectors sa0 and na0 by Eqs. (3.8) and (3.11), replacing F to Fe , as follows: 1
sa ¼ Fe sa0 ¼ sa0 FeT ; sa0 ¼ Fe sa ¼ sa FeT na ¼ FeT na0 ¼ na0 Fe1 ; na0 ¼ FeT na ¼ na Fe
9 > =
> T sa na ¼ Fe sa0 FeT na0 ¼ sa0 Fe FeT na0 ¼ sa0 na0 ¼ 0 ;
ð21:15Þ
noting Tv ¼ vTT Tij vj ¼ vj Tij . The rates of sa and na are given as follows:
sa ¼ Fe sa0 ¼ Fe Fe1 sa ¼ le sa ¼ sa leT
9 =
na ¼ FeT na0 ¼ FeT FeT na ¼ le na ¼ na le ;
T
ð21:16Þ
Needless to say, the current base vectors ðsa ; na Þ are no longer unit vectors.
21.1
Description of Strain Rate and Spin by Crystal Lattice Vectors
599
On the other hand, the simple shear strain ca along the slip system a is additively decomposed into the elastic shear strain cea and the plastic shear strain cpa as follows:
ca ¼ cea þ cpa ; c a þ c ea þ c pa
ð21:17Þ
The difference of plastic displacements dupa in both ends of infinitesimal line element dX, which is induced by the plastic shear strain cpa in the slip system a, is given by the following equation (see Fig. 21.2). dupa ¼ cpa na0 dX sa0 ¼ cpa sa0 na0 dX
ð21:18Þ
dX ¼ dX þ dupa ¼ I þ cpa sa0 na0 dX
ð21:19Þ
Then, dX changes to
in the intermediate configuration. Noting Eqs. (17.3)2 and (21.19), one has Fpa ¼ I þ cpa sa0 na0 ; by virtue of noting
Fpa1 ¼ I cpa sa0 na0
ð21:20Þ
I þ cpa sa0 na0 I cpa san na0 ¼ I due to na0 sa0 ¼ 0. Further,
pα d u pα = γ (nα0 • d X)sα0
γ pα nα0 • dX γ pα
nα0 • dX Fp 1 Plastic deformation
dX
dX
dX = FpdX
n0
n0
s0
s0
0
X
X 0
Fig. 21.2 Plastic displacement at end of infinitesimal line element dX, which is induced by plastic shear strain in slip system
600
I þ cpa sa0 na0
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
I cpa sa0 na0 ¼ c pa sa0 na0 c pa sa0 na0 cpa sa0 na0 ð21:21Þ
it follows that
F pa Fpa1 ¼ c pa sa0 na0
ð21:22Þ
The plastic velocity gradient based in the intermediate configuration is given by summing the plastic velocity gradients induced in all the relevant slip systems as follows: p
L ¼
n X
F pb Fpb1 ¼
n X
b¼1
sb0 nb0 c pb
ð21:23Þ
b¼1
where n is the number of slip systems. The substitution of Eq. (21.23) into Eq. (21.12) leads to p
p
p
L ¼ D þW ¼
n X
b
P þQ
b
c pb
ð21:24Þ
b¼1 n p P p b D ¼ sym L ¼ P c pb p
b¼1
p
W ¼ ant L
n P
¼
b
ð21:25Þ
Q c pb
b¼1
where ) a P ¼ sym sa0 na0 a Q ¼ ant sa0 na0
ð21:26Þ
Substituting Eq. (21.23) with Eq. (21.15) into Eq. (21.6) reads: lp ¼ dp þ wp ¼
n X
sb nb c pb
ð21:27Þ
b¼1
Inserting Eq. (21.27) into Eq. (21.9), one has dp ¼ sym½lp ¼
n P b¼1
w ¼ ant½l ¼ p
p
n P
a¼1
9 > pb c pb > = b
pb
q c
> > ;
ð21:28Þ
21.1
Description of Strain Rate and Spin by Crystal Lattice Vectors
601
with pa ¼ sym½sa na qa ¼ ant½sa na
ð21:29Þ
Substituting Eq. (21.27) into Eq. (21.5), one has le ¼ l l p ¼ l
n X
sb nb c
pb
¼l
b¼1
n X pb þ qb c pb
ð21:30Þ
b¼1
9 h i n P pb c pb > de ¼ sym½le ¼ sym F e Fe1 ¼ d > = b¼1 h i n P b pb > > we ¼ ant½le ¼ ant F e Fe1 ¼ w q c ;
ð21:31Þ
b¼1
21.2
Resolved Shear Stress (Rate)
The resolved shear stress (rate) is formulated for the hypoelastic-based plasticity and the multiplicative hyperelastic-based plasticity in this section. (a) Hypoelastic-based plasticity Let the resolved shear stress be described reffering to Asaro and Rice (1977) in the following. sa ¼ sa sna ¼ pa : s
ð21:32Þ
The time-differentiation of Eq. (21.32) leads to s a ¼ s a sna þ sa s na þ sa s n a
Substituting Eqs. (21.7) and (21.16), the right-hand side of this equation leads: s a sna þ sa s na þ sa s na ¼ ðde þ we Þsa sna þ sa s na sa sðde þ we ÞT na ¼ sa ðde þ we ÞT sna þ sa s na sa sðde we Þna ¼ sa s we s þ swe þ de s sde na
leading to
s a ¼ sa s na þ sa ðde s sde Þna
ð21:33Þ
602
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
where ð Þ designates the corotational rate observed from the substructure of material itself so that it is given by excluding the substructure spin we from the
material-time derivative as shown in Eg. (3.43). The corotational rate s is related to the elastic strain rate by the hypoelastic relation, i.e.
where
s ¼ E : de
ð21:34Þ
T s ¼ Re ReT sRe Re ¼ s we s þ swe
ð21:35Þ
and E is the overall elastic modulus tensor. Accounting for Eqs. (21.29) and (21.35) and the symmetry of s and de , one has
sa s na ¼sai s ij naj ¼ sai naj s ij ¼ sa na : s ¼ pa : s ¼pa : E :de ¼ paij Eijkl dkle ¼ Eklij paij dkle ¼ Epa :de sa ðde s sde Þna ¼ sai dire srs nas sai sir drse nas 1 a a e a a e 1 a a e ¼ s n ssr dir þ srs nas sai drie ns si sir dsr þ sri si ns drs 2 i s 2 1 1 ¼ f½ðsa na Þs:de ½ðna sa Þs:de g f½sðsa na Þ:de ½sðna sa Þ:de g 2
2 1 1 ¼ ðsa na na sa Þs s ðsa na na sa Þ :de 2 2 ¼ ðqa s sqa Þ: de ¼ ba : de
where ba qa s sqa
ð21:36Þ
Substituting these relations into Eq. (21.33), we have the relation between the resolved shear stress rate versus the global elastic strain rate as follows:
where
s a ¼ Na : de
ð21:37Þ
6 NaT N a E : pa þ b a ¼
ð21:38Þ
Further, substituting Eq. (21.31)1 into Eq. (21.37) reads: ! n P a a b pb s ¼N : d p c b¼1
ð21:39Þ
21.2
Resolved Shear Stress (Rate)
603
(b) Multiplicative hyperelastic-based plasticity The resolved shear stress in Eq. (21.32) is represented by the Mandel stress in the intermediate configuration as follows: sa ¼ sa0 Mna0 ¼ Ga0 : M
ð21:40Þ
Ga0 sa0 na0
ð21:41Þ
where
noting sa sna ¼ Fe Sa0 sFeT na0 ¼ sa0 FeT sFeT na0 ¼ sa0 Mna0
ð21:42Þ
by virtue of Eqs. (5.22) and (21.15). The time-differentiation of Eq. (21.40) leads to
na ¼ Ga : M s a ¼ sa0 M 0 0
ð21:43Þ
Here, assume the incompressible Neo-Hookean-type strain eneryg function in Eq. (7.38), i.e. 1 1 ð21:44Þ we C e ¼ lðIc e 3Þ ¼ l trC e 3 2 2 e
where Ice trC , leading to S¼
@we C e @C e
¼ lI
ð21:45Þ
and M ¼ lC e
ð21:46Þ
M ¼ l C e ¼ 2lsym C e L L p
ð21:47Þ
Then, noting Eq. (17.23), one has
Hence, it follows from Eq. (21.43) with Eqs. (21.47) and (21.25) that a
s ¼
sa0
(
lC
e
na0
¼
2lsa0
" #) n e P e b b pb na0 sym C L sym C P þQ c b¼1
ð21:48Þ
604
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
21.3
Stress Rate Versus Plastic Shear Strain Rate Relation
The stress rate versus plastic shear strain rate relation is formulated for the hypoelastic-based plasticity and the multiplicative hyperelastic-based plasticity in this section. (a) Hypoelastic-based plasticity The Jaumann rate of the Kirchhoff stress is given as w
s ¼ s ws þ sw
ð21:49Þ
The corotational rate s in Eq. (21.34) is related to s w as
w
s s we s þ swe ¼ s þ wp s swp
ð21:50Þ
noting
s we s þ swe ¼ s ðw wp Þs þ sðw wp Þ
ð21:51Þ
The substitution of Eq. (21.28) with Eq. (21.36) into Eq. (21.50) leads to
w
s¼s þ
n X
bb c pb
ð21:52Þ
b¼1
Further, substituting Eqs. (21.8)1 and (21.28)1 into Eq. (21.34), one has
s ¼ E : ðd dp Þ ¼ E :d
n X
E : pb c pb
ð21:53Þ
b¼1
The substitution of Eq. (21.53) into Eq. (21.52) reads:
sw ¼ E : d
n P
Nb c pb
ð21:54Þ
b¼1
(b) Mulitiplicative hyperelastic-based plasticity It follows from Eqs. (21.24) and (21.47) that ( " #) n X e e b b pb M ¼ 2l sym C L 2sym C P þQ c b¼1
However, this equation will not be used in the later formulation.
ð21:55Þ
21.4
Conventional Crystal Plasticity Model
21.4
605
Conventional Crystal Plasticity Model
The crystal plasticity model based on the conventional plasticity assuming the yield region enclosing the purely-elastic region is shown in this section.
21.4.1
Yield Condition and Flow Rule
The crystal shear yield condition describing the crystal shear yield region in the slip system a is given by j^sa j ¼ say
ð21:56Þ
^sa sa sak
ð21:57Þ
where
say ð[ 0Þ is the shear hardening function, referred to as the critical shear stress, and sak is the shear kinematic hardening variable in the slip system a. The material-time derivative of the yield condition in Eq. (21.56) reads: ^ na s a s ak ¼ s ay
ð21:58Þ
i.e.
na s ay þ s ak sa ¼ ^
ð21:59Þ
The associated flow rule is adopted for the plastic shear strain rate as follows: c pa ¼ ka ^ na ka 0 c pa ¼ ka
ð21:60Þ
where ^ na
^sa @ j^sa j ¼ @sa j^sa j
ðj^ na j ¼ 1 Þ
ð21:61Þ
606
21.4.2
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
Evolution of Isotropic Hardening
The rate of critical shear stress is specified by
s ay ¼
n P b¼1
n P hpab c pb ¼ hpab kb
ð21:62Þ
b¼1
where hpab is given by the following matrix which is the function of the plastic shear strain (Peirce et al. 1982). p p h ðc Þ for a ¼ b hpab ¼ qhp ðcp Þ þ ð1 qÞhp ðcp Þdab ¼ hpab ¼ ð1 q 1:4Þ qhp ðcp Þ for a 6¼ b ð21:63Þ hp ðcp Þ is given by the function h ðc Þ ¼ p
p
with cp
hp0 sech2
hp0 cp ss sy0
ð21:64Þ
n Z t X pb c dt b¼1
ð21:65Þ
0
based on the single slip law sy ðcp Þ ¼ sy0 þ ðss sy0 Þ tanh
hp0 cp ss sy0
hp0 and sy0 are the initial values of hp and say , i.e. hp0 ¼ hp ð0Þ and sy0 ¼ say ð0Þ, respectively, and ss is the saturation value of say , i.e. ss ¼ say ð1Þ. The function hp ðcp Þ is illustrated in Fig. 21.3.
h p (γ p ) p
h0
0 Fig. 21.3 Function hp ðcp Þ
γp
21.4
Conventional Crystal Plasticity Model
21.4.3
607
Evolution of Kinematic Hardening
The evolution rules of the kinematic hardening will be formulated for the hypoelastic-based plasticity and the multiplicative hyperelastis-based plasticity below. (a) Hypoelastic-based plasticity Adopt the following shear nonlinear-kinematic hardening rule taken account of the latent hardening (Harder 1999). sak ¼ sa sk na ¼ pa : sk sk ¼
n P
aa ðsa na þ na sa Þ ¼ 2
a¼1
a
a ¼ ck
n P a¼1
1 c a c pa aa bk s y
!
pa
¼ ck k
a
ð21:66Þ
aa pa ¼ sTk
1 a ^n a bk say
ð21:67Þ
! ð21:68Þ
a
where ck and bk are the material constants. It follows by substituting Eq. (21.67) into Eq. (21.66) that sak ¼ pa : 2
n X
ab pb ¼ 2
b¼1
where
n X
ab pa : pb ¼
b¼1
n X
ð21:69Þ
yab ab
b¼1
yab sa sb na nb þ sa nb na sb
ð21:70Þ
noting 1 pa : pb ¼ ðsa na þ na sa Þ : sb nb þ nb sb 4 1 a b a s s n nb þ sa nb na sb þ na ¼ 4 1 a b a s s n nb þ sa nb na sb ¼ 2
sb sa
n b þ na
nb s a
sb
ð21:71Þ
The time-derivative of Eq. (21.69) reads
a
sk ¼
n X b¼1
yab ab þ
n X y ab ab b¼1
ð21:72Þ
608
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
Here, noting Eq. (21.16), the time-derivative of yab in Eq. (21.70) is given by •
yαβ = (s•α • s β )(nα • n β ) + (sα • s• β )(nα • n β ) + (sα • nβ )(n• α • s β ) + (sα • s β )(nα • n• β ) •α
+(s
• • n β )(nα • s β ) + (sα • n β )(nα • n β ) + (sα
•
•
⊗ nα )(nα • s β ) + (sα • nβ )(nα • s β )
β = ( sα l e T • n β )(nα • n β ) − (sα • n β l e )(nα • n β ) − (sα • n β )(nα l e • n β ) − (sα • n β )(nα • n l e )
+( sα l e T ⊗ nα )(nα • s β ) − (sα ⊗ nα l e )(nα • s β ) − (sα ⊗ nα )(nα l e • s β ) + (sα ⊗ nα )(nα • sα l e T )
= 2d e : [(sα ⊗ s β )(nα • n β ) − (nα ⊗ n β )(sα • s β )]
=2d e : zαβ
ð21:73Þ zab sa sb na nb na nb sa sb
where
ð21:74Þ
Substituting Eq. (21.73) into Eq. (21.72), one has
s ka ¼
n X
yab a b þ
n X
b¼1
ð21:75Þ
2de : zab ab
b¼1
Then, it follows by substituting Eqs. (21.8), (21.28), (21.60) and (21.68) into Eq. (21.75) that a
s k ¼ ck
n X
!
yab k
b
^ n b
b¼1
i.e. a
sk ¼
n P
"
ck yab ^ nb
b¼1
1
b
bk sby
1 bk sbv
a
a
þ
n X
2 d
n X
c¼1
b
!
n p k ^ b
b
b
: zac ac
ð21:76Þ
b¼1
2^ n p : b b
#
n P
ac c
z a k
b
þ2
c¼1
n P
d : zab ab
b¼1
ð21:77Þ (b) Multiplicative hyperelastic-based plasticity We adopt the following relations in terms of the Mandel-like kinematic hardening _
variable M k in Eq. (17.74) instead of Eq. (21.66). sak
¼
sa0
_
M k na0
¼
_ Ga0 : M k ;
s ak
¼
sa0
_
M k na0
¼
Ga0
_
: Mk
ð21:78Þ
designating the tensor based in the kinematic hardening intermediate configuration by adding the over round-hat (_ ). Now, assume the Neo-Hookean hyperelastic equation in Eq. (7.38) for the kinematic hardening analogously to the elastic deformation in Eq. (21.44), i.e.
21.4
Conventional Crystal Plasticity Model
609
_ _ wk Cksp ¼ lk trCksp 3 _
Sk ¼
_ @wk Cksp _
@Cksp
ð21:79Þ
¼ lk I
ð21:80Þ
where lk is the material constant. Then, it follows from Eq. (17.74) with Eqs. (21.80) and (17.62) that _
Mk ¼ ¼
_
_ _ _ C pks Sk ¼ 2lk C pks ; M k ¼ p p p 4lk FpT ks D D kd Fks
_
p p 2lk C pks ¼ 4lk FpT ks D ks Fks ð21:81Þ
Here, we adopt the following relation based on the storage part of the kinematic hardening rate in Eq. (21.68), noting Eqs. (21.25) and (21.60). ! n X p p 1 p b b b Dks ¼ D Dkd ¼ P ^nb a ð21:82Þ k b s b y k b¼1 The substitution of Eq. (21.81) with Eq. (21.82) into Eq. (21.78) leads to p p a s ak ¼ 4lk sa0 FpT ks D ks Fks n0
ð21:83Þ
leading to a
sk ¼
p p a 4lk sa0 FpT ks Dks Fks n0
¼
4lk sa0 FpT ks
n P
b
P
^nb
b¼1
!
1 bk sby
a
b
kb Fpks na0 ð21:84Þ
21.4.4
Stress Rate Versus Strain Rate Relation
The stress rate versus strain rate relation is formulated for the hypoelastic-base plasticity and the multiplicative hyperelastic-based plasticity in this section. (a) Hypoelatic-based plasticity The substitutions of Eqs. (21.62) and (21.77) into Eq. (21.59) leads to a
s ¼ ^n
a
n X b¼1
hab k_ b þ
n X b¼1
"
ck yab ^n b
1
ab bk sby
!
2^n p : b b
n X c¼1
#
ac c
z a
kb þ 2
n X
d : zab ab
b¼1
ð21:85Þ
610
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
which is rewritten as
sa ¼
n P b¼1
^ gab ^ n hab þ ck yab ^ n
1
b
ð21:86Þ
d : zab ab
b¼1
where a
n P
^ gab kb þ 2
! 2^ nb p b :
ab b
bk sy
n X
zac ac ¼ 6 ^gba
Equating Eq. (21.39) with Eq. (21.86), it follows that ! n n n X X X a b b b ^ N : d p ^ d : zab ab ¼ n k gab kb þ 2 b¼1
b¼1
^ ab k ¼ m
N 2
b
a
b¼1
where
ð21:88Þ
b¼1
which is rewritten as n X
n X
! ab b
z a
ð21:89Þ
:d
b¼1
^ ba ^ ab ^ gab þ Na : pb ^ 6 m nb ¼ m
ð21:90Þ
The positive plastic multiplier is given from (21.89) as follows: ! n n X X a 1 b bc c ^ ab N 2 z a :d m k ¼ Substituting Eq. (21.91) into Eq. (21.54), it follows that s ¼E:d
ð21:91Þ
c¼1
b¼1
w
ð21:87Þ
c¼1
n X n X
N ^ n
a a
^ 1 m ab
N 2 b
a¼1 b¼1
n X
! bc c
z a
:d
c¼1
leading to
^ ep : d sw ¼ K
ð21:92Þ
where ^ ep
K E
n X n X a¼1 b¼1
N ^ n a a
^ 1 m ab
N 2 b
n X c¼1
bc g
z a
!
^ 6¼ K
epT
ð21:93Þ
21.4
Conventional Crystal Plasticity Model
611
The loading criterion for the plastic shear strain rate is given by ( c pa 6¼ 0 for j^sa j ¼ say and sa s a [ 0 pa c ¼ 0 for others
ð21:94Þ
(b) Multiplicative hyperelastic-based plasticity The substition of Eqs.(21.62) and (21.84) into Eq. (21.59) leads to ! n n X X 1 b b p a b p b a a a pT b s ¼^ n hab k þ 4lk s0 Fks P ^n a k Fks n0 bk sby b¼1 b¼1
ð21:95Þ
which is rewritten as
sa ¼
n P
^ ab kb G
ð21:96Þ
b¼1
where
!
^ ab ^ G na hpab þ 4lk sa0 FpT ks P
b
^ nb c k
1 bk sby
^ ba ab Fpks na0 ¼ 6 G
ð21:97Þ
Equating Eq. (21.48) with Eq. (21.96), one has n X
(
^ ab k ¼ G b
2lsa0
" #) n e eX b b b b n na0 sym C L sym C P þQ k ^
b¼1
b¼1 n X ¼ 2lsa0 sym C e L na0 2l fsa0
e b b sym C P b þ Q b na0 ^ n kg
b¼1
ð21:98Þ leading to n X
^ ab kb ¼ 2lsa0 sym Ce L na0 M
ð21:99Þ
b¼1
where ^ ab M
n X ^ ab þ 2lsa sym C e P b þ Q b na ^nb 6¼ M ^ ba G 0 0 b¼1
It follows from Eq. (21.99) that
ð21:100Þ
612
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity 1 a ^ ba kb ¼ 2lM s0 sym C e L na0
and the substitution of this equation into Eq. (21.60) leads to 1 a ^ ba s0 sym C e L na0 ^nb c pb ¼ 2lM The substition of Eq. (21.102) into Eq. (21.25) leads: ( P b 1 a ^ s0 sym C e L na0 ^nb D p ¼ 2l nb¼1 P M ba P b 1 a ^ s0 sym C e L na0 ^nb W p ¼ 2l n Q M b¼1
ð21:101Þ
ð21:102Þ
ð21:103Þ
ba
and p
L ¼ 2l
n P b b ^ 1 sa0 sym Ce L na0 ^nb P þQ M ba
ð21:104Þ
b¼1
Substituting Eq. (21.102) into Eq. (21.55), one has ( " #) n X e e a b e b b 1 a ^ M ¼ 2l sym C L 4lsym C P þ Q Mba s0 sym C L n0 ^n b¼1
ð21:105Þ The stress can be calculated by the time-integration of Eq. (21.105). However, it can be done more effectively by the procedure: The plastic velocity gradient p is calculated first by inputting the velocity gradient tensor L into tensor L Eq. (21.104). Then, the rate of the plastic deformation gradient tensor is given p Fp and thus Fp is updated by the numerical calculation described in by F p ¼ L Sect. 17.12. Then, the Mandel stress tensor is given from Eq. (21.46) by M ¼ lFpT CFp1
ð21:106Þ
e
noting C ¼ FeT Fe ¼ FpT CFp1 . The Cauchy stress tensor is calculated from the Mandel stress as follows: r ¼ FeT MF
eT
ð21:107Þ
referring to Table 5.1. Further, assuming that the spin for the kinematic hardening is not induced as done for Eq. (17.104), the dissipative part of the kinematic hardening is given from Eqs. (21.82) and (21.101) as follows: p ¼ D p ¼ L kd kd
n X
b 1 ab kb P b bk s y b¼1
ð21:108Þ
21.4
Conventional Crystal Plasticity Model
613
leading to p
Lkd ¼ 2l
n P
b
P
b¼1
1 bk sby
e 1 a ^ ba ab M s0 sym C L na0
ð21:109Þ
p
p ¼ Lkd Fpkd using Eq. (21.108) and then one can Then, Fpkd is updated by Fkd _ _ p pT p calculate Fpks ¼ Fp Fp1 kd . Then, Mk is calculated by substituting C ks ¼ Fks Fks _
k ¼ FpT M k FpT into the first equation in Eq. (21.81). Mk is calculated as M ks ks by Eq. (17.76) and the kinematic-hardening variable a in the current configuration is calculated as k FeT a ¼ FeT M
ð21:110Þ
by the similar manner to Eq. (21.107).
21.5
Subloading Crystal Plasticity Model
Now, incorporate the shear subloading region for the shear yield region in Eq. (21.56), which is described by the following relation based on the subloading concept (Hashiguchi 2015b). j^sa j ¼ r a say ; i.e:r a
j^sa j say
ð21:111Þ
where r a ð0 r a 1Þ is referred to as the normal-yield shear ratio which designates always the ratio of j^sa j to the critical shear stress say not only in the slipping process pa c 6¼ 0 but also in the non-slipping process c pa ¼ 0 . The associated flow rule to the subloading shear region is adopted for the plastic shear strain rate as follows: na ka 0 c pa ¼ ka c pa ¼ ka ^
ð21:112Þ
Although this equation is just identical to Eq. (21.60), the resolved shear stress sa lies on or inside the limit of the normal shear yield region in Eq. (21.56). The material-time derivative of Eq. (21.111) reads: ^ na s a s ak ¼ ca sya þ r a say
ð21:113Þ
614
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
i.e.
sa ¼ ^ na r a s ay þ s ak þ ^ na r a say
ð21:114Þ
Referring to the Sect. 9.2 on the subloading surface model, let us postulate for the crystalline shear strain rate as follows: The crystalline plastic shear strain rate is induced when the resolved stress approaches the critical shear stress but only the crystalline elastic shear strain rate is induced when the resolved stress recedes from the critical shear stress, while the resolved stress rate causes the crystalline elastic shear strain rate inevitably. In other words, the resolved shear stress approaches the critical shear stress when a crystalline plastic shear strain rate is induced but it recedes from the critical shear stress when only a crystalline elastic shear strain rate occurs. Here, note that the approaching degree of the resolved shear stress to the critical shear stress is described by the shear normal-yield ratio r a . Then, let the evolution equation of shear normal-yield ratio r a be given analogously to Eq. (9.9) for the normal-yield ratio R of the subloading surface model as follows: ⎧ •α • • • ⎪ r = U ( r α ) |γ pα | = U ( r α ) λ α for γ pα ≠ 0 ⎨ ⎪r α = |τ^α | / τ αy for other ⎩
where the function Uðr a Þ fulfills the conditions (see Fig. 21.4): 8 ! þ 1 for r a ¼ 0 (elastic state) > > < [ 0 for 0\ra \1 (subyield state) U ðr a Þ ¼ 0 for ra ¼ 1 (normal-yield state) > > : \0 for r a \1 (over normal-yield state)
U (r α )
• pα
γ
•
= 0 : rα < 0
• pα
γ
0
1
•
0 : rα > 0 • pα
γ
≠ 0:
r• α < 0
rα
Fig. 21.4 Function Uðr a Þ in the evolution rule of shear normal-yield ratio r a
ð21:115Þ
ð21:116Þ
21.5
Subloading Crystal Plasticity Model
615
or 8 ! þ 1 for 0 ra re (elastic state) > > < [ 0 for re \r a \1 (subyield state) Uðr a Þ ¼ 0 for ra ¼ 1 (normal-yield state) > > : \0 for ra [ 1 (over normal-yield state)
ð21:117Þ
re ð\1Þ is the material constant. Let the explicit function of Uðra Þ fulfilling Eq. (21.117) be given analogously to Eq. (9.14) as follows: p hr a re i a ð21:118Þ U ðr Þ ¼ uc cot 2 1 re where uc is the material constant. The smooth shear stress versus crystalline shear strain curve is depicted and the resolved shear stress is automatically attracted to the critical shear stress because of
r a \0 for ra [ 1 (over shear normal-yield state) in Eq. (21.115) with Eq. (21.116)4 or (21.117)4 as shown in Fig. 21.5.
u c → ∞ : Conventional plasticity •
τ
τ yα0
•
•
Subloading plasticity
γα
0
r• α = U ( r α )| γ• pα | for γ• pα ≠ 0 ⎧> 0 for r α < 1 ⎪⎪ U (r α ) ⎨= 0 for r α = 1 ⎪ α ⎪⎩< 0 for r > 1 Fig. 21.5 Resolved shear stress is automatically attracted to critical shear stress in plastic shear process
616
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
(a) Hypoelastic-based plasticity Substituting Eqs. (21.62), (21.77) and (21.115) into Eq. (21.114), one has the consistency condition for the subloading crystalline shear region. a
s ¼ ^n r
a a
n X
hab k þ b
b¼1
n X
"
b¼1 n X
þ2
n ck yab ^
!
1
b
2^ n p :
b
bk sby
b b
a
n X
# ac c
z a
kb
c¼1
ð21:119Þ
d:zab ab þ ^ na U ðr a Þsay k a
b¼1
which is rewritten as
sa ¼
n X
n X
gab kb þ 2
b¼1
ð21:120Þ
d:zab ab
b¼1
where ! n r hab þ ck yab ^ gab ^ n a a
b
1
a b
2^nb pb :
b
bk s y a a a þ^ n U ðr Þsy dab 6¼ kba
n X
zac ac
c¼1
ð21:121Þ
d:zab ab
ð21:122Þ
Equating Eq. (21.39) with (21.120), it follows that N : d a
n X
!
p ^ n k b b
¼
b
b¼1
n X
gab kb þ 2
b¼1
n X b¼1
which is rewritten as Na :d 2
n X
d:zab ab ¼
b¼1
where
n X
mab kb
ð21:123Þ
b¼1
Tab ab gab þ Na : pb ^ 6 m nb ¼ m
ð21:124Þ
The plastic shear strain rate is given from Eq. (21.123) as follows:
ka ¼
n X b¼1
b 1 m ab ðN 2
n X c¼1
zac ac Þ:d
ð21:125Þ
21.5
Subloading Crystal Plasticity Model
617
Substituting Eq. (21.112) with Eq. (21.125) into Eq. (21.54), it follows that
s w ¼ E:d
n X
na Na ^
a¼1
n X
b 1 m ab ðN 2
n X
zac ac Þ:d
c¼1
b¼1
leading to
sw ¼ K ep : d
ð21:126Þ
where ep E K
n X n X
b 1 Na ^ na m ab ðN 2
a¼1 b¼1
n X
epT zac ac Þ ¼ 6 K
ð21:127Þ
c¼1
The loading criterion for the plastic shear strain rate is given by the sign of the plastic multiplier in terms of the shear strain rate as follows:
c pa 6¼ 0 for k a [ 0
) ð21:128Þ
c pa ¼ 0 for other
The stress calculation by the forward-Euler calculation method is performed as follows:
(1) First calculate the plastic multipliers ka by solving Eq. (21.123) for the input of the strain rate d.
(2) If k a is positive, calculate the plastic shear strain rate c pa ¼ na ka , the critical shear stress rate say by Eq. (21.62) and the kinematic hardening rate s ak by
Eq. (21.77), the rate of normal-yield shear ratio r a by Eq. (21.115), the resolved shear stress rate sa by Eq. (21.39), the Jaumann rate of the Kirchhoff w stress s by Eq. (21.54) and the rates of the base vectors s a , na of slip system by substituting Eq. (21.30) into Eq. (21.16). Then, update all these variables.
(3) If ka is negative, set s ay ¼ 0, s ka ¼ 0, and calculate s a by Eq. (21.39), s w by
Eq. (21.54) and s a , na by substituting Eq. (21.30) into Eq. (21.16) under setting c pa ¼ 0. Then, update all these variables. Thereafter, update r a by Eq. (21.111). (4) Move to the calculation for the next incremental step in which the updated values obtained in the above-mentioned processes are substituted.
618
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
(b) Multiplicative hyperelastic-based plasticity The substitution of Eqs.(21.62), (21.82), (21.83) and (21.115) into Eq. (21.114) leads to a
s ¼^ n r
a a
n X
hab k
b
þ 4lk sa0
FpT ks
b¼1
n X
P
1
^n ck
b
b
bk sby
b¼1
! a
b
k b Fpks na0 þ ^ na U ðra Þsay ka
ð21:129Þ which is rewritten as
sa ¼
n X
G ab kb
ð21:130Þ
b¼1
where !
Gab
^ na r a hab þ 4lk sa0 Fpks P b
^ nb c k
1 bk sby
ab Fpks na0 þ ^na U ðr a Þsay dab ¼ 6 gba ð21:131Þ
Equating Eq. (21.48) with Eq. (21.130), it follows that n X
(
ab kb ¼2lsa G 0
b¼1
" #) n eX b b b b sym C L sym C P þ Q k ^n na0
e
b¼1
n n h e i o X e b b ¼2lsa0 sym C L na0 2l sa0 sym C P þ Q na0 ^nb kb b¼1
ð21:132Þ leading to n X
M ab kb ¼ 2lsa0 sym C e L na0
ð21:133Þ
b¼1
where M ab Gab þ 2lsa0 sym C e P b þ Q b na0 ^nb 6¼ M ba
ð21:134Þ
It follows from Eq. (21.133) that
kb ¼ 2lM
1 a ba s0
sym C e L na0
ð21:135Þ
21.5
Subloading Crystal Plasticity Model
619
and the substitution of this equation into Eq. (21.60) leads to 1 c pb ¼ 2lM ba sa0 sym C e L na0 ^na
ð21:136Þ
The substition of Eq. (21.136) into Eq. (21.25) leads: 9 1 P b M ba sa0 sym C e L na0 ^na > > = b¼1 n P 1 > W p ¼ 2l Q b M ba sa0 sym C e L na0 ^na > ; D p ¼ 2l
n P
ð21:137Þ
b¼1
and L p ¼ 2l
n 1 P P b þ Q b M ba sa0 sym C e L na0 n^a
ð21:138Þ
b¼1
The stress can be calculated by the method described at the end of Sect. 21.4.4b. Further, the dissipative part of the kinematic hardening is given from Eqs. (21.108) and (21.135) as follows: p ¼ p ¼D L kd kd
n X
b 1 ab k b P b bk s y b¼1
ð21:139Þ
leading to p
Lkd ¼ 2l
n P b¼1
1
b
P
bk sbv
e 1 ab M ba sa0 sym C L na0
ð21:140Þ
p Then, Fpkd is updated by Fkd ¼ L pkd Fpkd using Eq. (21.140) and then one can _
calculate Fpks ¼ Fp Fp1 kd . Then, M, Mk and a are calculated by the identical method described at the end in Sect. 21.4.
21.6
Subloading-Overstress Crystal Plasticity Model
Based on the extension of the overstress model by the subloading surface model in Eq. (14.29), let the viscoplastic slip rate c vpa be given by
c vpa ¼ Ca ^ na
ð21:141Þ
620
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
where a n r rsa 1 na C ¼ lcs 1 ra =ðcm rs Þ
ð21:142Þ
a
or 1 exp n r a rsa 1 1 sinhðnhR Rs iÞ ; Ca ¼ C ¼ a lcs lcs 1 ra =ðcm rs Þ 1 r =ð cm r s Þ a
ð21:143Þ
where lcs is the material constant standing for the crystalline viscous coefficient, cm ð 1Þ is the material constant specifying the maximum value of r a and nð 1Þ is the material constant. The evolution rule of the static normal-yield ratio rSa 0 rSa 1 is given by 8 < rs a ¼ U rsa c vpa ¼ U rsa Ca for r a [ rsa ðCa ¼ 6 0Þ : rSa ¼ r a ¼ sa sak =say for others ðCa ¼ 0Þ
ð21:144Þ
following Eq. (21.115), where the function U rSa is given as a a p rS re U rS ¼ uc cot 2 1 re
ð21:145Þ
The smooth resolved shear stress versus crystalline plastic shear strain curve is described always as shown in Fig. 21.6. The isotropic hardening variable evolves following Eq. (21.62) as follow:
s ay ¼
n P b¼1
b hvp ab C
ð21:146Þ
p vp where hvp ab is given by Eq. (21.63) with the replacement of c to c as follows:
vp vp vp vp p hvp ab ¼ qh ðc Þ þ ð1 qÞh ðc Þdab ¼ hab
vp vp h ðc Þ for a ¼ b ¼ ð1 q 1:4Þ qhvp ðcvp Þ for a 6¼ b hvp ðcvp Þ is given by the function 2 hvp ðcvp Þ ¼ hvp 0 sech
vp hvp 0 c ss sy 0
ð21:147Þ
ð21:148Þ
21.6
Subloading-Overstress Crystal Plasticity Model
621
Impact loading • γα
τα
•
γ α increases Overstress (r α rαs )τ αy
Quasi-static loading • γα 0
r α = r sα: Quasi-static loading • γα 0 γα
0
Fig. 21.6 Resolved shear stress versus crystalline plastic shear strain curve predicted by subloading-overstress-crystal plasticity model
with cvp
n Z t n Z t X X c_ vpb dt ¼ Cb dt b¼1
0
b¼1
ð21:149Þ
0
vp vp a hvp and say , i.e. hvp 0 and sy0 are the initial values of h 0 ¼ h ð0Þ and sy0 ¼ sy ð0Þ, a a respectively, and ss is the saturation value of sy , i.e. ss ¼ sy ð1Þ. The rates of the isotropic hardening variable and the static normal-yield ratio are given by Eqs. (21.144) and (21.146) for both of the hypoelasticity and the multiplicative finite elastoplasticity.
(a) Hypoelastic-based viscoplasticity The kinematic hardening variable is given by
s ka
¼
n P
"
ck yab ^ n
b¼1
b
1 bk sby
n p : ap 2^ b b
n P c¼1
# ac c
b
z aC
þ2
n P
d : zab ab
b¼1
ð21:150Þ from Eq. (21.77).
s a and s w are given by Eqs. (21.39)and (21.54) with the replacement of the plastic shear strain rate c pb to the viscoplastic shear strain rate Cb ^nb as follows:
622
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
a
s ¼N : d a
n X
! b b
p C ^n
ð21:151Þ
Nb Cb ^nb
ð21:152Þ
b
b¼1
sw ¼ E : d
n X b¼1
The elastic velocity gradient is given following Eq. (21.30) by le ¼ l lp ¼ l
n X
pb þ qb Cb ^nb
ð21:153Þ
b¼1
The deformation analysis by the forward-Euler calculation method is performed as follows:
(1) For r a [ rsa , calculate the crystalline slip rate c vpa by Eqs. (21.142) or (21.143), the rate of critical shear stress s ay by Eq. (21.146), the rate of shear kinematic
hardening variable s ak by Eq. (21.150), the rate of resolved shear stress s a by Eq. (21.151), the rate of subloading shear ratio r as by Eqs. (21.144), the
Jaumann rate of the Kirchhoff stress s w by Eq. (21.152) and the rates of the base vectors s a , na of slip system by substituting Eq. (21.153) into Eq. (21.16). Thereafter, calculate ra by Eq. (21.111). (2) For the other leading to c vpa ¼ 0, set s ya ¼ 0, s ak ¼ 0, and calculate s a by
Eq. (21.151), s w by Eq. (21.152) and n a , s a by substituting Eq. (21.153) into Eq. (21.16) under setting c vpa ¼ 0. Then, update all these variables. Further, calculate r a by Eq. (21.111) and set rsa ¼ r a . (3) Move to the calculation for the next incremental step in which the updated values obtained in the above-mentioned processes are substituted. (b) Multiplicative hyperelastic-based viscoplasticity The kinematic hardening variable is given by a
sk ¼ from Eq. (21.83).
4lk sa0
FpT ks
n P b¼1
! b
P
^ n b
1 bk sby
a
b
Cb Fpks na0
ð21:154Þ
Substituting Eq. (21.141) into Eq. (21.24) with the replacement of c pb to c vpb , it follows that
21.6
Subloading-Overstress Crystal Plasticity Model
vp
L ¼2
n X
b
623
P þQ
b
Cb ^nb
ð21:155Þ
b¼1
The rate of the plastic deformation gradient tensor is given by Fvp ¼ L vp Fvp and then Fvp is updated by the numerical calculation described in Sect. 17.12. Then, the Mandel stress tensor is given by M ¼ lFvpT CFvp1
ð21:156Þ
from Eq. (21.106), where Fe is calculated by Fe ¼ FFvp1 . The Cauchy stress tensor is calculated by Eq. (21.107). Further, the dissipative part of the kinematic hardening is given from Eqs. (21.108) and (21.141) as follows: vp L vp kd ¼ Dkd ¼
n X
Pb
b¼1
1 bk sby
ab Cb
ð21:157Þ
vp vp vp Then, Fvp kd is updated by F kd ¼ L kd Fkd using Eq. (21.157) and then one can _vp _ vpT vp vp vp1 calculate Fvp ¼ F F . Then, M is calculated by substituting C ¼ F F ks ks kd ks ks _
k ¼ FpT Mk FpT in into the first equation in Eq. (21.81). Mk is calculated by M ks ks Eq. (17.76) and the kinematic-hardening variable a in the current configuration is calculated by Eq. (21.110).
21.7
Extension to Description of Cyclic Loading Behavior
The crystal plasticity model formulated in the preceding sections is based on the initial subloading surface model in which the similarity-center of the shear normal-yield and the shear subloading regions, i.e. the shear elastic-core is fixed at the shear kinematic hardening variable point. Therefore, unrealistically large shear strain accumulation is predicted, while open hysteresis loops are depicted as illustratively shown in Fig. 21.7. As described in Sect. 11.2, it will be extended to describe cyclic loading behavior by letting the similarity-center of the shear normal-yield and the shear subloading regions move with a plastic shear strain as shown in Fig. 21.7. Incorporating this fact, the extended shear crystalline model will be described in this section. The subloading shear region is given instead of Eq. (21.111) as follows: jsa j ¼ ra say
ð21:158Þ
624
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
τα
τα
0
0
γα
γα
Initial subloading model
Extended subloading model for cyclic loading Similarity-center (elastic-core)
Fig. 21.7 Modification of subloading model to describe cyclic loading behavior
where s a sa s ak
ð21:159Þ
sak stands for the conjugate (similar) point in the shear subloading shear region to the point sak in the normal-yield shear region. By letting sac denote the similarity-center of the shear normal-yield and the shear subloading regions, which is called shear elastic-core since the most elastic shear behavior is induced when the resolved shear stress lies on it fulfilling r a ¼ 0, the following relation holds by virtue of the similarity of the shear subloading region to the shear normal-yield region (see Fig. 21.8). ð21:160Þ sac s ak ¼ r a sac sak It follows from Eqs. (21.159) and (21.160) that sak ¼ sac r a^sac _a
s a ¼ s þ r a^sac
ð21:161Þ ð21:162Þ
where ^sac sac sak _a s sa sac
ð21:163Þ
21.7
Extension to Description of Cyclic Loading Behavior
τˆαy
625
τˆαy ( ≡ τ αy − τ kα) τ α (≡ r ατˆαy )
τα
(
τ kα τ kα
cα
cˆ α
)
c α τ) α τ α
Slip system
c α ≡ c α − τ kα = r α cˆ α cˆα ≡ cα − τ αk = ℜ cα (τ yα − τ kα ) ) τ α τ α cα
Fig. 21.8 Resolved and critical shear stresses, shear back-stress and shear elastic-core in slip system
The associated flow rule for the extended shear subloading region is adopted for the plastic shear strain rate as follows: c pa ¼ ka na ka 0 c pa ¼ ka
ð21:164Þ
where na
@ js a j sa ¼ ¼ signðs a Þðjn a j ¼ 1Þ @sa js a j
ð21:165Þ
In what follows, we limit to the multiplicative finite strain theory because the formulation in the hypoelasticity is rather complicated. The isotropic shear hardening rate is given by Eq. (21.62). The kinematic shear hardening rate is given by extending Eq. (21.84) as follows: a
sk ¼
4lk sa0 FpT ks
n P b¼1
b
P
nb
1 bk sby
! a
b
kb Fpks na0
ð21:166Þ
626
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
Now, let the following shear elastic-core region be introduced, which always passes through the shear elastic-core ca and maintains the similarity to the shear normal-yield surface with respect to the shear kinematic-hardening variable aa .
cˆ α = ℜ cα τ αy , i.e. ℜ cα = cˆ α / τ αy
ð21:167Þ
^c a ca aa
ð21:168Þ
where
ℜ cα designates the ratio of the size of the shear elastic-core region to the shear normal-yield region (see Fig. 21.8) so that let it be called the shear elastic-core yield ratio. Then, let it be postulated that the shear elastic-core can never reach the shear normal-yield region designating the fully-plastic shear state so that the shear elastic-core does not go over the following limit shear elastic-core region. ^c a ¼ vc say
ð21:169Þ
where vc ð\1Þ is material parameter and the following inequality must be satisfied.
cˆ α ≤ χcτ αy , i.e. ℜ cα ≤ χ c
ð21:170Þ The evolution rule of the shear elastic-core is given simply as follows (Hashiguchi 2017): α
α
c• α = cc ( γ• pα− ℜχc | γ• pα | nˆ cα ) = cc λ α ( n α − ℜχc nˆ cα ) •
c
c
ð21:171Þ
where cc is the material constant and ^ nac
^ca a ^ nc ¼ 1 a j^c j
ð21:172Þ
noting that ca does not go over the limit shear elastic-core region, satisfying the inequality Eq. (21.170) as known from α
nˆ cα c• α = cc ( γ• pα− ℜχc | γ• pα | nˆ cα ) = cc λ• α (nˆ cα n α −1) ≤ 0 for ℜ cα = χ c c
ð21:173Þ
We adopt the following relations analogously to Eq. (21.78). ( ( (• (• τ cα = sα0 • M cn α0 = G α0 : M c , τ• αc = sα0 • M c n α0 = G α0 : M c
ð21:174Þ
designating the tensor based in the elastic-core intermediate configuration by adding the over inverse-round-hat ð^ Þ.
21.7
Extension to Description of Cyclic Loading Behavior
627
Here, assume the Neo-Hookean hyperelastic equation in Eq. (7.38) for the elastic-core analogously to the kinematic hardening in Eq. (21.79), i.e. ^p
^p
^p
^
w ðCcs Þ ¼ lc ðtrCcs 3Þ; S c ¼ c
@wc ðCcs Þ ^p
@Ccs
¼ lc I
ð21:175Þ
where lc is the material constant. Then, it follows analogously to Eq.(21.181) that ^p ^
^
^p
^
p p Mc ¼ Ccs Sc ¼ 2lc Ccs ; M c ¼ 4lc FpT cS D cs Fcs
ð21:176Þ
Here, it follows analogously to Eq. (21.82) with Eq. (21.171) that n
D cs = ∑ P p
β =1
β
β
(n β − ℜχ c nˆ cβ ) λ• β
ð21:177Þ
c
The substitution of Eq. (21.176) with Eq. (21.177) into Eq. (21.174) leads to n
β
β =1
c
pT p ℜ τ• c = 4 μ csα0 • F cspT D csp Fcs n α0 = 4 μc sα0 • Fcs ∑ P β (n β − χ c nˆ cβ ) λ Fcsn α0 α
p
•β
ð21:178Þ
The material-time derivative of Eq. (21.158) reads: n a s a s ak ¼ r a s ay þ r a say
ð21:179Þ
where s ak is described from Eq. (21.161) as
s ak ¼ r a s ak þ ð1 ra Þ s ac r a ^sac
ð21:180Þ
Substituting Eq. (21.180) into Eq. (21.179), one has na s a r a s ak þ ð1 r a Þ s ac r a ^sac ¼ r a s ay þ r a say
ð21:181Þ
which is rewritten as
n a s a ¼ n a r a s ak þ n a ð1 ra Þ s ac n a r a ^sac þ r a s ay þ r a say
ð21:182Þ
s a ¼ n a r a s ay þ ca s ak þ ð1 ra Þ s ac þ n a say ^sac r a
ð21:183Þ
i.e.
628
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
Here, noting the relation na say
^sac
a _a s sak sac sak sa jsa j sac sa sac s ¼ a a a¼ ¼ ¼ ac a a r r r r js j r
ð21:184Þ
Equation (21.183) is rewritten as
s a ¼ na r a s ay þ r a s ak þ ð1 ra Þ c a þ
ra _ a s ra c
ð21:185Þ
Substituting Eqs. (21.62), (21.115), (21.166) and (21.178) into Eq. (21.185), it follows that
which is rewritten as
sa ¼
n X
~ ab kb G
ð21:187Þ
b¼1
where 1 p α β )Fkpsn α0 G% αβ ≡ n α r α hαβ + 4rα μk sα0 • FksT P β ( n β − b kτ yβ α α pT + (1 − r α ) μk sα0 • Fcs P β ( n β − ℜ c nˆ cβ ) Fcsp n α0 + U ( rα ) τ)cα δαβ χ r
c
(21.188)
The full constitutive relation in the hypoelasticity is quite complicated. Then, the calculation method will be shown in the multiplicative finite strain theory below. The following relations hold analogously to Eqs. (21.135)–(21.138). e a 1 a L n ; ~ ba S0 sym C kb ¼ 2lM 0
p ¼ 2l D
e a a 1 a L n n ~ ba c pb ¼ 2lM s0 sym C 0
e a a9 L bM n n > ~ 1 sa0 sym C > P > 0 ba = b¼1
p ¼ 2l W
ð21:189Þ
n P
n e a a> P bM L n n > ~ 1 a sym C > Q ; 0 ba
b¼1
ð21:190Þ
21.7
Extension to Description of Cyclic Loading Behavior n X
629
a a 1 a b M L b þ Q n n ~ ba s0 sym C P 0
ð21:191Þ
e b P b na nb 6¼ M þQ ~ ab þ 2lsa sym C ~ ba G 0 0
ð21:192Þ
p ¼ 2l L
b¼1
where ~ ab M
n X b¼1
The stress and the kinematic hardening variable can be calculated by the method described at the end of of Sect.21.5(b). Further, the elastic-core Mc and c are calculated by the similar manner used for the kinematic hardening variable Mk and a.
21.8
Uniqueness of Slip Rate Mode
Hill and Rice (1972) proved that the sufficient condition for the uniqueness of the combination of slip rates is the positive-definiteness of the matrix in Eq. (21.124) in all the slip systems as shown in the following, while the superscript p specifying the plastic sliding is omitted since the rigid-plastic sliding is considered. Suppose to impose the two strain rates d and d, and designate the slip rates for
these strain rates as c a and c , respectively, in the two slip modes, and denote their differences as follows:
D ca ¼ ca ca
ð21:193Þ
Dd ¼ dd
ð21:194Þ
The following inequality must be satisfied from Eqs. (21.59) and (21.62), ignoring the kinematic hardening for simplicity.
na s a
n X
hab c b na
ð21:195Þ
b¼1
where na
@ j sa j sa ¼ a ¼ signðsa Þðjna j ¼ 1Þ a @s js j
while the equality holds when all the slip systems are activiated.
ð21:196Þ
630
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
The substitution of Eq. (21.39) into Eq. (21.195) leads to N : d a
n X
!
b
na
b
p c
b¼1
n X hab c b na
ð21:197Þ
b¼1
resulting in " N : d a
n X
!
b b
p c
b¼1
#
n X
b
na 0
hab c
b¼1
which is decsibed as N :d a
n X
! b
na 0
Mab c
ð21:198Þ
b¼1
where Mab hab þ Na : pb
ð21:199Þ
First, we assume that the slip system a is active in both modes, i.e. a
c 6¼ 0;
c a 6¼ 0, so that the following equations hold from Eq. (21.198). N :d a
n P
!
Mab c
9 > > 6 0> n ¼ 0 for c ¼ > > > =
b
b¼1
Na :d
n P
a
a
!
> > > > a > 6 0> n ¼ 0 for c ¼ ;
ð21:200Þ
a
Mab c b
b¼1
which leads to Na : Dd
n X
Mab D c b ¼ 0
ð21:201Þ
b¼1
where Dd dd and D c b c b c b . Multiplying D c a to Eq. (21.201), we have
Na :DdD c a ¼
n X
Mab D c b D c a
ð21:202Þ
b¼1
Next, we assume that the slip system a is active in the first mode but it is inactive
in the second mode, i.e. c a 6¼ 0; c a ¼ 0, so that
21.8
Uniqueness of Slip Rate Mode
631
9 b a a > = M c ¼ 0 for c ¼ 6 0 n ab b¼1 P n Mab c b na \0 for c a ¼ 0 > ; Na :d b¼1
Na :d
Pn
ð21:203Þ
which leads to !
n X
N :Dd a
b
Mab D c
na [ 0
ð21:204Þ
b¼1
Noting
na D c a ¼ na c a [ 0 Eq. (21.204) leads to N :Dd a
n X
! b
Mab D c
na na D c a [ 0
b¼1
from which it follows that n X
Na :DdD c a [
Mab D_cb D c a
ð21:205Þ
b¼1
Consider the inverse case that the slip system a is inactive in the first mode and
active in the second mode, i.e. c a ¼ 0; c a 6¼ 0, so that Na : d
n P b¼1
N :d a
n P
!
Mab c_ b na \0 for c ! _ b
Mab c
b¼1
a
9 > > ¼0> > =
> > > 6 0> n ¼ 0 for c ¼ ; _a
a
from which it follows that Na :Dd
n X
! b
Mab D c
na \0
b¼1
for which, noting
na D c a ¼ na c a \0
ð21:206Þ
632
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
it follows that N :Dd a
n X
!
Mab D c
b
na n a D c a [ 0
b¼1
Consequently, the inequlaity in (21.205) holds also in this case. Furthermore, in the case that the slip system a is inactive in both of the first and
the second modes, i.e. c a ¼ 0; c a ¼ 0 leading to D c a ¼ 0, so that the equality in Eq. (21.202) holds. Consequently, the follwing inequality holds in all the above-mentioned four cases.
Na : DdD c a
n X
Mab D c b D c a
ð21:207Þ
b¼1
Taking the total sum of Eq. (21.207) for all of slip systems, the following inequality holds. n X a¼1
Na :DdD c a
n X n X
Mab D c b D c a
ð21:208Þ
a¼1 b¼1
where Dd ¼ O holds for given strain rate, and if Mab is the positive-deinite matrix, the right-hand side in Eq. (21.208) is non-negative, so that the uniqueness of slip mode, i.e. D c a ¼ 0 holds. In other words, Mab must be positive-definite in order that the uniqueness of slip mode, i.e. the uniqueness of c a holds for given strain rate d. Then, the matrix Mab is called the effective slip-systems hardening moduli. The matrix Mab is asymmetric, i.e. Mab 6¼ Mba and thus it is not the positive-definite marix, so that the uniqueness of slip mode does not holds in general. The uniqueness of the matrix Mab depends on the hardening coefficient, state of stress and the number and the directions of critical shear stress sensitively. It is not guaranteed and its tendency is remarkable for a higher latent hardening (Hill 1966; Hill and Rice 1972; Havner 1982; Asaro 1983; Franciosi and Zaoui 1991).
21.9
Various Schemes for Calculation of Shear Strain Rates
Big computational time is required for calculating shear strain rates in numerous slip systems. Then, various schemes for the improvement of the calculation have been proposed to date. Main schemes for the improvement will be sown in this section.
21.9
Various Schemes for Calculation of Shear Strain Rates
21.9.1
633
Singular Value Decomposition
It is required to solve Eq. (21.89) in order to calculate shear strain rates in slip systems directly from macroscopic strain rate applied to crystalline. However, the ^ ab is not positive-definite, so that there does not exist a unique solution in matrix m general as described in the last section. The singular value decomposition is used to calculate the solution with the shortest path (Golub and Van Loan 2013; Press et al. 1988). It has been applied to the crystal plasticity by Anand and Kothari (1996) and used widely by Miehe and Schroder (2001), Knockaert et al. (2000), Yoshida and Kuroda (2012), etc. The singular value decomposition is explained in this subsection. The general second-order tensor T in the n-dimensional space which is asymmetric and thus cannot be led to the spectral decomposition in general. On the other hand, designating the eigenvectors of the symmetric positive-definite tensors TTT and TT T as uq and vq ðq ¼ 1; 2; ; nÞ, respectively, and their eigenvalues r2q because they are positive, it follows that TTT ¼ T T¼ T
n P q¼1 n P q¼1
9 > r2q uq uq > = > r2q vq vq > ;
ð21:209Þ
i.e. TTT uq ¼ r2q uq TT Tvq ¼ r2q vq
) ð21:210Þ
while uq , vq and rq are not the eigenvectors and eigenvalues except for the symmetric tensor T. Exploiting the eigenvalues and vectors defined above, the tensor T can be led to the following singular value decomposition. ð21:211Þ
T ¼ URVT
where R is the diagonalized tensor with the components r1 ; r2 ; ; rn and thus it is described as follows: ⎡σ 1 0 0 ⎤ ⎢0 σ ⎥ 0⎥ 2 ⎢ ð21:212Þ Σ = diag(σ 1 , σ 2 , • ••, σ n ) = ⎢ ⎥ ⎢ ⎥ ⎢⎣0 0 σ n ⎥⎦
L L M MOM L
634
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
provided that the order of the magnitudes of these components is U and V are the orthogonal tensors by lining the eigenvectors in column (horizontally).
U = ⎢⎣u1
V = ⎣⎢ v1
L u n ⎤⎥ ⎫⎪ L u n ⎥⎪ uL ⎪ M M M O M ⎥⎥ ⎪ L unn ⎥⎦ ⎪⎪ ⎬ ⎧ v ⎫ ⎡v v Lv n ⎤ ⎪ ⎥ ⎪ v ⎪ ⎢v v Lv n ⎥ ⎪⎪ ⎪ ⎪ ⎢ v L v n ⎦⎥ = ⎨ ⎬ = ⎢ ⎥⎪ ⎪ M ⎪ ⎢M M O M ⎥ ⎪ 2
⎧u1 ⎫ ⎡u11 u12 ⎪u ⎪ ⎢u u ⎪ 2 ⎪ ⎢ 21 22 u n ⎥⎦ = ⎨ ⎬ = ⎢ ⎪ ⎪ ⎢ ⎪u ⎪ ⎢ ⎩ n ⎭ ⎣un1 u n2 1
11
12
2
21
22
1
2
ð21:213Þ
1
2
2
⎪v ⎪ ⎩ n⎭
⎢v v ⎣ n1 n2
L vnn ⎥⎦ ⎪⎭
where the components of each vector are lined up in row (vertically), satisfying UUT ¼ UT U ¼ VVT ¼ VT V ¼ IðkUk ¼ kVk ¼ 1Þ
ð21:214Þ
noting UUT ij ¼ Uir Ujr ¼ ðui er Þ uj er ¼ ui uj er er ¼ ui uj ¼ dij . It follows from Eqs. (21.211) and (21.214) that TTT ¼ URVT VRUT ¼ UR2 UT TT T ¼ VRUT URVT ¼ VR2 VT
ð21:215Þ
The pseudo-inverse tensor Ty of T is defined by Ty VRy UT where
Σ† = diag(1/ σ 1 , 1/σ 2 ,
L1/σ r , 0L
ð21:216Þ
L L0⎤⎥ L L0⎥ ⎥ M M O M M M⎥ ⎥ L L 0⎥ L L 0⎥⎥ M M M MOM ⎥⎥ L L0⎥⎦
⎡σ 1−1 0 00 ⎢ −1 00 ⎢0 σ 2 ⎢ ⎢ ⎢ σ r−1 0 0) = ⎢0 0 ⎢0 0 0 0 ⎢ ⎢ ⎢ 0 0 ⎢⎣0 0
ð21:217Þ
21.9
Various Schemes for Calculation of Shear Strain Rates
635
provided that we set 1=ri ¼ 0ði ¼ 1; 2; ; rÞ for ri ¼ 0ði ¼ r þ 1; 2; ; nÞ, obviously fulfilling RRy ¼ I. It is confirmed that the following equation holds from Eqs. (21.211), (21.214) and (21.216). TTy ¼ URVT VRy UT ¼ I
ð21:218Þ
Consider the following tensor equation with the vectors x and c. Tx ¼ c
ð21:219Þ
If T is the singular tensor, there exist numerous solutions for x. The vector x is expressed noting Eqs. (21.216) and (21.218) as follows: x ¼ Ty c ¼ VRy UT c ¼ Vdiag 1=rj UT c
ð21:220Þ
which is called the singular value decomposition and calculated from the right to the left. The solution obtained by the singular value decomposition is unique and possesses the shortest path among numerous solutions satisfying the original equation as will be proved as follows: There exists the zero-dimensional subspace of the vector x projected to zero vector if T is the singular tensor. Designating the solution of Eq. (21.220) for an arbitrary tensor Ry possessing non-zero component only in the zero component in
Eq. (21.217) by x0 , it follows for the vector x þ x0 noting Eqs. (21.216) and (21.220) that kx þ x0 k ¼Ty c þ x0 ¼ VRy UT c þ x0 ¼V Ry UT c þ VT x0 ¼ kVkRy UT c þ VT x0 ð21:221Þ ¼Ry UT c þ VT x0
The components in the second term are non-zero in the zero components in the first terms and they are orthogonal and thus independent to each other in the ndimensional orthogonal coordinate system, so that the minimum of the quantity in Eq. (21.221) holds for x0 ¼ 0 leading to the vector x as the singular value decomposition. Consider the following tensor based on the above-mentioned singular value decomposition. T ¼ ðT þ eIÞ
ð21:222Þ
636
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
where T is the nonsingular tensor made by adding the infinitesimal perturbation tensor eI to the singular tensor T. Adopting it in Eq. (21.219), we have x ¼ ðT þ eIÞ1 c
ð21:223Þ
from which we can obtain the similar solution to that due to the singular value decomposition. This is the simplest solution of the singular equation and called the diagonal shift method (Miehe and Schroder 2001). Number of unknown quantities on shear strain rates is larger than nine given equations on macroscopic strain rate components in the crystal plasticity, so that solution is not determined uniquely. In such case, we may solve by supplementing the lines composed of zero components to the matrix T and the zero components to the vector c for the difference of numbers in unknown quantities and given equations. By applying the singular value decomposition to the simultaneous equation made by this method, we can obtain the solution with the shortest path. It corresponds to the minimum shear principle by Taylor (1938).
21.9.2
Regularized Schmid Law
The yield surface in the slip systems is formed by plural intersecting planes in the stress space so that it possesses the sharp corner. Then, the shear strain rates must be calculated in each slip systems. In order to avoid this complicated work, the regularized Schmid law has been studied, in which the yield surface with rounded-off corners is formulated and the only one plastic multiplier by applying the associated flow rule to that yield surface is calculated. The yield condition in slip system is described by j sa j 1¼0 say
ð21:224Þ
The slip systems described by Eq. (21.224) give rise to the yield surfaces composed of the planes in number of slip systems, exhibiting the sharp corners at their intersections. A single smooth yield surface can be formulated by smoothing the envelope of the yield surfaces in Eq. (21.224) (Gambin 2001; Gambin and Barlat 1997; Darrieulat and Piot 1996; Zamiri et al. 2007; Zamiri and Pourboghrat 2010). This is regarded as the invocation of the method to derive the Mises yield surface from the Tresca yield condition by Hosford (1974, 2009). Zamiri et al. (2007) proposed the following simple yield surface.
21.9
Various Schemes for Calculation of Shear Strain Rates
2n1 11=m 0 2n1 11=m n n a a a a X X s s s p :s f ¼@ 1A ¼ @ 1A ¼ 0 a sa a sa s s y a¼1 y y a¼1 y
637
0
ð21:225Þ
noting Eq. (21.32). mð 1Þ is the material constant for smoothing the corner of yield surface, while the larger m is, the smoother the yield surface is. While Eq. (21.225) is of the power form, in order to avoid the problem on the numerical calculation caused by the power function, Zamiri and Pourboghrat (2010) proposed the yield surface in the logarithm-exponential function. ( " ! #) pa : s n X 1 ¼0 ð21:226Þ exp q a 1 f ¼ ln sy q a¼1 where qð 1Þ is the material constant for smoothing the corner of yield surface, while the larger q is, the smoother the yield surface is. The plastic strain rate for the above-mentioned yield condition with the associated flow rule is given by
dp ¼ k
@f ðk 0Þ @s
ð21:227Þ
The plastic multiplier k is calculated by the formulation based on the consistency condition of the yield condition (Gambin 2001; Gambin and Barlat 1997) or the return-mapping (Zamiri et al. 2007; Zamiri and Pourboghrat 2010). The above-mentioned regularized Schmid law reduces the calculation time since only one plastic multiplier for the single yield surface has only to be calculated. However, it would deviate from the primary purpose of the crystal plasticity for deriving the macroscopic behavior from the microscopic physical law, since it replaces the yield conditions in multi slip systems to the global yield surface.
21.9.3
On Creep-Type Crystal Plasticity Model
The crystal elastoplastic constitutive equation described in the last section possesses the difficulties: (1) The yield judgment whether the resolved shear stresses sa reach the critical shear stress say is required. (2) Particular algorithm to pull-back the resolved shear stresses to the critical shear stress must be incorporated. These procedures must be executed in numerous slip systems and thus the analysis by this constitutive equation is so complicated as actually impossible.
638
21 Hypoelastic- and Multiplicative Hyperelastic-Based Crystal Plasticity
Then, the crystal plasticity analyses by the creep model is widely used as will be described below. Nakada and Keh (1966) first advocated and ten years later Hutchinson (1976) presented the following creep-type rate-dependent equation of crystalline slip rate, which are widely adopted after the review report by Peirce et al. (1982, 1983) for single crystals and Asaro and Needleman (1985) for polycrystal. ca
c
¼ c ca 0
sa say
! ð1=mÞ1 sa a sy
ð21:228Þ
where c ca 0 is the reference rate of shearing which is taken usually same in all slip systems and m is the material constant, while for m\0:02 the creep slip rate c ca is a induced abruptly when the magnitude of shear stress js j reaches the shear-yield stress say . All the slip systems are active and thus the selection of active ones is not required for Eq. (21.228). However, Eq. (21.228) possesses the following fundamental impertinences. (1) The creep strain rate depends only on the shear stress because it falls within the framework of the creep-type viscoplasticity. Therefore, the creep crystalline slip rate is determined only by the current shear stress with time and thus an arbitrary deformation rate cannot be given to the material obeying Eq. (21.228) which is independent of the stress rate as far as a large elastic deformation of crystal lattice is not incorporated. In other words, it cannot be applied to rigid-viscoplastic materials. In facts, however, the crystalline slip rate would have to depend not only on the shear stress but also on its rate in a quasi-static deformation at room temperature. On the other hand, the rate-independent equation of plastic slip rate c pa must depend on both the shear stress and its rate (magnitude) in the slip system because it falls within the framework of the plasticity. (2) It belongs to the creep type model without a definite yield stress among the viscoplastic models. Therefore, it predicts always the creep deformation except for the stress-free state, and thus the creep slip rate is induced even when the magnitude of shear stress, jsa j, decreases as illustrated in Fig. 21.9 in which a general trend is intelligibly shown for a moderate value of m. It is unrealistic as known from the fact that any metallic solid has never collapsed under their own weight at room temperature as was indicated in p. 201 in Havner (1992). Needless to say, it cannot be applied to the description of cyclic loading behavior since it does not possess a loading criterion and thus it predicts identical deformation behavior in the reloading process and in the unloading process. Therefore, it predicts excessively large mechanical ratcheting as illustrated in Fig. 21.9 for the pulsating loading. In general, the viscoplastic deformation behavior cannot be described pertinently by the creep-type model and instead it can be predicted realistically by the overstress-type model possessing the loading criterion as explained in Chap. 14.
21.9
Various Schemes for Calculation of Shear Strain Rates
639
τα
0
γ cα
Fig. 21.9 Shear stress versus crystalline slip curves accompanied with excessive mechanical ratcheting predicted for pulsating loading by creep model (Peirce et al. 1983)
As examined above, the creep-type equation of crystalline slip contains fundamental defects. Then, it has been strongly desired over the last half century to find the physically and numerically pertinent rate-independent equation of crystalline slip rate. This fact is declared emotionally as “The various viscoplastic, finite-strain aggregate calculations reviewed in this section, and similar ones in the literature, are computationally impressive (although which approximate polycrystal model is superior appears to be an open question). However, one hopes that there may soon evolve a theory of rate-dependent crystalline slip in metals that would leave such great structural landmarks as the Eiffel Tower, Empire State Building, and Golden Gate Bridge still standing” in p. 204 of Havner (1992). The landmark would have been found in the subloading crystal plasticity model described in Sect. 21.7. All of the afore-mentioned defects in the rate-dependent crystalline slip rate are dissolved by subloading crystal plasticity model. The crucial importance for the incorporation of the subloading concept is manifested most distinctly in the crystal plasticity analysis which is required to calculate slips in numerous number of slip systems, while the yield judgment is not required and the stress is automatically attracted to the yield surface only in the subloading crystal plasticity model.
Chapter 22
Constitutive Equation for Friction: Subloading-Friction Model
All bodies in the natural world are exposed to friction phenomena, contacting with other bodies, except for bodies floating in a vacuum. Therefore, it is indispensable to analyze friction phenomena rigorously in addition to the deformation behavior of bodies themselves in analyses of boundary value problems. The friction phenomenon can be formulated as a constitutive relation between sliding displacement vector and contact traction vector, or their rates, in a similar form to the elastoplastic constitutive equation of materials. A constitutive equation for friction with the transition from the static to the kinetic friction and vice versa and the orthotropic and rotational anisotropy is described in this chapter. The stick–slip phenomenon is also delineated, which is an unstable and intermittent motion caused by the friction and thus important for the prediction of earthquake, vibration of machinery, etc.
22.1
History of Constitutive Equation for Friction
Friction influences all the natural phenomena as described above and it is the typical irreversible phenomenon connected directly to the second law of thermodynamics, so that it has been studied by numerous natural scientists since ancient times. Aristotle in ancient Greek has recognized the difference between the static friction and the kinetic friction. The term “Tribology” was coined by Leonardo da Vinci and the term “Friction” was used first by Isaac Newton. The term “radiation friction” was coined by Albert Einstein. Thus, the friction phenomenon has been studied extensively over the long years. Then, the terms “static friction” and “kinetic friction” are described even in the text books in middle and high schools. However, their mathematical expression had to wait until the proposition of the subloading-friction model by Hashiguchi et al. (2005). Still now, however, only the classical Coulomb friction is described in literatures except for the book by the present author and installed in the commercial FEM software, e. g. Abaqus, Marc, © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4_22
641
642
22
Constitutive Equation for Friction: Subloading-Friction Model
Nastran, LS-Dyna, etc. The recent history of the studies on the friction will be reviewed in the following. Constitutive equations of friction within the framework of elastoplasticity were formulated as rigid-plasticity (Seguchi et al. 1974; Fredriksson 1976). Subsequently, they were extended to elasto-perfect-plasticity (Michalowski and Mroz 1978; Oden and Pires 1983a, b; Curnier 1984; Cheng and Kikuchi 1985; Kikuchi and Oden 1988; Wriggers et al. 1990; Peric and Owen 1992; Anand 1993; Zhong 1993; Mroz and Stupkiewicz 1994; Wriggers 2003). Further, isotropic hardening was introduced (Oden and Martines 1986; Gearing et al. 2001) to describe the test results (cf. e.g. Courtney-Pratt and Eisner 1957). However, the interior of the sliding-yield surface was assumed as an elastic domain. Therefore, the plastic sliding velocity induced by the rate of contact traction inside the sliding-yield surface cannot be described. Needless to say, the accumulation of plastic (irreversible) sliding displacement induced by the cyclic loading of contact traction within the sliding-yield surface cannot be described by these models. They might be called the conventional friction model in accordance with the classification of plastic constitutive models for deformation of solids by Drucker (1988). Therefore, they are incapable of describing an accumulation of sliding displacement during cyclic loading of contact stress (traction) in addition to the incapability of describing the transition from the static to the kinetic friction and the recovery of the static friction. In addition, the simple friction model (Anand, 1993) falling within the framework of the creep model without the sliding yield surface was formulated but it is unacceptable in the general sliding velocity, since the creep is induced in any low stress level as was explained in Sect. 14.3 and thus it was applied only to the simulation of the sliding behavior at the high sliding velocity (Gearing et al., 2001). Eventually, it is irrelevant to the usual sliding behavior including the quasi-static sliding. It is widely known that when bodies at rest begin to slide, a high friction coefficient appears first, which is called the static friction. Subsequently, a friction coefficient decreases approaching a stationary value, called the kinetic friction. Furthermore, if the sliding ceases for a while and then starts again, the friction coefficient recovers and similar behavior to that of the initial sliding is recovered as has been indicated by some workers (Dokos 1946; Rabinowicz 1951, 1958; Howe et al. 1955; Derjaguin et al. 1957; Brockley and Davis 1968; Kato et al. 1972; Horowitz and Ruina 1989; Ferrero and Barrau 1997; Bureau et al. 2001). The recovery of friction coefficient has been formulated using equations directly including a time elapsed after the stop of sliding (cf. Rabinowicz 1951; Howe et al. 1955; Brockley and Davis 1968; Kato et al. 1972; Horowitz and Ruina 1989; Bureau et al. 2001). In past relevant studies, however, the inclusion of time itself leads to the loss of objectivity in constitutive equations, as is evident from the fact that the evaluation of elapsed time is accompanied with the ambiguity in the judgment about when the sliding commences and ceases, especially in the state that the sliding velocity varies in a low-velocity regime. Here, it should be noticed that the friction phenomenon has to be described in terms of the sliding velocity, the contact traction and internal variables without the inclusion of time itself.
22.1
History of Constitutive Equation for Friction
643
The reduction of the friction coefficient from the static to kinetic friction and the recovery of the friction coefficient as described above are the fundamental characteristics in friction phenomena, which have been widely recognized for a long time. Difference of the static and kinetic frictions often reaches up to several tens of percent, and thus the formulation taking account of these characteristics is of importance for analyses of practical problems in engineering. The constitutive equation describing these fundamental friction behavior has been formulated based on the subloading surface model and thus it is called the subloading-friction model (Hashiguchi et al. 2005; Hashiguchi and Ozaki 2008a). In addition, the difference of friction coefficients is observed in opposite sliding directions. It can be described by the rotation of a sliding-yield surface in the contact traction vector space, noting that the anisotropy of soils has been described by the rotation of a yield surface, as described in Sect. 22.11. Further, the range of friction coefficient depends on the sliding direction. It would be describable by the concept of orthotropy of the sliding-yield surface (Mroz and Stupkiewicz 1994). The subloading-friction model has been extended so as to describe these anisotropy (Hashiguchi 2006; Hashiguchi and Ozaki 2008b; Ozaki et al. 2012). The rate-and-state friction model (cf. e.g. Dieterich 1978, 1979; Ruina 1980, 1983; Rice and Ruina 1983; Scholz 1998; Rice et al. 2001; Kame et al. 2013) the basic idea of which is the dependence of the contact shear stress or friction coefficient on the rate of sliding and some state variables based on experimental data (cf. Dieterich 1979; Ruina 1980) has been widely used for the prediction of earthquake phenomena. An earthquake is a typical irreversible phenomenon which can be described appropriately by elasoplasticity but the rate-and-state friction model is not based on elastoplasticity. The earthquake disaster prevention cannot be attained forever as far as the rate-and-state model is used for the prediction of earthquakes. It should be emphasized that the irreversible deformation phenomena can be formulated rigorously within the framework of the elastoplasticity which has been developed to describe the irreversible deformation of solids in the history of the applied mechanics. Consequently, the subloading-friction model is formulated within the framework of the elastoplasticity aiming at establishing the rate-independent and -dependent constitutive equation of friction with the generality and the universality in the three-dimensional finite sliding under monotonic and cyclic frictional loadings.
22.2
Sliding Displacement and Contact Traction
The sliding displacement vector u, which is defined as the sliding displacement of the counter (slave) body to the main (master) body, is orthogonally decomposed into the normal sliding displacement vector un and the tangential sliding displacement vector ut to the contact surface as follows:
644
22
Constitutive Equation for Friction: Subloading-Friction Model
u ¼ un þ ut
ð22:1Þ
where
un ¼ ðu nÞn ¼ ðn nÞu ¼ un n ut ¼ u un ¼ ðI n nÞu
ð22:2Þ
n being the unit outward-normal vector of the surface of main body and (
un n un ¼ n u ut ¼ tu u ¼ kut k; tu
ut ðn tu ¼ 0; ktu k ¼ 1Þ kut k
ð22:3Þ
The minus sign is added for un to be positive when the counter body approaches the main body. The sliding displacement vector u can be exactly decomposed into the elastic (reversible) sliding displacement ue and the plastic (irreversible) sliding displacement up in the additive form even for the finite sliding displacement, i.e. u ¼ ue þ up ( e u ¼ uen þ uet up ¼ upn þ upt
ð22:4Þ ð22:5Þ
where (
uen ¼ ðue nÞn ¼ ðn nÞue ¼ uen n uet ¼ ue uen ¼ ðIn nÞue
(
upn ¼ ðup nÞn ¼ ðn nÞup ¼ upn n upt ¼ up upn ¼ ðIn nÞup
ð22:6Þ
ð22:7Þ
setting uen n uen ¼ n ue ;
upn n upn ¼ n up
ð22:8Þ
The elastic sliding displacement vector ue is formulated by the hyperelastic relation to the current contact stress vector f,where ue is calculated by subtracting the plastic sliding displacement up from the total sliding displacement vector u. Note here that the exact additive decomposition of the sliding displacement vector holds, although the deformation gradient tensor defined by the ratio of the current to the initial infinitesimal line-element vectors is obliged to be decomposed multiplicatively into the elastic and the plastic parts as studied in the preceding chapters.
22.2
Sliding Displacement and Contact Traction
645
The large elastoplastic sliding behavior involving nonlinear sliding under rotation of sliding contact surface can be described exactly by the hyperelastic-based plastic sliding formulation (Hashiguchi 2018d, 2020). On the other hand, the hypoelatic-based plastic sliding formulation relating the sliding velocity v to the
corotational contact stress rate f (Hashiguchi et al. 2005; Hashiguchi and Ozaki 2008a; Hashiguchi 2017) is limited to the description of infinitesimal elastic sliding and requires the cumbersome time-integration of the corotational rate of contact stress vector. The contact traction vector f acting on the main body is additively decomposed into the normal traction vector f n and the tangential traction vector f t as follows (see Fig. 22.1): f ¼ f n þ f t ¼ fn n þ ft tf
ð22:9Þ
where
f n ðn fÞn ¼ ðn nÞf ¼ fn n f t ff n ¼ ðIn nÞf ¼ ft tf ðn f t ¼ 0Þ
8 < fn n f : ft tf f ¼ jjf t jj; tf
ft ðn tf ¼ 0; jjtf jj ¼ 1Þ jjf t jj
ð22:10Þ
ð22:11Þ
The minus sign is added for fn to be positive when the compressive stress acts to the main body by the counter body. The contact traction vector f, f n and f t can be calculated from the Cauchy stress r applied in the contact bodies by virtue of the Cauchy’s fundamental theorem in Eq. (5.2) as follows: f ¼ rn f n ¼ ðn r nÞn ¼ ðn nÞr n f t ¼ ðI n nÞr n
fn f
ft Contact surface
Counter body
e3 n
e2 e1
Main body
Fig. 22.1 Contact stress and sliding velocity
un
u
ut
ð22:12Þ
646
22.3
22
Constitutive Equation for Friction: Subloading-Friction Model
Hyperelastic Sliding Behavior
Let the contact stress vector f be given by the hyperelastic relation with the elastic sliding energy function uðue Þ as follows: f¼
@uðue Þ @ue
ð22:13Þ
Then, the work w done during the elastic sliding is uniquely determined by the elastic sliding displacement ue before and after the elastic sliding as follows: Z w¼
ue ue0
Z f due ¼
ue
e @uðue Þ e e u e e du ¼ ½uðu Þue ¼ uðu Þ ¼ uðu0 Þ e 0 e @u u0
The simplest function uðue Þ is given by the quadratic form: e
uðue Þ ¼ ue Eu =2
ð22:14Þ
where the second-order tensor E designates the constant elastic tangent contact T stiffness modulus fulfilling the symmetry E ¼ E . The substitution of Eq. (22.14) into Eq. (22.13) leads to e
1
f ¼ Eu ; ue ¼ E f
ð22:15Þ
Assuming the isotropy on the contact surface, i.e. the independence of frictional property to a sliding direction on the contact surface and introducing the normalized rectangular coordinate system ðe1 ; e2 ; e3 Þ ¼ ðe1 ; e2 ; nÞ fixed to the contact surface, the elastic contact tangent stiffness modulus tensor E is given as follows: 8 < E ¼ at ðI n nÞ þ an n n ¼ at ðe1 e1 þ e2 e2 Þ þ an n n 1 1 1 1 ð22:16Þ 1 : E ¼ ðI n nÞ þ n n ¼ ðe1 e1 þ e2 e2 Þ þ n n at an at an where an and at are the normal and tangential contact elastic moduli, respectively. Equation (22.15) with Eq. (22.16) leads to 8 e e < f ¼ at ut þ an un 1 1 : ue ¼ f t þ f n at an
ð22:17Þ
Further, the contact stress and the elastic sliding displacement are described in the rectangular coordinate system as follows:
22.3
Hyperelastic Sliding Behavior
647
f ¼ f1 e1 þ f2 e2 fn n ue ¼ ue1 e1 þ ue2 e2 uen n
ð22:18Þ
The substitution of Eqs.(22.16) and (22.18) into Eq. (22.15) leads to
resulting in
noting e1 e2 ¼ e1 n ¼ e2 n ¼ 0, which is expressed in the matrix representation as follows: 8 9 2 > at = < f1 > f2 ¼ 4 0 > ; : > 0 fn
0 at 0
3 8 ue 9 0 > = < 1> 5 0 ue2 ; > > a n : ue ; n
8 e9 2 > 1=at = < u1 > ue2 ¼ 4 0 > ; : e> 0 un
0 1=at 0
38 f 9 > 0 = < 1> 5 0 f2 > > 1=an : fn ; ð22:19Þ
Here, note that one does not need to adopt a corotational rate but one has only to use the time derivative for the contact stress vector f by the fact: The contact stress vector f is calculated from the hyperelastic equation with the substitution of the elastic displacement vector ue which is obtained by subtracting the plastic displacement vector up from the displacement vector u.
22.4
Elastoplastic Sliding Velocity
The plastic sliding velocity is formulated based on the subloading concept and the plastic potential theory in this section.
22.4.1
Sliding Normal-Yield and Subloading Surfaces
Assume the following sliding-yield surface with the isotropic hardening/softening, which describes the sliding-yield condition.
648
22
Constitutive Equation for Friction: Subloading-Friction Model
f ðfÞ ¼ l
ð22:20Þ
l designates the expansion (hardening) or the contraction (softening) of the sliding yield surface and is called the sliding hardening function. The sliding-yield stress function f ðfÞ for the Coulomb friction law is given by f ðfÞ ¼ ft =fn
ð22:21Þ
for which l specifies the coefficient of friction which is the dimensionless variable. The variation of l depending on the state of contact traction and/or the history of sliding process is defined by the friction hardening/softening rule, which will be described in Sect. 22.4.2. Now, incorporate the subloading surface concept: The plastic sliding velocity develops as the contact stress approaches the sliding normal-yield surface, renamed the sliding normal-yield surface. Then, introduce the sliding-subloading surface which is similar to the sliding normal-yield surface and passes through the current contact stress. Further, introduce the sliding normal-yield ratio defined by the ratio of the size of the sliding-subloading surface to that of the sliding normal-yield surface, which plays the role to designate the approaching degree of the contact stress to the sliding normal-yield surface. The sliding-subloading surface is represented by the following equation. f ðfÞ ¼ rl
ð22:22Þ
where rð0 r 1Þ is the sliding normal-yield ratio. The sliding-subloading surface for the Coulomb friction law in Eq. (22.22) with Eq. (22.21) is shown in Fig. 22.2. The time-differentiation of Eq. (22.22) leads to the consistency condition for the sliding subloading surface: @f ðfÞ f ¼rl þ r l @f
ð22:23Þ
Here, it is required to formulate the evolution rules of the sliding hardening function l and the sliding normal-yield ratio r as will be done in the subsequent sections.
22.4.2
Evolution Rule of Sliding Hardening Function
The followings might be stated from the results of experiments. (i) The sliding hardening function (friction coefficient for the Coulomb friction law) first reaches the maximal value of static-friction and then decreases to the minimum stationary value of kinetic-friction. Physically, this phenomenon might be interpreted to result from separations of the adhesions of surface asperities between contact bodies due to the sliding (cf. Bowden and Tabor
22.4
Elastoplastic Sliding Velocity
649
fn
e3 n 1
r e1
f e2
tf
f (f )
f ( f (f ) )n (n f
ft
nt
Contact surface
( f (f ) )t f
f (f )
f
f (f )
f
)n
( f (f ) )n f
Sliding-subloading surface Sliding- normal yield surface
ft / fn = r
ft / fn = Fig. 22.2 Coulomb-type sliding normal-yield and sliding-subloading surfaces
1958). Note here that a real contact area between tips of asperities is far smaller than apparent contact area between bodies (cf. Bay and Wanheim 1976). Then, let it be assumed that the plastic sliding causes the decrease of the friction coefficient, i.e. the plastic softening. (ii) The sliding hardening function recovers gradually with the elapse of time and the identical behavior as the initial sliding behavior exhibiting the static friction is reproduced if sufficient time elapses after the sliding ceases. Physically, this phenomenon might be interpreted to result from the reconstructions of the adhesions of surface asperities during the elapsed time under a quite high contact pressure between edges of surface asperities. Then, let it be assumed that the time-elapse causes the increase of the sliding hardening function, i.e. the viscoplastic hardening. Taking account of these facts, let the evolution rule of the sliding hardening/ softening function l be postulated as follows: l ¼ jðl lk Þup þ nðls lÞðlk l ls Þ
ð22:24Þ
where ls and lk are material constants designating the maximum and the minimum values of l for static friction and kinetic friction, respectively. j is the material constant specifying the decrease of the sliding hardening function l plastic displacement kdup k, and n is the material constant specifying the increase of (recovery) of l by the elapsed time increment dt. The first and second terms in Eq. (22.24) are relevant to the destruction and reconstruction, respectively, of the adhesion between
650
22
Constitutive Equation for Friction: Subloading-Friction Model
surface asperities. The variable l is regarded as the hardening variable, i.e. the size of the sliding yield surface. Equation (22.24) is revised from the past formulation (Hashiguchi 2006; Hashiguchi and Ozaki 2008a, etc.). Equation (22.24) is rewritten in the incremental form as follows dl ¼ jðl lk Þjjdu p jj þ nðls lÞdt |fflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflffl} |fflfflfflfflfflfflffl{zfflfflfflfflfflfflffl} Negative
ð22:25Þ
Positive
from which the following characteristics for the evolution of the sliding hardening function are deduced. (a) In a fast sliding process ðdl=dt ! 1; dup =dt ! 1Þ for which the creep part of the second term in the right-hand side is negligible in Eq. (22.25), the sliding hardening function l decreases obeying the following equation. l0 for up ¼ up0 p p ; ð22:26Þ l ¼ ðl0 lk Þ exp½jðu u0 Þ þ lk ¼ lk for up ! 1 dl ¼ jðl lk Þ ¼ dup
jðls lk Þ for l ¼ ls 0 for l ¼ lk
ð22:27Þ
(b) In the quasi-static sliding process ðdl=dt ffi 0; dup =dt ffi 0Þ for which the terms other than the creep part of the second term in the right-hand side are negligible in Eq. (22.25), the sliding hardening function l increases obeying the following exponential equation. l0 for t ¼ t0 ð22:28Þ l ¼ ls ðls l0 Þ exp½nðtt0 Þ ¼ ls for t ! 1 dl ¼ nðls lÞ ¼ dt
nðls lk Þ for l ¼ lk 0 for l ¼ ls
ð22:29Þ
denoting the initial values of t and l as t0 ad l0 , respectively. The variation of the sliding hardening function based on Eqs. (22.26) and (22.28) is shown in Fig. 22.3.
Fig. 22.3 Variation of sliding hardening function
22.4
Elastoplastic Sliding Velocity
22.4.3
651
Evolution Rule of Sliding Normal-Yield Ratio
On the basis of the above-mentioned fundamental postulate of elastoplastic sliding, the rate of the normal sliding-yield ratio r must satisfy the following conditions (see Fig. 22.4): 8 ! þ 1 for r ¼ 0: quasi-elastic sliding state > > < [ 0 for 0\r\1: sliding sub-yield state r for u p 6¼ 0; ¼ 0 for r ¼ 1: sliding normal-yield state > > : ð\0 for r [ 1: sliding over normal-yield stateÞ
r
8 0 p
u = 0 : r 0
p
u 0 : r > 0 p
u 0 : r 0
0
ut
Fig. 22.4 Plastic sliding velocity based on the subloading surface concept
652
22
Constitutive Equation for Friction: Subloading-Friction Model
U (r )
e
p u 0, u 0 p
u 0
0
1
p u 0
r
Fig. 22.5 Function UðrÞ for the evolution rule of r
8 ! þ 1 for r ¼ 0 (quasi-elastic sliding stateÞ > > < [ 0 for 0\r \1 ðsliding sub-yield state) UðrÞ : ¼ 0 for r ¼ 1 ðsliding normal-yield stateÞ > > : ð\0 for r [ 1: sliding over normal-yield stateÞ
ð22:34Þ
which can be given analogously to Eq. (9.12) by the equation p UðrÞ ¼ ~ u cotð rÞ 2
ð22:35Þ
where ~ u is the material constant designating the increment of the sliding normal-yield ratio r, i.e. dr for a certain plastic sliding increment jjdup jj. The contact stress versus the sliding displacement curve becomes the smoother, i.e. gentle slope for the smaller value of ~ u, while the abrupt elastic–plastic transition is described for ~u ! 1. Equation (22.32) with Eq. (22.35) can be analytically integrated in the case of a monotonic sliding process as 8 n p h p io 2 p > 1 p > ~ cos r cos u Þ exp r ¼ u ðu > 0 0 > p 2 2 < ; p > > 2 1 cos 2 r0 p > p > ln : u u0 ¼ p ~ u cos p2 r
ð22:36Þ
R where up is the accumulation of the plastic sliding displacement up ¼ jj u p jjdt in the monotic loading process, and r0 and up0 are the initial values of r and up , respectively. The analytical integration would be beneficial in numerical calculations (Hashiguchi 2013). The general trend of the effect of ~u on UðrÞ is shown schematically in Figs. 22.6 and 22.7. The contact stress is automatically attracted to the sliding normal-yield surface in the plastic sliding process and it is pulled back to that surface even when it goes over the surface in numerical calculation because of r \0 for r [ 1 from Eq. (22.32) with Eq. (22.34) as seen in Fig. 22.8.
22.4
Elastoplastic Sliding Velocity
653
U (r )
u% decreases
0
r
1
Fig. 22.6 Influence of material parameter ~ u on the function UðrÞ
u%
ft / f n
u% decreases
ut
0
Fig. 22.7 Influence of the material parameter ~ u on contact stress versus sliding displacement curve
ft / f n
r = U (r ) || u p || r0
0
U (r ) 0 for r 1 U (r ) = 0 for r = 1
U (r ) 0 for r 1
ut
Fig. 22.8 Contact stress controlling function in subloading-friction model: contact stress is automatically attracted to sliding normal-yield surface in plastic-sliding process
654
22
22.4.4
Constitutive Equation for Friction: Subloading-Friction Model
Elastoplastic Sliding Velocity
The partial derivative of the sliding-yield stress function is given by @f ðfÞ @f ðfÞ @f t @f ðfÞ @f n @f ðfÞ @f ðfÞ þ ¼ ðIn nÞ þ nn ¼ @f @f t @f @f n @f @f t @f n
ð22:37Þ
noting 9 @f n @½ðn nÞf > > ¼ ¼nn = @f @f > @f t @½ðIn nÞf > ¼ ¼ ðIn nÞI ¼ In n ; @f @f
ð22:38Þ
and it follows from Eq. (22.11) that @fn @ðf nÞ ¼ ¼ nI ¼ n @f @f @ft @jjf t jj @jjf t jj @f t f t @½ðIn nÞf ft ¼ ¼ ðIn nÞ tf ¼ ¼ @f jjf t jj @f @f t @f jjf t jj @f
9 > > = > > ;
ð22:39Þ
The substitution of Eqs. (22.24) and (22.32) into Eq. (22.23) leads to p p @f ðfÞ f ¼r½kðl lk Þu þ nðls lÞ þ UðrÞu l @f
ð22:40Þ
Now, assume that the direction of plastic sliding velocity is tangential to the contact plane and outward-normal to the curve generated by the intersection of the sliding subloading surface and the constant normal traction plane f n ¼ const:, leading to the tangential associated flow rule (see Fig. 22.2): u p ¼ k nt ð k 0Þ ðu p ¼ k; n u p ¼ 0Þ
ð22:41Þ
where
@f ðfÞ @f t t
ðnt ¼ 1; n nt ¼ 0Þ
ð22:42Þ
@f ðfÞ @f ðfÞ @f ðfÞ @f ðfÞ n ¼ ðI n nÞ @f t @f @f @f
ð22:43Þ
nt
@f ðfÞ @f
with
22.4
Elastoplastic Sliding Velocity
655
where k and nt are the magnitude and the direction, respectively, of the plastic sliding velocity. The substitution of Eq. (22.41) into Eq. (22.40) reads: @f ðfÞ f ¼ r½kðl lk Þ k þ nðls lÞ þ UðrÞ k l @f
ð22:44Þ
i.e. @f ðfÞ p c f ¼km þm @f
ð22:45Þ
where mp kðl lk Þr þ UðrÞl;
mc nðls lÞrð 0Þ
ð22:46Þ
are relevant to the plastic and the creep sliding velocity, respectively. It is obtained from Eqs. (22.41) and (22.45) that @f ðfÞ @f ðfÞ c c f m f m p @f @f k¼ ;u ¼ nt mp mp
ð22:47Þ
Substituting the rate forms of Eqs. (22.15) and (22.47) into the rate form of Eq. (22.4), the sliding velocity is given by
@f ðfÞ c f m f þ @f p nt m
1
u ¼ E
It follows from Eq. (22.48) that @f ðfÞ c f m @f ðfÞ @f ðfÞ @f ðfÞ @f Eu ¼ f þ Ent mp @f @f @f leading to @f ðfÞ @f ðfÞ c c f m f m @f ðfÞ @f ðfÞ p @f c @f Eu ¼ m m þ Ent p p m m @f @f
ð22:48Þ
656
22
Constitutive Equation for Friction: Subloading-Friction Model
i.e. @f ðfÞ Eu @f
@f ðfÞ ¼ m k m þ k Ent @f p
c
@f ðfÞ p c ¼ m þ Ent k m @f
from which the plastic multiplier in terms of the sliding velocity, denoted by the
symbol K, is given as follows: @f ðfÞ c E u m @f K¼ ; @f ðfÞ mp þ Ent @f
@f ðfÞ c E u m p @f u ¼ nt @f ðfÞ mp þ Ent @f
ð22:49Þ
The inverse relation of Eq. (22.48) is given by substituting the rate form of Eq. (22.4) with Eq. (22.49) into the rate form of Eq. (22.15) as follows: 0
1 @f ðfÞ Ent E C B mc @f Cu þ Ent f¼B @E A @f ðfÞ @f ðfÞ p p Ent m þ Ent m þ @f @f
ð22:50Þ
The contact stress f can be calculated by the time-integration of Eq. (22.50). However, it can be calculated by the hyperelastic equation with the substitution of the elastic displacement ue which is obtained by subtracting the plastic displacement vector up from the displacement vector u. The loading criterion for the plastic sliding rate is given by 8 < u p ¼ 6 0 :p u ¼0
for K [ 0 or
@f ðfÞ c E u m @f
ð22:51Þ
for other
The physical background of this loading criterion is described in the next section. The contact stress function for the isotropic sliding-yield surface is described as f ðfÞ ¼ f ðft ; fn Þ
ð22:52Þ
for which the following partial derivatives hold, noting Eq. (22.39). @f ðfÞ @f ðft ; fn Þ @f ðft ; fn Þ @ft @f ðft ; fn Þ @fn @f ðft ; fn Þ @f ðft ; fn Þ þ ¼ ¼ ¼ tf n @f @f @f @f @ft @fn @ft @fn ð22:53Þ
22.4
Elastoplastic Sliding Velocity
657
It follows by substituting Eq. (22.53) into Eq. (22.42) that nt ¼ tf
ð22:54Þ
Therefore, the direction of the plastic sliding velocity coincides with the direction of the tangential traction. As a particular form of f ðfÞ in Eq. (22.52), let the following Coulomb sliding-yield function be adopted. f ðfÞ ¼
ft fn
ð22:55Þ
It follows for Eqs. (22.53), (22.54) and (22.55) that @f ðfÞ 1 @f ðfÞ ft ¼ ; ¼ 2 fn @ft fn @fn
ð22:56Þ
@f ðfÞ 1 ft tf þ n ¼ fn @f fn
ð22:57Þ
Ent ¼ at tf
ð22:58Þ
Then, we have
E
@f ðfÞ 1 ft 1 ft ¼ ½at ðIn nÞ þ an n n tf þ n ¼ at tf þ an n ð22:59Þ @f fn fn fn fn
@f ðfÞ 1 ft at t f þ n at t f ¼ Etf ¼ ð22:60Þ fn fn @f fn
@f ðfÞ 1 ft a t t f þ an n u Eu ¼ fn @f fn
ð22:61Þ
Substituting Eqs. (22.16), (22.57)–(22.61) into Eqs. (22.48) and (22.50) reduces to
1 ft c t þ n f m f 1 1 fn fn u¼ ðIn nÞ þ n n f þ tf mp at an
3 * 1 ða t þ a ft nÞ u mc + t f n 6 7 fn fn f ¼ ½at ðIn nÞ þ an n n6 u a tf 7 t a 4 5 t mp þ fn
ð22:62Þ
2
ð22:63Þ
658
22
Constitutive Equation for Friction: Subloading-Friction Model
or
ep
f¼E uþ
mc
at at t f fn
1 ft a t t f þ an n at t f fn fn ep E at ðIn nÞ þ an n n a t mp þ fn mp þ
ð22:64Þ
ð22:65Þ
In the coordinate system with the base ðe1 ; e2 ; nÞ, noting In n ¼ ðe1 e1 þ e2 e2 þ n nÞn n ¼ e1 e1 þ e2 e2 tf ¼ tf1 e1 þ tf2 e2
tf1 ¼
f1 f1 ffi; ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffi 2 jjf t jj f1 þ f22
tf2 ¼
ð22:66Þ !
f2 f2 ffi ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffi 2 jjf t jj f1 þ f22 ð22:67Þ
E ¼ at ðe1 e1 þ e2 e2 Þ þ an n n
ð22:68Þ
Equations (22.57)–(22.61) are expressed as
@f ðfÞ 1 ft tf1 e1 þ tf2 e2 þ n ¼ fn @f fn
ð22:69Þ
Etf ¼ at ðtf1 e1 þ tf2 e2 Þ E
ð22:70Þ
@f ðfÞ 1 ft at ðtf1 e1 þ tf2 e2 Þ þ an n ¼ fn @f fn
E u ¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n u
@f ðfÞ 1 ft at ðtf1 e1 þ tf2 e2 Þ þ an n u Eu ¼ fn @f fn noting uT v ¼ u Tv for T ¼ TT . Substituting Eqs. (22.66) and (22.67) into Eq. (22.63) reduces to
ð22:71Þ ð22:72Þ ð22:73Þ
22.4
Elastoplastic Sliding Velocity
659
f ¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n
* 1 a ðt e þ t e Þ þ a ft n u mc + t f1 1 f2 2 n fn f ðtf1 e1 þ tf2 e2 Þ ½ u at n at p m þ fn
¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n u
* 1 a ðt e þ t e Þ þ a ft n u mc + t f1 1 f2 2 n fn f at n ðtf1 e1 þ tf2 e2 Þ a t mp þ fn
ð22:74Þ
¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n u
ft 1 a ðt e þ t e Þ þ a n u t f1 1 f2 2 n fn fn Sat ðtf1 e1 þ tf2 e2 Þ at p m þ fn mc þ Sat ðtf1 e1 þ tf2 e2 Þ at mp þ fn with the loading criterion S¼
8 < :
1 for K [ 0 or
1 ft at ðtf1 e1 þ tf2 e2 Þ þ an n u mc [ 0 fn fn
ð22:75Þ
0 for other
as will be described in Sect. 22.5. Unrealistically high tangential contact stress exceeding the maximum shear stress of the base materials is calculated in case of a high normal contact stress by the above-mentioned subloading-friction model adopting the Coulomb friction condition. This problem can be avoided by inputting a limit normal stress as a material constant into the constitutive equation of friction when the normal contact stress exceeds the limit normal stress. The friction constitutive equation taking account of this fact will be given in Sect. 22.9.
22.5
Loading Criterion
The loading criterion is required for the plastic sliding velocity. It will be formulated in the following, which is similar to that for the plastic strain rate described in Sect. 8.5.
660
22
Constitutive Equation for Friction: Subloading-Friction Model
Now, note the following facts: 1. In the loading (plastic sliding) process, it is required noting Eqs. (22.47) and (22.49) that 9 @f ðfÞ c > > f m > > > > k ¼ @f [ 0 = p m for u p ¼0 ð22:76Þ @f ðfÞ > c > E u m > > > K ¼ @f @f ðfÞ [0> ; mp þ @f Ent
2. In the unloading (elastic sliding) process u p ¼ 0, it holds that @f ðfÞ @f
f \0
ð22:77Þ
and thus @f ðfÞ k ¼ @f
f mc
[0
mp
for mp \0
ð22:78Þ
noting mc 0 (Eq. (22.46)). On the other hand, substituting u ¼ u e leading to
½@f ðfÞ=@f E u ¼ ½@f ðfÞ=@f E u e ¼ ½@f ðfÞ=@f f into Eq. (22.49), it holds
for K that @f ðfÞ c f m @f K¼ \0 @f ðfÞ Ent mp þ @f
ð22:79Þ
in this process. Here, note that the elastic modulus tensor E is the positive definite tensor fulfilling ½@f ðfÞ=@f E½@f ðfÞ=@f [ 0. Here, we assume that nt is not far different from @f ðfÞ=@f, leading to mp þ
@f ðfÞ @f
Ent [ 0
ð22:80Þ
Then, the following inequalities hold from Eqs. (22.78) and (22.79) in the plastic
unloading process (u p ¼ 0; u e 6¼ 0).
k 0 and K 0 when m [ 0 p
9 > > > =
k ! 1 or in determinate and K 0 when mp ¼ 0 > > > ; k 0 and K 0 when mp \0
ð22:81Þ
22.5
Loading Criterion
661
Therefore, the sign of k at the moment of unloading from the state mp 0 is not
necessarily negative. On the other hand, K is negative in the unloading process. Consequently, the distinction between a loading and an unloading processes cannot
be judged by the sign of k but can be done by the sign of K . Inconclusion the loading criterion is given in general as follows: 8
=
f n ¼ const:; tf ¼ const:;
f n ¼ 0; f ¼ ft ¼ ft tf
un ¼ u
e n
¼ u
p n
¼ 0;
u
e
¼
u et tf ;
u ¼ ut tf
> ;
ð22:84Þ
662
22
22.7.1
Constitutive Equation for Friction: Subloading-Friction Model
Relation of Tangential Contact Stress Rate and Sliding Velocity
Equations (22.62) and (22.63) for the Coulomb sliding yield condition are reduced to the following relations in the one-dimensional sliding under the conditions in
Eq. (22.84), denoting ft1 ! f t ; Eq. (22.46).
ft ¼at ut at
vt1 ! vt ð f t2 ¼ f n ¼ 0; ut2 ¼ un ¼ 0Þ and noting
ft ft mc nðls lÞr 1 fn 1 fn f ut ¼ ft þ ¼ þ t mp at at kðl lk Þr þ UðrÞl
ð22:85Þ
at *at v mc + * + v nðls lÞr fn t fn t at ¼at ut at at mp þ kðl lk Þr þ UðrÞl þ fn fn
ð22:86Þ
For the particular case that the sliding velocity is high, so that the creep term can be ignored, i.e. mc ¼ 0, Eqs. (22.85) and (22.86) are reduced to
ut ¼
1 1 þ at fn m p
ft ¼at u t
a2t
f
t
¼
ut p m fn þ at
1 1 þ at fn ½kðl lk Þr þ UðrÞl
1
¼at 1 1þ
1 ¼ at 1 1 þ ðfn =at Þmp !
fn ½kðl lk Þr þ UðrÞl at
ft
ð22:87Þ
ut ð22:88Þ
ut
The relation between the tangential components of the contact stress vector and the displacement vector is schematically shown in Fig. 22.9 for Eq. (22.88) concerning with a high sliding velocity process in which the creep hardening of the second term in Eq. (22.86) is negligible. The relation predicted by the conventional friction model with the sliding-yield surface enclosing an elastic domain is also shown as bold curves 0 y k. In the subloading-friction model, the softening term jðl lk Þrð 0Þ increases monotonically from the negative value to zero and inversely the sliding normal-yield term Ulð 0Þ decreases monotonically from the infinite value to zero in the denominator of the plastic sliding velocity in second term in the bracket in Eq. (22.88). In the initial stage of sliding, the plastic modulus
22.7
Fundamental Mechanical Behavior of Subloading-Friction Model
663
Initial normal sliding-yield surface at static friction
ft / fn = s
t ft
ft
y 1 p
Current subloading-sliding surface
y
ft / fn = p
Final normal sliding-yield and subloading-sliding surfaces at kinetic friction state
f k
ut
0
k 0
ft / fn = k
o f n = const.
fn
Elastic Conventional friction model Elastoplastic
Subloading-friction model
mp
U ( 0)
ut
r
k ( k ) r ( 0)
o
p
k
0
1
ft
f t = t (1
1 ) u 1 ( fn t )m p t
244
k3 m p 14 )r U k (4 { 0
0
mp
0
0
k ( k ) r
U
0
s
(Softening)
k
Fig. 22.9 Prediction of linear sliding behavior from the static to the kinetic friction by the conventional friction and the subloading friction models at a high sliding rate without the creep hardening
is positive, i.e. mp [ 0 so that the tangential contact traction increases but thereafter these terms cancel mutually leading to mp ¼ 0 at which the tangential contact traction reaches the peak, i.e. the static friction point p. Thereafter, the softening term increases gradually from negative to zero but the sliding normal-yield term decreases rapidly resulting in mp \0 so that the tangential contact traction decreases to the kinetic friction point k.
664
22.7.2
22
Constitutive Equation for Friction: Subloading-Friction Model
Numerical Experiments and Comparisons with Test Data
Numerical experiments and comparisons with test data for the subloading-friction model are shown below for Eqs. (22.87) and (22.88) with Eq. (22.35). u, an , at and the initial value l0 of The seven material constants of ls , lk , j, n, ~ the friction coefficient are included in the present model. Material parameters are selected as follows: l0 ¼ ls ¼ 0:4; lk ¼ 0:2 j ¼50 mm1 ; n¼ 0:025 s1 ~ u ¼ 1000 mm1 an ¼ at ¼ 1000 kN/mm3 The influence of the sliding velocity on the relation of the contact stress ratio ft =fn versus the tangential sliding displacement ut are shown in Fig. 22.10. Smooth transitions from the static friction to the kinetic friction and the decreases of the friction coefficient are shown. Faster decrease of friction coefficient is shown for higher sliding velocity. The recovery of the static friction coefficient from the kinetic friction with the elapsed time t after the stop of sliding is shown in Fig. 22.11. In the calculation, the constant sliding velocity vt ¼ 0:1 mm/s is given in the first stage reaching the kinetic friction and then the tangential contact stress is unloaded to zero. After the cessation of sliding for several elapsed times, the same sliding velocity in the first stage is given again. The recovery is larger for a longer stationary time.
vt = 0.0001mm/s
0.4
vt = 0.001mm/s vt = 0.01mm/s
0.2
vt = 0.1mm/s
ft /fn
0.3
0.1 0.0 0
0.02
Fig. 22.10 Influence of sliding velocity
0.04 0.06 ut [mm]
0.08
0.1
22.7
Fundamental Mechanical Behavior of Subloading-Friction Model
665
0.4 100 s 50 s 20 s 10 s
ft /fn
0.3
0.2 0s
0.1 Stationary contact
0.0
0
0.05
0.1 0.15 ut [mm]
0.2
Fig. 22.11 Influence of stationary time
The influence of the material constant ~ u in Eq. (22.35) for the evolution rule of the sliding normal-ratio r in Eq. (22.32) on the variation of the contact stress ratio ft =fn is shown in Fig. 22.12, where the material constants and the initial value are the same as those in Fig. 22.10. The calculated results for low and high sliding velocities are shown in Fig. 22.12a, b, respectively. As shown in this figure, smoother transition from the elastic to the plastic state is shown for smaller values of ~u. The accumulations of sliding displacement under the cyclic loading of tangential contact traction of 80%, i.e., ft ¼ 0 0:8ls fn for the two levels of sliding velocities are shown in Fig. 22.13, where the material parameters are chosen same as in Figs. 22.10, 22.11 and 22.12. The sliding displacement increases as the velocity is larger since the recovery of friction requires time. In particular, the occurrence of infinite sliding under a high rate of sliding is shown in Fig. 22.13b. The accumulation cannot be predicted at all by the conventional friction model capable of predicting only elastic sliding for the contact stress lower than the sliding-yield condition. The serious accidents of automobiles, rail ways, airplanes, etc. and also the defects in atomic power plants occur frequently by the loosening of fastening elements, e.g. bolts and nuts. It should be emphasized that the incorporation of the subloading-friction model is inevitable in the mechanical designs of machine elements for prevention of these accidents. The comparison with test data for the reduction process of friction coefficient from the static- to kinetic-friction is shown in Fig. 22.14. The test curve for sliding between roughly polished steel surfaces (Ferrero and Barrau 1997) under the quite low sliding velocity vt 0:0002 mm/s is simulated well enough by the present model, where the material parameters are selected as follows:
666
22
0.5 0.4
Constitutive Equation for Friction: Subloading-Friction Model
u% 5, 000mm 1 u% 1, 000mm 1
ft /fn
0.3
u% 250mm 1 u% 100mm 1 u% 50mm 1 u% 25mm 1 u% 10mm 1
0.2 0.1 0.0
0
0.04
0.02
0.06 0.08 u t [mm]
0.1
(a ) u t = 0.0001 mm/s
0.5 u% 5, 000mm 1 u% 1, 000mm 1 u% 250mm 1
0.4
ft /fn
0.3 0.2
u% 100mm 1 u% 50mm 1 u% 25mm 1 u% 10mm
0.1 0.0
0
0.02
0.04
0.06 0.08 u t [mm]
0.1
(b) u t = 0.01 mm/s Fig. 22.12 Influences of the material constant in the evolution rule of normal sliding-yield surface on the relation of friction stress ratio versus tangential sliding displacement for two levels of tangential sliding velocity
l0 ¼ ls ¼ 0:58; lk ¼ 0:38 j ¼ 92:11 mm1 ; n ¼ 0:0086 s1 ~ u ¼ 2000 mm1 an ¼ at ¼ 1000 kN=mm3
22.7
Fundamental Mechanical Behavior of Subloading-Friction Model
667
0.4
Conventional friction model ( u )
ft /fn
80 %
0.2
0 0
0.1
u t [mm]
0.2
•
(a) u t 0.0001 mm/s 0.4
Conventional friction model ( u )
ft /fn
80 %
0.2
0 0
0.1
u t [mm]
0.2
•
(b) u t 0.01 mm/s Fig. 22.13 Influence of sliding velocity on accumulation of sliding displacement under cyclic loading
under the condition
fn ¼ 10 MPa; u t ¼ 0:0002 mm=s The comparison with test data for the recovery of friction stress ratio by the stop of sliding on the way of the reduction process from the static- to kinetic-friction is depicted in Fig. 22.15. The test curves for sliding between roughly polished steel surfaces (Ferrero and Barrau 1997) under the infinitesimal sliding velocity
668
22
Constitutive Equation for Friction: Subloading-Friction Model
0.6
ft / f
n
Calculation Experiment
0.4
0.2 0
0.04
ut [mm]
0.08
Fig. 22.14 Comparison with test data (Ferrero and Barrau 1997) for reduction of friction stress ratio under linear sliding
0.6 Calculation Experiment ft / f
n
vt 0.0002 mm/s
0.4
0.2
0
Stationary time
Stationary time
400 s
20 s
0.02
ut [mm]
0.04
Fig. 22.15 Comparison with test data (Ferrero and Barrau 1997) for influence of stationary time on recovery of friction stress ratio
u t 0:0002 mm=s and the stationary time 20 and 400 s are simulated sufficiently well by the present model, where the material parameters are selected same as for the calculation in Fig. 22.14. The quite close simulation of test data for the frtction behavior between metals and sands by the subloading-friction model is shown in Ozaki et al. (2022).
22.8
22.8
Stick–Slip Phenomenon
669
Stick–Slip Phenomenon
When the sliding between solid bodies proceeds in a low velocity, the unstable motion leading to the intermittent vibration phenomenon is induced, which is referred to as the stick–slip motion. The unstable motion influences on the performance of machinery and wear, fatigue, durability and acoustic emission systems. It should be avoided in elements such as gears and bearings to produce the smooth movement. In addition, the stick–slip motion is observed in earthquakes in a continental plate sliding type. In what follows, let the stick–slip motion be simulated by the subloading-friction model described in the preceding sections. The stick–slip analysis delineated in this section was executed pertinently by Ozaki and Hashiguchi (2010). The simplest Coulomb friction condition is used again in this section. First, let the stick–slip motion be examined qualitatively by the simplest example, i.e., the one-dimensional sliding as shown in Fig. 22.16, while the movement of slider is shown in Fig. 22.17a and the variation of tangential contact stress and the sliding displacement of slider with the time is shown in Fig. 22.17b. The slider is connected to the spring which is pulled in a constant velocity. (1) The slider stops until the spring force reaches the value causing the static friction as shown in the processes (1–0) (1 − s). (2) At the moment when the spring force reaches the value causing the static friction, the slider moves until spring force decreases to the value corresponding to the kinetic friction as shown in the process (1 − k). (3) The identical phenomena to the processes (1 − m) (1 − k) is repeated as shown in the processes (2 − m) (3 − k) and so on. We now describe mathematically the stick–slip instability, taking account of the acceleration of slider. Denoting the constant velocity at the spring-end induced by the driver by V R and the displacement of deriver by U, the spring elongation is given as U ut ð¼ ðV vt ÞdtÞ. The equation of motion is given by KðU ut Þ ft S ¼ Mat
ð22:89Þ
where M is the mass of the slider, K is the spring stiffness, S is the nominal contact area, and at and ut are the acceleration and the displacement, respectively, of the slider, which are relative values with respect to the fixed base. The tangential contact stress ft is estimated by Eq. (22.86). Now, we show the numerical simulation of the test result for the stick–slip behavior measured by Baumberger et al. (1994). The structure of test apparatus is as follows: Mass of slider: M ¼ 0:8 kg, Contact area: S ¼ 720 mm2 , Spring constant: K ¼ 58 N/mm, Velocity of deriver of spring: V ¼ 0:001 mm/s.
670
22
Constitutive Equation for Friction: Subloading-Friction Model ft = 0
(1-0)
0
(1-m)
_
ft s ft s
(1-s)
u t(1)
ft k
(1-k)
ft k ft s
(2-m)
ft s
(2-s)
u t(2)
ft k
(3-k)
(3-m)
ft k ft s
ut
vt
U
V
(a) Movement of slider.
ft
(1-s)
ft s
(2-s)
(1-m) ft k
(2-m) (1-k)
(3-m)
(3-k) time
(1-0)
ut
(3-k) (3-m)
u t(2) u t(1)
(1-k) (2-m) (2-s)
(1-0) (1-m) (1-s)
time
(b) Variations of force and movement of slider.
Fig. 22.16 One-dimensional stick–slip phenomenon
The comparison with test data is shown in Fig. 22.17, where the material parameters are chosen as follows: l0 ¼ ls ¼ 0:4; lk ¼ 0:2
22.8
Stick–Slip Phenomenon
u t [mm]
0.20
671
= 0.5mm−1 , = 25s −1 = 0.5mm−1 , = 0 = 0.0mm−1, = 25s−1
0.15
0.10 Baumberger et al. (1994)
0.05
0.00
0
0.05
0.1 0.15 U [mm]
0.2
(a) Relation between displacements of slider and deriver
0.4
ft /fn
0.3 0.2
= 0.5mm−1 , = 25s−1 = 0.5mm−1 , = 0 = 0.0mm−1, = 25s −1
0.1
0.0
0
0.05
0.1
0.15
0.2
u [mm] (b) Relation between contact stress ratio versus sliding displacement.
Fig. 22.17 Simulation of stick–slip phenomenon with comparison to test result after Baumberger et al. (1994), (Ozaki and Hashiguchi 2010)
j ¼50; n ¼ 0:25 s1 ~ u ¼ 1; 000 mm1 an ¼ at ¼ 1; 000 N/mm3 The influences of the mass of slider and the spring constant are shown in Figs. 22.18 and 22.19. The larger mass of slider and the weaker spring constant induce the more intense stick–slip behavior as observed in this figure. Now, examine the influence of driver velocity on the stick–slip motion. Then, the variations of spring elongation under the increase and decrease of driver velocity are shown in Fig. 22.20 for K ¼ 50 N=mm, M ¼ 1 kg, S ¼ 1000 mm2 . The material parameters are chosen to be same as the ones for Fig. 22.17. The
22
U u t [mm]
U u t [mm]
U u t [mm]
U u t [mm]
672
0.3
Constitutive Equation for Friction: Subloading-Friction Model
M 3.0 kg
0.2 0.1 0 0 0.3
200
400
600
800
400
600
800
400
600
800
M 1.5 kg
0.2 0.1 0 0 0.3
0.2
200 M 0.75 kg
0.1 0 0 0.3
200 M 0.5 kg
0.2 0.1 0 0
200
400 800 600 Elapsed time [s]
1000
Fig. 22.18 Influence of slider mass on stick–slip behavior (Ozaki and Hashiguchi 2010)
responses for the cases that the driver velocity increases from V ¼ 0:0005 mm=s and decreases from V ¼ 0:01 mm=s linearly are shown in this figure. It is observed that the stick–slip movement converges as the driver velocity increases and that it commences and amplifies as the velocity decreases. The prediction of earthquake occurrence has been done by the unreliable method by the rate-and-state model (cf. e.g. Dieterich 1978, 1979; Ruina 1980, 1983; Rice and Ruina 1983; Scholz 1998; Rice et al. 2001; Kame et al. 2013) which is quite irrational as described in Sect. 22.1. The reliable prediction of the earthquake would be able to be realized by incorporating the subloading-friction model.
22.9
Friction Condition with Saturation of Tangential Contact Stress
It is recognized from experiments (cf. Bay and Wanheim 1976; Dunkin and Kim 1996; Stupkiewicz and Mroz 1999; Gearing et al. 2001) that the friction-coefficient decreases with the increase of normal contact stress. However, this property cannot
22.9
Friction Condition with Limitation of Tangential Contact Stress
U u t [mm]
U u t [mm]
U u t [mm]
U u t [mm]
0.4
673
K 10 N/mm
0.2 0 0 0.4
200
400
600
8001000
600
8001000
600
8001000
K 20 N/mm
0.2 0 0 0.4
200
400
K 40 N/mm
0.2 0 0 0.4
200
400
K 80 N/mm
0.2 0 0
200
400 800 600 Elapsed time [s]
10
Fig. 22.19 Influence of spring constant on stick–slip behavior (Ozaki and Hashiguchi 2010)
be described by adopting the Coulomb sliding-yield surface in which the tangential contact stress increases linearly with the normal contact stress, resulting in the constant friction coefficient. Actually, the unrealistic result that the tangential contact stress higher than the shear yield stress of the base materials applies at the high normal contact surface and thus the base materials adhere to each other, leading to their peeling-off by the sliding, is calculated by the calculation using the Coulomb sliding-yield surface. Actually, this problem can be seen in the tightening of bolts and nuts and in the friction between the landward plate and the seaward plate during the submerge of the seaward plate which is essential for predicting earthquake occurrence. Then, the concrete formulations of the subloading-friction model to take account of this property have been proposed by Hashiguchi et al. (2005) and Ozaki et al. (2020). However, the former formulation (Hashiguchi et al. 2005) involves the physical irrationality that the tangent friction-coefficient becomes minus in a high normal contact stress region. The latter formulation (Ozaki, 2020) is irrelevant to solid materials as metals but limited to the rubbers. The rigorous subloading-friction model incorporating the yield stress function f ðfÞ by which the tangential contact stress saturates with the increase of the normal
674
22
Constitutive Equation for Friction: Subloading-Friction Model
V [mm/s]
0.01
U u t [mm]
0.00 0 0.10
400
200
600
800 1000
0.05 Unstable Stable 0.00
0
200
400 800 600 Elapsed time [s]
1000
(a) Linear increasing process of deriver velocity
V [mm/s]
0.01
U u t [mm]
0.00 0
0.10
400
200
600
800
0.05 0.00
Unstable Stable 0
200
400 800 600 Elapsed time [s]
1000
(b) Linear decreasing process of deriver velocity
Fig. 22.20 Variation of spring elongation under linear variation of deriver velocity (Ozaki and Hashiguchi 2010)
contact stress will be formulated in the following (Hashiguchi, 2021b; Hashiguchi and Ueno, 2022). The following sliding-yield stress function f ðfÞ ¼ f ðft ; fn Þ in Eq. (22.20), i.e. f ðft ; fn Þ ¼ l for the sliding normal-yield surface is incorporated (Hashiguchi 2021b) ft gn ðfn Þ
ð22:90Þ
i.e: ft ¼ gn l
ð22:91Þ
ðf ðfÞ ¼Þf ðft ; fn Þ ¼ leading to the sliding normal-yield surface: ft =gn ¼ l;
22.9
Friction Condition with Limitation of Tangential Contact Stress
675
Fig. 22.21 Variation of sliding hardening function
where gn ð0 gn 1Þ is the function of fn satisfying the conditions gn
g0n
¼ 0 for fn ¼ 0 ! 1ðft ¼ lÞ for fn ! 1
dgn dfn
(
¼ cn for fn ¼ 0 ! 0 for fn ! 1
ð22:92Þ
ð22:93Þ
leading to ( ¼ cn l for fn ¼ 0 @ft 0 ¼ gn l @fn ! 0 for fn ! 1
ð22:94Þ
cn is the material constant with the inverse dimension of stress. Equation (22.91) represents the sliding normal-yield surface with the fusiform shape which expands from the origin to the positive direction of the normal contact stress fn in the tree-dimensional stress space ðft1 ; ft2 ; fn Þ as shown in Fig. 22.21. The tangential contact stress ft on the sliding normal-yield surface increases with the normal contact stress fn and saturates at ft ¼ l, while it approaches faster to the saturation (maximum) value l for a larger value of cn as shown in Fig. 22.22a which is depicted for l ¼ const. The sliding normal-yield and the sliding
676
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.22 Sliding normal-yield and subloading surfaces with limitation of tangential contact stress
subloading surfaces satisfying Eqs. (22.92) and (22.93) are shown in Fig. 22.22b. The Coulomb friction-yield condition can be realized approximately by setting l ! 1 (large value) and cn ffi 0 (small value) leading to cn l ffi const: for fn 1. The following differentiations hold for Eq. (22.90). @f ðfÞ 1 ¼ ; @ft gn
@f ðfÞ ft ¼ 2 g0n gn @fn
ð22:95Þ
In what follows, the constitutive equation will be shown for Eq. (22.103). Now, we have @f ðfÞ 1 ¼ 2 ðgn tf þ ft g0n nÞ @f gn noting Eq. (22.53) with Eq. (22.95).
ð22:96Þ
22.9
Friction Condition with Limitation of Tangential Contact Stress
677
Further, we have Ent ¼ Etf ¼ at tf
E
ð22:97Þ
@f ðfÞ 1 1 ¼ ½at ðIn nÞ þ an n n 2 ðgn tf þ ft g0n nÞ ¼ 2 ðat gn tf þ an ft g0n nÞ @f gn gn ð22:98Þ @f ðfÞ @f
Etf ¼
1 at ðat gn tf þ an ft g0n nÞ at tf ¼ 2 gn gn
ð22:99Þ
@f ðfÞ 1 E u ¼ ðat gn tf þ an ft g0n nÞ u @f g2n
ð22:100Þ
noting uT v ¼ u Tv for T ¼ TT . The substitutions of Eqs. (22.96)–(22.100) into Eqs. (22.48) and (22.50) with Eq. (22.54) lead to 1 ðat gn tf þ an ft g0n nÞ f mc 2 1 1 g u¼ ðIn nÞ þ n n f þ n tf ð22:101Þ mp at an
* 1 ða g t þ a f g0 nÞ u mc + # t n f n t n g2 tf f ¼ ½at ðIn nÞ þ an n n u at n mp þ at =gn "
ð22:102Þ or
f ¼E
ep
ep
u þ at
mp
at t f
E at ðIn nÞ þ an n n
mc tf þ at =gn
ð22:103Þ
1 0 ðat gn tf þ an ft gn nÞ 2 gn mp þ at =gn
ð22:104Þ
The examples of the explicit function satisfying Eqs. (22.92) and (22.93) are given by Exponential function: gn ¼ 1 expðcn fn Þ;
g0n ¼ cn expðcn fn Þ
ð22:105Þ
678
22
Constitutive Equation for Friction: Subloading-Friction Model
Tangent-hyperbolic function: gn ¼ tanhðcn fn Þ;
g0n ¼ cn sech2 ðcn fn Þ
ð22:106Þ
In what follows, based on Eqs. (22.96)–(22.102), the explicit equations for the function gn in Eq. (22.105) are shown in the coordinate system ðe1 ; e2 ; nÞ. @f ðfÞ 1 ¼ 2 ½gn ðtf1 e1 þ tf2 e2 Þ þ cn ft expðcn fn Þn @f gn E
@f ðfÞ 1 ¼ 2 ½at gn ðtf1 e1 þ tf2 e2 Þ þ an cn ft expðcn fn Þn @f gn
E u ¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n u @f ðfÞ 1 E u ¼ ½a g ðt e þ t e Þ þ a c f expðc f Þn u t f 1 f 2 n n t n n n 1 2 @f g2n
ð22:107Þ ð22:108Þ
ð22:109Þ ð22:110Þ
f ¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n " * 1 ½a ðt e þ t e Þ þ a c ðf =g Þ expðc f Þn u # mc + t f1 1 f2 2 n n t n n n gn ðtf1 e1 þ tf2 e2 Þ u at mp þ at =gn
¼ ½at ðe1 e1 þ e2 e2 Þ þ an n n u * 1 ½a ðt e þ t e Þ þ a c ðf =g Þ expðc f Þn u mc + t f1 1 f1 2 n n t n n n gn at ðtf1 e1 þ tf2 e2 Þ mp þ at =gn
ð22:111Þ The fiction model formulated in this subsection possesses only one more material constant cn in addition to the seven material constants ls , lk , j, n, ~u, an and at involved in the Coulomb friction model in Sect. 22.4.4. The computer program for friction analysis based on the formulations described so far is given in the Appendix L(d). The numerical experiments and the comparison with the test data (Hashiguchi and Ueno, 2020) will be shown below.
22.9
Friction Condition with Limitation of Tangential Contact Stress
679
(a) Numerical experiments The principal mechanical responses of the present subloading-friction model will be examined by performing the numerical experiments in this section. The material parameters are chosen with the two levels of the static sliding hardening functions ls as follows: an ¼ 5; 000 MPa mm3 ; at ¼ 5; 000 MPa mm3 ; ls ¼ 10; 000 MPa and 7; 500 MPa; lk ¼ 7; 000 MPa
j ¼ 2:14; n ¼ 0:011 s1 ; u ¼ 15 mm1 ; cn ¼ 0:0002 MPa1
and the tangential sliding velocity is chosen as u t ¼ 1 mm/60s. Firstly, the tangential contact stress paths at the five levels of constant normal contact stress are shown in Fig. 22.23 in which the sliding normal-yield surface for the static and the kinetic isotropic hardening functions are depicted by the red and the black curve, respectively. Needless to say, the maximum contact tangential stress is larger for the larger static sliding hardening function. Next, the variations of tangential contact stress at the five levels of normal contact stress in the monotonic sliding followed by the reverse sliding for the different static sliding- hardening functions are shown in Fig. 22.24. The variations of tangential contact stress with the five levels of stationary sliding time just after the unloading for the different static sliding-hardening functions in the monotonic and the reciprocal sliding are shown in Figs. 22.25 and 22.26, respectively. The more remarkable recoveries of the tangential contact stress by the longer stationary sliding time are shown in these figures. The variations of the tangential contact stress in the pulsating sliding with the five levels of stationary sliding just after the unloading to the zero tangential contact
Fig. 22.23 Tangential contact stress paths at five levels of normal contact stress for different static sliding-hardening functions. Red short bar in each stress path indicates the maximum contact tangential stress, and the black one indicates the contact tangential stress at infinite sliding displacement
680
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.24 Variations of tangential contact stress at five levels of normal contact stress in monotonic sliding followed by the reverse sliding for different static sliding-hardening functions
Fig. 22.25 Variations of tangential contact stress with five levels of stationary sliding time just after unloading for different static sliding-hardening functions
stress is shown in Fig. 22.27. The recovery of the tangential contact stress with the increase of sliding time is shown, while it decreases gradually with the increase of number of cycle. (b) Simulation of test results The simulation of the test data for the linear sliding behavior for the boric acid lubricated A16111-T4 and tool steel interface after Gearing et al. (2001) is shown in Fig. 22.28 to verify the applicability of the present model to the description of real friction phenomenon. The linear sliding behavior at the constant normal contact stress is the most basic behavior among various sliding behaviors but strangely the test data is quite few as the authors could find only this test data. Fortunately, this test data is suitable for the verification of the present model, since the rather high
22.9
Friction Condition with Limitation of Tangential Contact Stress
681
Fig. 22.26 Variations of tangential contact stress during reciprocal sliding with five levels of stationary sliding time just after unloading for different static sliding-hardening functions
Fig. 22.27 Variations of tangential contact stress in pulsating contact tangential stress with five levels of stationary sliding time just after unloading for different static sliding-hardening functions
normal contact stress up to 600 MPa is applied in this data. The relations of the contact tangential stress vs. the sliding displacement for the five levels of the constant normal contact stresses are represented, where the values of the material parameters are chosen as follows: an ¼ 500 MPa mm3 ; at ¼ 500 MPa mm3 ; ls ¼ 500 MPa; lk ¼ 63 MPa;
j ¼ 1:56; n ¼ 0:22 s1 ; u ¼ 15 mm1 ; cn ¼ 0:0002 MPa1
682
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.28 Comparison with test data of boric acid lubricated A16111-T4/tool steel interface after Gearing et al. (2001) for tangential contact stress vs. sliding displacement relations at six levels of normal contact stress, where calculated results are shown by solid lines
while the tangential sliding velocity ut is 1 mm s-1 in the test data. The variations of the tangential contact stress are closely simulated by the present friction model as shown in Fig. 22.28. The comparison for the relations of the tangential contact stress vs. the normal contact stress for the four levels of the sliding displacement is shown in Fig. 22.29, which is depicted from the test and the calculated values. The fact that the ratio of the tangential contact stress to the normal contact stress is not constant but decreases gradually with the increase of the normal contact stress is simulated closely. The incorporation of the present model is required to describe the friction behavior in this test data accurately. A test datum specifying an explicit sliding displacement less than 1mm is not shown in the figure of the paper: Gearing et al. (2001). Then, only the calculated result for the displacement 0.5mm is shown by the dashed curve for the reference in Fig. 22.29. The quite close simulation of the test result (Gearing et al. 2001) is attained by the present model. Then, the applicability of the present friction model to the prediction of real friction behavior between metals would be verified. Besides, the simulation of this test data was performed also by Gearing et al. (2001) themselves.
22.9
Friction Condition with Limitation of Tangential Contact Stress
683
Fig. 22.29 Comparison with test data of boric acid lubricated A16111-T4/tool steel interface after Gearing et al. (2001) for tangential contact stress vs. normal contact stress relations in four levels of sliding displacement, where calculated results are shown by solid lines. The dashed curve shows the calculated result for the sliding displacement 0.5 mm
However, the friction model proposed by Anand (1993) used to the simulation is physically impertinent belonging to the creep model which is approximately applicable to the sliding behavior at high rate but inapplicable to the sliding behavior at the moderate rate including the quasi-static sliding behavior as delineated in Hashiguchi (2017, 2020), while the sliding velocity in their test data
(Gearing et al. 2001) is rather high as ut is 1 mm/s . On the other hand, the subloading-friction model is concerned with the sliding behaviour in the general sliding velocity.
22.10
Subloading-Overstress Friction Model
In the subloading friction model, the sliding velocity is postulated to consist of elastic and plastic sliding velocities. It leads to negative rate-sensitivity (i.e., a decrease in friction resistance with increasing sliding velocity) because the adhesion of surface asperities is lost quickly at high sliding velocities and the recovery of adhesion requires time. In contrast, the positive rate sensitivity (i.e. an increase in the friction resistance with increasing sliding velocity) is observed in the sliding
684
22
Constitutive Equation for Friction: Subloading-Friction Model
between lubricated surfaces or between soft solids, e.g. indium, Teflon and various polymers, which is often called the fluid friction or the wet friction. Its theoretical prediction is of importance for mechanical designs of lubricated machinery elements such as gears and bearings in which low friction resistances are desirable and of wheel tires travelling on wet roads in which high friction resistances are required for high breaking performances, and so forth. In what follows, the subloading-overstress friction model to describe not only negative but also positive rate sensitivities (Hashiguchi et al. 2016) will be given by introducing the notion of the overstress which has been adopted in the description of viscoplastic deformation behaviour in Chap. 14.
22.10.1
Subloading-Overstress (Viscoelastic) Friction Model
The constitutive equation described in the former sections will be extended to the subloading-overstress model in this section. First, we introduce the viscoplastic sliding displacement vector uvp instead of the plastic sliding displacement vector up in Eq. (22.4), i.e. u ¼ ue þ uvp :
ð22:112Þ
The viscoplastic strain rate is given by
u
vp
¼ Cnt
ð22:113Þ
leading to 1
u ¼E
f þ Cnt ;
ð22:114Þ
and
f ¼ E u CEnt ;
ð22:115Þ
The explicit function of C is given by the power function 1 hr rs in gv ^r r
ð22:116Þ
1 henðrrs Þ 1i 1 sinhðnhr rs iÞ ;C¼ ^r r ^r r gv gv
ð22:117Þ
C¼ or the exponential function C¼
22.10
Subloading-Overstress Friction Model
685
where gv is the material constant specifying the time required for the occurrence of a unit viscoplastic strain rate and ^r ð 1Þ is the material constant specifying the maximum value of the sliding-normal-yield ratio which is realized in the impact
sliding process ðjj u vp jj ! 1Þ. rs ð0 rs 1Þ is the subloading sliding-yield ratio calculated using Eqs. (22.32) and (22.33) by replacing the plastic sliding dis
placement rate u p with the viscoplastic sliding rate u vp , i.e.,
rs ¼ Uðrs Þjj u
vp
f ðfÞ l
rs ¼
jj
for
u
for
u
vp
vp
6¼ 0;
ð22:118Þ
¼0;
ð22:119Þ
with p Uðrs Þ ¼ ~ u cotð rs Þ: 2
ð22:120Þ
Thus, the viscoplastic sliding is induced by overstress f ðfÞrs l from the subloading friction surface: f ðfÞ¼rs l; i:e: r¼rs ;
ð22:121Þ
so that a smooth elastic–viscoplastic transition is described. Simulations reproducing test data are difficult to perform using Eq. (22.113) with a viscoplastic term in the power form but, as will be described later, simulations with high accuracy are possible using the equations with the exponential form
¼
vp
u
1 henðrrs Þ 1i nt ; ^r r gv
1
u ¼E
f þ
f ¼E u
ð22:122Þ
1 henðrrs Þ 1i nt ; ^r r gv
ð22:123Þ
1 henðrrs Þ 1i Ent ; ^r r gv
ð22:124Þ
which will be used hereinafter. In the Coulomb friction condition, the evolution of the friction coefficient l is
given by replacing the plastic sliding displacement rate u p to the viscoplastic
sliding displacement rate u vp in Eq. (22.24) as follows:
l ¼ jðl lk Þjj u
vp
jj þ nðls lÞðlk l ls Þ
ð22:125Þ
In what follows, let the friction behavior for the Coulomb friction condition be examined.
686
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.30 Response of the generalized subloading-friction model
The various behaviours derived from the subloading-overstress friction model for fluid friction are schematically shown in Fig. 22.30 in which the normalized overstress divided by the normal contact stress fn is written simply by the term “overstress”. The dynamic loading sliding-yield ratio r coincides with the subloading sliding-yield ratio rs , i.e. r ¼ rs in quasi-static sliding and increases above rs with increasing sliding velocity. However, r does not rise above ^r , i.e., r ^r , with the equality r ¼ ^r being realised only for impact sliding. Equations (22.124) and (22.125) are described in incremental forms as df ¼ Edu
1 henðrrs Þ 1i Ent dt ^r r gv
dl¼ jðl lk Þjjdu vp jj þ nðls lÞ dt ðlk l ls Þ
ð22:126Þ ð22:127Þ
We infer the following properties from Eq. (22.127) for the coefficient of friction l in monotonic sliding at constant sliding velocity (see Fig. 22.31). (1) In the quasi-static sliding process ðdl=dt ffi 0; jjduvp jj=dt ffi 0, jjdujj=dt ffi 0; jjdfjj=dt ffi 0Þ in which the terms other than the second terms in the right-hand sides are negligible in Eqs. (22.126) and (22.127), we have l ¼ ls from Eq. (22.127) and r ¼ rs from Eq. (22.126), resulting in ft =fn ¼ rs ls due to Eq. (22.22) with Eq. (22.21). Therefore, the contact stress moves satisfying the subloading surface in quasi-static sliding process, so that the original subloading-friction model is reproduced in that process. In other words, the subloading-overstress friction model is generalized so as to contain the subloading-friction model formulated for the dry friction in the fomer sections. Therefore, the subloading-friction model can be disused by using the subloading overstress friction model.
22.10
Subloading-Overstress Friction Model
687
Fig. 22.31 Influence of sliding velocity at constant sliding velocity process
(2) In the fast sliding process ðdl=dt ! 1; jjduvp jj=dt ! 1; jjdujj=dt ! 1; jjdfjj=dt ! 1Þ for which the creep part of the second term in the right-hand side is negligible in Eq. (22.127), the friction coefficient l decreases with the viscoplastic sliding displacement uvp after reaching a peak. Friction resistance depends on the destruction of the adhesion of surface asperities and the shearing of the viscous fluid lying between contact surfaces. The former is dominant in the case of dry friction whereas the latter is dominant in the case of fluid friction since the surfaces are in direct contact with each other in the case of dry friction but are in indirect contact via the viscous medium in the case of fluid friction. We thus infer the following differences between the responses of dry and fluid friction in the monotonic sliding process at a constant sliding velocity (see Fig. 22.32). (3) The friction resistance in dry friction is obviously larger compared with that in fluid friction, because both the static coefficient of friction ls and the kinetic coefficient of friction lk in the former are larger compared with those in the latter. In addition, the difference between a peak and a bottom contact stress ratios is larger in the dry friction than in the fluid friction. (4) The sliding displacement in the transition from static friction to kinetic friction in the case of dry friction is far smaller than that in the case of fluid friction because the sliding displacement required for destruction of the adhesion of surface asperities is far smaller than that required for shear flow in a viscous fluid. Therefore, the contact stress ratio decreases more gradually in the case of fluid friction than in the case of dry friction.
688
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.32 Comparison of responses in dry and fluid frictions at constant sliding velocities
(5) The friction coefficient l decreases with increasing sliding displacement as described in (1). Then, the dry friction exhibits the negative rate-sensitivity since the contact stress ratio ft =fn is given by the friction coefficient itself. On the other hand, the fluid friction exhibits the positive rate-sensitivity since the contact stress ratio is given by the friction coefficient plus the overstress, while the overstress increases with sliding velocity inducing a higher viscous resistance. The friction resistance is kept constant in the quasi-static sliding process in which the friction coefficient is kept to be the static-friction, and from which the friction resistance decreases and increases in the dry and the fluid friction, respectively, with the increment of sliding velocity. (6) Consequently, the contact stress ratio ft =ðlfn Þ approaches unity (i.e., r ¼ 1: the normal sliding yield state) in the case of dry friction but becomes greater than unity (i.e., r [ 1: over the sliding normal-yield state) in fluid friction. It should be emphasized again that it is not required to use Eqs. (22.48) and/or Eq. (22.50) in the subloading friction model even for calculation of dry friction behaviour. In facts, the dry friction behaviour is generated in the quasi-static sliding process in the subloading overstress friction model. Consequently, we need only to use Eqs. (22.123) and/or (22.124) with Eqs. (22.125) and (22.118)–(22.120) in the subloading overstress friction model for calculations of general sliding behaviour involving both dry and fluid frictions. Additionally, there exists an intermediate friction between dry and fluid friction where contact surfaces are lubricated slightly, leading to a behaviour in which there is no effect of the sliding velocity on the contact stress ratio under constant sliding velocities. The sliding behaviour ranging from dry to fluid friction covering negative to positive rate sensitivities can be described universally by the generalised subloading friction model. The return-mapping projection in the numerical analysis for the rate-independent elastoplastic deformation (e.g., Simo and Hughes 1998;
22.10
Subloading-Overstress Friction Model
689
Hashiguchi 2013b) is regarded to be an application of the overstress concept. Consequently, the subloading overstress friction model is the generalisation of the subloading friction model. In addition, the Stribeck curve (Stribeck 1902), which has been used widely to provide a qualitative relationship between frictional resistance ft and variable gv vt =fn , would also be able to be drawn by changing appropriately the material constants an ; at ; ls ; lk ; j; n; ~u; gv ; n and by extending the limit dynamic loading sliding-yield ratio ^r as cs ^r ¼ exp ð Þ; fn
ð22:128Þ
where cs is a material parameter. It follows from Eq. (22.128) that ^r ! 1 for fn ! 0 and ^r ! 1 for fn ! 1. Therefore, ^r becomes larger for a smaller normal contact stress fn , noting that the thickness of the fluid layer between contact surfaces increases approaching the fluid friction for a smaller normal contact stress. In contrast, ^r approaches unity so that the dynamic loading sliding-yield ratio r can become only slightly larger than the subloading sliding-yield ratio rs for an infinitely large normal contact stress fn approaching dry friction.
22.10.2
Numerical Experiments
The basic mechanical properties of the subloading overstress friction model formulated in the preceding section are examined below in numerical experiments using the constitutive Eq. (22.123) with Eq. (22.120) and adopting Eq. (22.21) for the Coulomb-type sliding-yield stress function f ðfÞ. In numerical simulations, the material constants in the subloading friction model are chosen as follows: Elastic sliding constant: an ¼ at ¼ 1000 GPa mm1 Static friction coefficient: ls ¼ 0:097; Kinetic friction coefficient: lk ¼ 0:085 Sliding-softening constant: j ¼ 1:176 mm1 Creep-hardening constant: n ¼ 206:2 min1 Normal friction-yield ratio evolution constant: ~ u ¼ 80 mm1 Furthermore, the three material constants gv , n and ^r added in the subloading overstress friction model are changed in the following five levels, fixing the other two material constants to be gv ¼ 1000, n ¼ 16 and ^r ¼ 1:6. Viscoplastic coefficient: gv ¼ 10; 100; 1000; 10000; 100000 min.mm1 Viscoplastic power coefficient: n ¼ 4; 8; 16; 32; 64 Limit dynamic loading frictionyield ratio: ^r ¼ 1:2; 1:6; 2:4; 3:2; 6:4
690
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.33 Influence of material parameters in subloading overstress friction model
The sliding velocity is set to be 0.1, 1, 100 and 1000 mm/min. The calculated results are shown in Fig. 22.33. The contact stress ratio ft =fn is larger for higher sliding velocities, exhibiting positive rate sensitivity. Higher contact stress ratios are predicted for larger values of gv and ^r for lower values of n. Curves of the contact stress ratio versus sliding displacement for various sliding velocities are shown in Fig. 22.34, indicating positive rate sensitivity, where the material constants are
22.10
Subloading-Overstress Friction Model
691
Fig. 22.33 (continued)
chosen as gv ¼ 1000; n ¼ 16 and ^r ¼ 1:6. The effect of the sliding velocity on the peak (maximum) and bottom (minimum) contact stress ratios in the contact stress ratio versus sliding displacement curves, which are read from Fig. 22.34, are plotted in Fig. 22.35. Here, the peak and bottom contact stress ratios increase with sliding velocity in a certain velocity range, whereas they converge to certain values outside this range.
692
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.34 Influence of sliding velocity
Fig. 22.35 Influence of sliding velocity on contact stress ratio
22.10.3
Comparison with Test Data
To examine the validity of the generalised subloading overstress model, a lubricated friction test was performed. A schematic diagram of the test apparatus is shown in Fig. 22.36. The test plate, which is placed between the tools, is subjected to the constant normal load (5kN) and is pulled up at constant velocity. The pulling force is measured by the load cell (maximum load: 100kN) attached to the upper part of the test plate. The test plate is made of galvannealed steel sheet with the friction area of width 30 mm, height 300 mm and thickness 0.7 mm. The tool steel SKD-11 of width 40 mm, height 30 mm and thickness 20 mm is used for the tools which grasp the test plate. Therefore, the friction contact area for the normal contact stress
22.10
Subloading-Overstress Friction Model
693
30mm
300mm
Fig. 22.36 Illustration of friction test apparatus
is 900 mm2 and therefore the normal contact stress is 5.56 MPa, whereas the friction contact area for the tangential contact stress is 1,800 mm2 . The friction surfaces were polished and coated with anti-rust oil prior to the tests. The pulling-up velocity of the test plate is set at five levels: 1; 10; 50; 100; 200 mm=min: The friction test was adopted supposing the press forming of thin sheet metal in the metal forming process. The measured relationship between the contact stress ratio ft =fn and the tangential sliding displacement ut is shown in Fig. 22.37a. The contact stress ratio first peaks and then gradually falls to a stationary value, exhibiting a positive rate sensitivity, i.e. larger contact stress ratios at higher sliding velocities. The simulation of the above-mentioned test result using Eq. (22.124) with Eqs. (22.21) and (22.120) is shown by the solid lines in Fig. 22.37a, using the following values for the material constants. an ¼ at ¼ 1000 GPamm1 ls ¼ 0:097; lk ¼ 0:085 j ¼ 58:82 mm1 ; n ¼ 1306 min1 ~ u ¼ 80 mm1 gv ¼ 950 min.mm1 ; n ¼ 16; ^r ¼ 1:8 The simulation of the test result using Eq. (22.116) with the power function is shown by the solid lines in Fig. 22.37b, using the following values for material constants.
694
22
Constitutive Equation for Friction: Subloading-Friction Model 0.16
0.16
0.12
0.12
ft / fn
ft / fn 0.08
0.08
200mm/min 100mm/min 50mm/min 10mm/min 1mm/min
0.04
200mm/min 100mm/min 50mm/min 10mm/min 1mm/min
0.04
0.00 0.0
2.0
4.0
6.0
ut (mm)
8.0
10.0
0.0
(a) Eq. (63)
2.0
4.0
6.0
ut (mm)
8.0
10.0
(b) Eq. (58) with power function
Fig. 22.37 Comparison with test data for various levels of sliding velocity
an ¼ at ¼ 1000 GPamm1 ls ¼ 0:097; lk ¼ 0:085 j ¼ 11:76 mm1 ; n ¼ 2062 min1 ~ u ¼ 80 mm1 gv ¼ 950 min.mm1 ; n ¼ 16;^r ¼ 1:8 These material constants are chosen so as to simulate the test result as closely as possible. However, it is impossible even to restrict the range from the highest and the lowest curves to the range in the test curves. Equation (22.117) with the exponential function would be much more appropriate than Eq. (22.116) with the power function for the prediction of real friction behaviour. However, the clarification of the definite physical background should be continued for the future. Based on Fig. 22.37a, a comparison of the calculated and test results for the influence of the sliding velocity on the peak (maximum) and bottom (minimum) values in the contact stress ratio versus sliding displacement curves is shown in Fig. 22.38, where the close coincidence can be observed.
22.11
Extension to Rotational and Orthotropic Anisotropy
The constitutive equation of friction explained in the preceding sections has been extended to describe the anisotropy by the rotation and orthotropy of sliding-yield surface (Hashiguchi 2007; Hashiguchi and Ozaki 2007) and its validity was verified by comparisons with experiments (Ozaki et al. 2012). The variation of friction behavior responding to the sliding direction is predicted pertinently by the extended constitutive equation of friction as will be described in this section.
22.11
Extension to Rotational and Orthotropic Anisotropy
695
Fig. 22.38 Comparison of calculated and test results for influence of sliding velocity on contact stress ratio
The simple surface asperity model is illustrated in order to obtain the insight for the anisotropy in Fig. 22.39. Here, the inclination of surface asperities to a particular direction would lead to rotational anisotropy. In addition, the anisotropic shapes and intervals of surface asperities would lead to the orthotropic anisotropy. Now, choosing the bases e1 and e2 in the directions of the maximum and the minimum principal directions of anisotropy, respectively, and letting e3 coincide with n to make the right-hand coordinate system fei g, the contact stress vector f and the rotational anisotropy vector b can be written as f ¼ f1 e1 þ f2 e2 þ f3 e3 b ¼ b1 e1 þ b2 e2 þ b3 e3
ð22:129Þ
In what follows, it is assumed that the rotational anisotropy vector b is fixed on
the contact stress surface, i.e. b 1 ¼ b 2 ¼ b 3 ¼ 0. Equation (22.129) is rewritten by f1 ¼ ft1 ; f2 ¼ ft2 ; ft3 ¼ fn and b1 ¼
bt1 ; b2 ¼ bt2 ; b3 ¼ 0 as follows: f ¼ ft1 e1 þ ft2 e2 fn e3 b ¼ bt1 e1 þ bt2 e2
ð22:130Þ
In what follows, the extension of the subloading-friction model with the Coulomb friction condition to the anisotropy will be considered in the following.
696
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.39 Surface asperity model suggesting the rotational and the orthotropic anisotropy
Invoking the orthotropic anisotropy proposed by Mroz and Stupkiewicz (1994), the sliding normal-yield and the sliding subloading-surfaces with the orthotropic and the rotational hardenings (see Fig. 22.40) are given by ^ vc ¼ l
ð22:131Þ
^ vc ¼ rl
ð22:132Þ
where ^ vc
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ^ ^g g ^ v1c 1 ; v^2 c 2 v1c2 þ ^ v2c2 ; ^ C1 C2
g1 ^ g1
ft ft 1 2 fn ; g2 fn g1 b1 ; ^ g2
ð22:133Þ
)
g2 b2
ð22:134Þ
C1 and C2 are the material constants designating the orthotropic anisotropy, whereas the e1 direction is chosen for the long axis of ellipsoid in the cross section of sliding-yield surface so that l designates the friction coefficient for C1 ¼ C2 ¼ 1 and b1 ¼ b2 ¼ 0 leading to the isotropic sliding-yield surface.
22.11
Extension to Rotational and Orthotropic Anisotropy
697
Fig. 22.40 Anisotropic sliding normal-yield and subloading-surfaces
The partial derivatives for Eq. (22.133) are given as follows: 9 > > > > > > > > > > > ^c i v @^ vci @ðfti =fn bi Þ=Ci fti > > ¼ ¼ 2 ¼ ðno sumÞ > > > @fn @fn fn Ci fn > > > = ^ v @^ vc 1 ci ^f ¼ 2^ v ¼ ci > vc @^ vci 2^ vc ci ^ > > > > > > vc i 1 vc i ^ @^ vc @^ vc @^ 1 ^fci > > ¼ ¼ ¼ ðno sumÞ > > > ^ vc fn Ci fn Ci @fti @^ vci @fti > > > @^ > vc i @^ vc @^ vc vc i @^ vc @^ 1 ^ ^ > > ; ¼ þ ¼ ð f c i vc 1 þ f c i vc 2 Þ > fn @fn @^ vci @fn @^ vci @fn
@^ vci @ðfti =fn bi Þ=Ci 1 ¼ ¼ ðno sumÞ @fti @fti fn Ci
Further, it holds from Eqs. (22.16) and (22.135) that
@^ vc 1 ^fc1 ^fc2 ^ ^ ¼ e þ e þ ðfc1 vc1 þ fc2 vc2 Þn @f fn C1 1 C2 2
ð22:135Þ
ð22:136Þ
698
22
ðIn nÞ
E
Constitutive Equation for Friction: Subloading-Friction Model
@^ vc 1 ^fc1 ^fc2 ^f v þ ^f v Þn ¼ ðIn nÞ e þ e þ ð c1 c1 c2 c2 @f fn C1 1 C2 2
1 ^fc1 ^fc2 e1 þ e2 ¼ C2 fn C1
@^ vc ¼ ½an e3 e3 þ at ðe1 e1 þ e2 e2 Þ @f
1 ^fc1 ^fc2 ^ ^ e þ e þ ðfc1 vc1 þ fc2 vc2 Þn fn C1 1 C2 2 ^
fc1 ^fc2 1 ^ ^ a þ a e þ e ¼ t n ðfc1 vc1 þ fc2 vc2 Þn C1 1 C2 2 fn ^f c1 C1
ð22:138Þ
^f
e1 þ Cc22 e2 nt ¼ rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ^f ð Cc11 Þ2
ð22:137Þ
ð22:139Þ
^f þ ð Cc22 Þ2
^f ^f ^
c1 e1 þ Cc22 e2 fc1 ^fc2 @^ vc 1 C 1 at E nt ¼ e þ e þ an ð^fc1 vc1 þ ^fc2 vc2 Þn rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi @f C1 1 C2 2 fn ^f ^f ð Cc11 Þ2 þ ð Cc22 Þ2 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ^ 2 ^ 2 fc1 fc 2 at ¼ þ fn C1 C2
ð22:140Þ Substituting Eqs. (22.16), (22.130) and (22.136)–(22.140) into Eqs. (22.48) and (22.50), we obtain the sliding velocity vs. contact stress rate and its inverse relation as follows:
u ¼
¼
1 1 n n þ ðIn nÞ ð ft1 e1 þ ft1 e2 fn nÞ an at
1 ^fc1 ^fc2 ^f ^f e1 þ e2 þ ð^fc1 vc1 þ ^fc2 vc2 Þn ð ft1 e1 þ ft2 e2 fn nÞ mc c1 c2 fn C1 C2 C1 e 1 þ C 2 e 2 s ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi þ p
2 2 m ^f ^f c1 c2 þ C1 C2 1 1 fn n ð ft1 e1 þ ft2 e2 Þ at an
þ
^ 1 f c1 fn ½ C1 ft1
þ
^f c2 C2 ft2
ð^fc1 vc1 þ ^fc2 vc2 Þ fn mc mp
^f
^f
c1 c2 C1 e 1 þ C2 e 2 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
^f c1 C1
2
þ
^f c2 C2
2
ð22:141Þ
22.11
Extension to Rotational and Orthotropic Anisotropy
699
f ¼ fan n n þ at ðe1 e1 þ e2 e2 Þg½ u 1 e1 þ u 2 e2 u n n ^
fc1 ^fc2 1 ^ ^ at þ an ðfc1 vc1 þ fc2 vc2 Þn e þ C1 1 C2 fn ^f ^f c1 # * c + e1 þ c2 e2 ðu 1 e1 þ u 2 e2 u nÞ m C1 C2 n sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ^ 2 ^ 2 ^ 2 ^ 2 f f f c1 f c2 c c a 1 2 t m p þ fn þ þ C1 C2 C1 C2
¼ at ð u 1 e1 þ u 2 e2 Þ an u n n ^
^ c+ * 1 a fc1 u þ fc2 u ^f ^f ^ ^ c1 c2 t 1 2 an ðfc1 vc1 þ fc2 vc2 Þvn m fn C1 C2 C1 e1 þ C2 e2 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi at ^ 2 ^ 2 ^ 2 ^ 2 f c1 f c2 f c1 fc2 mp þ afnt þ þ C1 C2 C1 C2
ð22:142Þ The calculation for sliding with the orthotropic anisotropy must be performed in the coordinate system with the principal axes of orthotropic anisotropy, i.e. ðe1 ; e2 ; nÞ. We examine below the basic response of the present friction model by numerical experiments and comparisons with test data for the linear sliding phenomenon
without a normal sliding velocity leading to u n ¼ 0. The contact stress rate vs. sliding velocity relation is given by substituting
un ¼ 0 into Eq. (22.142) as
f ¼ at ð u 1 e1 þ u 2 e2 Þ ^
^ ^f ^f c+ c1 c2 * 1 a fc1 u þ fc2 u m e þ e t 1 2 C1 C2 fn C1 1 C2 2 s ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi at ^ 2 ^ 2 ^ 2 ^ 2 fc 1 fc2 fc1 fc 2 mp þ afnt þ þ C1 C2 C1 C2
ð22:143Þ
where mp and mc are given by Eq. (22.46). In what follows, we demonstrate the basic response of the present anisotropic friction model through numerical experiments for the linear sliding phenomenon without a rigid-body rotation under a constant normal traction. The calculations described in the following have been executed by Dr. Shingo Ozaki, Yokohama National Univ. The nine material parameters and the initial value are selected as
700
22
Constitutive Equation for Friction: Subloading-Friction Model
Fig. 22.41 Influence of rotational anisotropy on relation of contact stress ratio (Ozaki et al. 2012)
l0 ¼ ls ¼ 0:4; lk ¼ 0:2 j ¼ 0:5 mm1 ; n¼ 25s1 ~ u ¼ 1; 000 mm1 an ¼ at ¼ 1; 000 N=mm3 under the condition fn ¼ 10 MPa; vt ¼ 1:0 mm=s The variations in the contact stress ratio ft =fn with the tangential sliding displacement ut is shown in Fig. 22.41. Here, we assume the three sets of the parameters for orthotropy C1 and C2 without the rotational anisotropy, i.e., b ¼ 0. Then, the
constant sliding velocity of the magnitude ut ¼ 1:0 mm/s was given in the directions
22.11
Extension to Rotational and Orthotropic Anisotropy
701
Fig. 22.42 Influence of rotational anisotropy on contact stress (Ozaki and Hashiguchi 2012)
0 ; 45 and 90 from the base vector e1 of orthotropy. As shown in this figure, the friction behavior varies with the sliding direction because of orthotropic anisotropy. The influence of the rotational anisotropy, i.e., the parameter b ¼ fb1 ; b2 ; 0g on the relation of the contact stress ratio with the tangential sliding displacement is shown in Fig. 22.42. In this calculation, we set the orthotropic anisotropic parameters as C1 ¼ 1:0 and C2 ¼ 0:8, and set the rotational hardening parameters b1 as 0.0, 0.05, and 0.1 and b2 ¼ 0:0. Then, the constant tangential sliding rate
ut ¼ 1:0 mm/s is given into mutually opposite directions. It is confirmed that the frictional properties for mutually opposite directions of sliding are different from each other. The description of the differences in friction coefficients in opposite directions is important in biomimetic textures and in drive systems of off-the-road vehicles and robots. Some other verifications of the pertinence of the present model by the comparisons with test data are referred to Ozaki et al. (2012).
Final Remarks
The continuum mechanics and the elastoplasticity are comprehensively described in this book. In particular, the subloading surface model is described in detail through this book. The basic features of the subloading surface model are summarized as follows: (1) It is based on the quite natural concept that the plastic deformation is developed as the stress approaches the yield surface and thus it possesses the high generality and the capability of describing accurately irreversible deformations. (2) It fulfills the smoothness condition, describing always continuous variation of tangent stiffness modulus and thus depicting the smooth elastic-plastic transition. (3) It possesses the automatic controlling functions to attract the stress to the normal-yield surface, so that the stress is pulled-back to the yield surface when it goes over the surface in numerical calculations. Further, the plastic strain is also automatically pulled-back to the stagnation surface. By virtue of these rigorous physical backgrounds, the subloading surface model is endowed with the following rigorous descriptions distinguishable from the other elastoplastic constitutive models. 1. Plastic strain rate is predicted even for any low stress level. Then, cyclic loading behavior is predicted accurately even for infinitesimal loading amplitude. The other models, e.g. the cylindrical yield (Chaboche), the multi and the two surface models are incapable of describing the cyclic loading behavior appropriately. 2. Stress is automatically pulled-back to the yield surface when it goes out from the yield surface for the large strain increment in the explicit (stress integration) method.
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
703
704
Final Remarks
3. Deformation of various solids unlimited to metals, e.g. soils, rocks, concrete, glass, and polymers can be described. The other cyclic plasticity models are limited to the description of metal deformation. 4. Inelastic strain rate induced by the stress rate tangential to the yield surface, i.e. tangential-inelastic strain rate is described appropriately, fulfilling the continuity and smoothness conditions, which is required to describe the non-proportional loading behavior. All the other models assume the purely-elastic domain so that they predict the tangential-inelastic strain rate induced suddenly at the moment when the stress reaches the yield surface if the tangential-inelastic strain rate is incorporated, violating the continuity and the smoothness conditions at the moment. 5. The subloading surface model is extended to the subloading-overstress model to describe the monotonic and cyclic deformation behavior in the general rate from the quasi-static to the dynamic and impact deformation and thus subloading surface model for the elastoplastic deformation can be discussed. The other models are incapable of describing the viscoplastic deformation behavior for cyclic loading and at high rate, predicting an elastic response for an impact loading. 6. Damage phenomenon under cyclic loading is described appropriately, which leads to a softening behavior in general. The other cyclic plasticity models are incapable of describing the cyclic damage phenomenon with a softening pertinently. 7. Constitutive equation of fatigue would be described pertinently, which is required to describe plastic strain rate induced in low stress level and small stress amplitude. The other models are incapable of describing the fatigue phenomenon, predicting only an elastic strain rate in low stress level and small stress amplitude. 8. Constitutive equation of phase-transformation of metals can be described pertinently. The other models are incapable of describing the phasetransformation phenomenon pertinently. 9. Friction phenomenon is described pertinently, including the transition from static to kinetic friction by the sliding, the recovery of static friction with elapse of time and the both of positive and negative rate-sensitivities. The negative and the positive rate sensitivities are relevant to the dry and the lubricated (fluid) friction, respectively. The other model is incapable of describing these friction phenomena. 10. Crystal plasticity analysis can be executed pertinently, in which calculation of slips in numerous number of slip systems are required. The yield judgment is unnecessary since the smooth elastic-plastic transition is described and the resolved shear stress is automatically pulled-back to the critical shear stress. The other models are inapplicable to the crystal plasticity analysis rigorously because the yield judgment and the operation to pull back the resolved shear stresses to the critical shear stress are required in numerous number of slip systems. Then, impertinent analysis using the creep model has been performed widely after Pierce et al. (1982, 1983). It is impertinent to adopt the rate-dependent model for the rate-independent deformation phenomenon. In
Final Remarks
705
addition, it should be noted that the creep model is impertinent such that it predicts a creep shear strain rate even in unloading process of the resolved shear stress from the critical shear stress. The subloading-overstress model should be adopted in the rate-dependent crystal plasticity analysis. 11. The multiplicative hyperelastic-based plasticity can be formulated readily based on the subloading surface model, which is capable of describing the cyclic loading behavior exactly for finite elastoplastic deformation/rotation. The multiplicative hyperelastic-based plasticity cannot be formulated by the other cyclic plasticity models. Thus, the subloading surface model is endowed inherently with the high generality and flexibility for the description of irreversible mechanical phenomena of solids. It has been attained by the incorporation of the normal-yield ratio R which contains the plenty of important information on the description of the irreversible (plastic and viscoplastic) deformation/sliding under the monotonic and the cyclic loading process. Eventually, it can be concluded that Subloading surface concept must be the universal natural law for irreversible deformation/sliding phenomenon of solids, which is inevitable to describe the irreversible mechanical phenomena in wide classes of solids, ranging from monotonic to cyclic loadings, from quasi-static to impact loadings, from infinitesimal to finite deformations and from micro to macro phenomena. Then, the elastoplasticity theory will be developed steadily (breaking through the stagnation) by incorporating the exact formulation of the subloading surface model, although it has stagnated during a half century since the study on the unconventional (cyclic) elastoplasticity aiming at description of the plastic strain rate caused by the rate of stress inside the yield surface has started in the 1960s. The Hashiguchi (subloading surface) model is implemented to the commercial FEM software Marc marketed by MSC Software Corporation, so that Marc users can apply this model to deformation analyses. The implementation was highly supported by Dr. Motohatu Tateishi, MSC Soft. Corp., Japan. The computer programs for the subloading-elastoplastic and the subloading-overstress models for metals based on the subloading surface model are provided in Appendix L. On explicit and implicit numerical calculation methods of state variables The numerical calculation is not described in this book, while it is described in the former books (Hashiguchi, 2017, 2000; Hashiguchi and Yamakawa, 2012) in detail. The calculation of the state variables, i.e. the stress, the strain and the internal variables in each material point can be performed by the two different methods. One is the explicit method, i.e. the forward-Euler method and the other is the implicit method by the return-mapping projection. The merits/demerits of these methods are commented below briefly. The merits of the explicit method are 1) The calculation can be performed immediately if the explicit relation of the stress rate and the strain rate is formulated and 2) The calculation can be performed easily even for the complex constitutive
706
Final Remarks
equations representing the precise material properties and the deformation behaviors under the complex and/or dynamic loading condition and 3) The constitutive equation can be modified and thus the formulation can be performed easily and 4) The determination of the material constants can be performed directly for the stress rate—strain rate relation. Noticeably, the subloading surface model is furnished with the distinguished abilities to pull-back the stress to the yield surface automatically (Sect. 9.3) and to control the isotropic hardening stagnation surface so as to involve the plastic strain (Sect. 12.2) and the Mullins-damage surface so as to involve the viscoelastic strain tensor (Sect. 18.3). Therefore, the large incremental steps can be input in the numerical calculation by the subloading surface model in the explicit method. On the other hand, the merit of the implicit calculation is no more than the high efficiency and accuracy of calculation. However, the cumbersome preparations to derive the partial-derivatives of the yield condition and the evolution equations of internal variables by the independent variables are required. Therefore, the adoption of the implicit method to the numerical calculation for the constitutive relation representing precise material properties would not be easy. In fact, the application of the implicit method has been almost limited to the Mises metals but hardly applied to the other materials, e.g. soils, rocks, concrete and glass exhibiting the plastic compressibility/pressure dependency and the dependence on the third deviatoric stress invariant of the yield stress. Consequently, the explicit method would be suitable to the calculation to survey the detailed mathematical response of the constitutive equation and to the FE analysis of the solids and structures composed of materials with the constitutive equation taken account of the precise constitutive properties. On the other hand, the implicit method would be suitable to the FE analysis of complex structures composed of simple materials for which the above-mentioned preparations are attained already. However, there is no commercial software adopting the implicit method by the return-mapping up to present. The subloading surface model for the elastoplastic deformation of metals in the explicit method is installed into the commercial software Marc form the 2017 version which can be used by all Mark users. On Various Irrational Models and Formulations Diffusing Widely It is quite regretful for the sound development of applied mechanics and technologies in engineering that various irrational theories have been proposed and are diffused widely. Typical examples are shown below. Cyclic Kinematic Hardening Models with purely-elastic domain: Chaboche model Although the subloading surface model is widely used for soils, it has not yet been widely applied to metals. The reason for this background is that the cylindrical yield surface models, i.e. the Chaboche (Ohno) model is widely used for metals and the Yoshida model is used for the springback analysis. However, they are incapable of describing the cyclic loading behavior under the general stress amplitude even for
Final Remarks
707
metals, since they assume a yield surface enclosing a purely elastic domain and lack the basic structure for the cyclic plasticity as described in Chap. 10. It is much desired that the researchers and the engineers involved in elastoplastic deformation analysis will wake up to this fact as soon as possible and carry out a rational deformation analysis for the sound development of solid mechanics and practical engineering. Historically rare and sad fact in solid mechanics: numerous number of papers on the modifications of the Chaboche model have been issued resulting in various quite complicated formulations but all of them have ended in failure and will never succeed to describe the plastic/viscoplastic deformation behavior pertinently forever as far as the crucially-important variable, i.e. the normal-yield ratio R introduced in the subloading surface model is not adopted in the formulation. This fact has severely prevented the sound development of plastic/viscoplastic theories during a half century and must be the worst disgrace in the history of solid mechanics. Creep Model The creep model for the rate-dependent irreversible deformation is irrational such that it always predicts the irreversible deformation in any low stress level, while its basic structure is irrelevant to that of the plastic constitutive equation. Therefore, the users of the creep model are obliged to use the plastic constitutive model for the quasi-static deformation but the creep model for the time-dependent deformation. In other words, the creep model is the typical ad voc. model for the rate-dependent deformation behavior only for a high rate of deformation. Unfortunately, however, it is applied widely for the prediction of irreversible deformation of not only macroscopic deformation but also crystal plasticity analysis. The overstress model must be used instead of the creep model. Earthquake Prediction and Rate-and-State Model The earthquake is the irreversible (plastic) sliding phenomenon and thus it must be formulated within the framework of the elasoplasticity. Unfortunately, however, the primitive rate-and-state model (e.g. Dieterich 1978; Ruina 1980, 1983; Rice and Ruina 1983; Rice et al. 2001) which ignores this crucial fact is used widely yet for the earthquake prediction. The earthquake disasters will never prevented as far as the rate-and-state model is used for earthquake prediction. The earthquake prediction should be attained by applying the subloading-friction model with the saturation of the tangential contact stress described in Sect. 22.9, since a quite high contact pressure far exceeding the shear strengths of the continental and the ocean plates is applied between them. Formulation of Plastic Flow Rule by Second Law of Thermodynamics The thermodynamic consideration with the statistical mechanics is the effective tool in formulating the basic equation for deformation behavior of simple materials, e.g. the entropic elasticity leading to the Neo-Hookean elastic equation. However, it would not be valid to the formulation of the complex mechanical behavior including the inelastic deformation behavior.
708
Final Remarks
Nevertheless, the formulation for the associated flow rule of plastic strain rate under the statement that it can be derived from the second law of thermodynamics is fashioned widely (J. A. Lemaitre, S. Murakami, A. Menzel, J. L. Chaboche, G. Z. Voyiadjis, I. N. Vladimirov, M. Wallin, etc.). However, it is to be the typical predetermined harmony of the associated flow, so that any novel valuable result has not been derived, and thus inhibits the serious consideration for the flow rule. The second law of thermodynamics is a major law that must be satisfied by all natural irreversible phenomena. On the other hand, there is no necessity in the associated flow rule which is merely the assumption. In fact, the non-associated flow rule with a plastic potential surface different from a yield surface has been adopted widely for frictional materials, e.g. soils and rocks based on experimental observations. It would be unable to prove that the non-associated flow rule violates the Clausius-Duhem inequality (Hashiguchi 2001b). It would be reckless to expect to derive a concrete equation for the complex behavior in plastic deformation of solids from such a fundamental law. In addition, the thermodynamic description hinders the understanding of the theory and thus it should be stopped also from an educational point of view. In facts, there is no necessity to include the thermodynamic explanation and the formulations are understood more straightly by excluding the thermodynamic description. Natural scientists are required to take up sincere attitude towards the deep consideration whether theories or models flashed into mind are really acceptable. All the defects seen in the above-mentioned irrational models and formulations would have been solved out already by the subloading surface model. Truth will be recognized and accepted in history. What a lot of unreasonable models have been proposed and used, and thus what a lot of waste have been repeated ! The waste is real in the Chaboche model, the Mroz model, the Dafalias model, the bounding surface model with radial-mapping, the repeated modifications of the Armstrong-Frederick kinematic hardening rule, the creep model, the rate-and-state friction model, the two-scale model for fatigue, etc. However, a long period of time would be required for the epoch-making model differing basically from the ordinary models to be accepted widely.
Appendix A: Projection of Area
Consider the projection of an area having the unit normal vector n onto the surface having the normal vector m in Fig. A.1. Now, suppose the plane (habcd in Fig. A.1) which contains the unit normal vectors m and n. Then, consider the line ef obtained by cutting the area by this plane. Further, divide the area having the unit normal vector n to the narrow bands perpendicular to this line and their projections onto the surface having the normal vector m. The lengths of projected bands are same as the those of the original bands but the projected widths db are obtained by multiplying the scalar product of the unit normal vectors, i.e. m n to the original widths db. Eventually, the projected areas da are related to the original areas da as follows: da ¼ m nda
ðA:1Þ
a e
n b da
db
d
m
f
db = n mdb da = n mda
c
Fig. A.1 Projection of area © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
709
Appendix B: Logarithmic Spin
The logarithmic spin proposed by Xiao and his colleagues (Xiao 1995; Xiao et al. 1997, 1999) will be delineated below. We adopt the Eulerian-logarithmic strain tensor eð0Þ in Eq. (4.42) and its coro ð0ÞE
tational rate e
with the Eulerian spin XE . Further, consider to add a spin x to XE ð0ÞLog
so that the corotational rate e d, i.e.
with the spin XE þ x coincides to the strain rate
d ¼ eð0ÞLog eð0Þ þ eð0Þ XLog XLog eð0Þ
¼ eð0Þ þ eð0Þ ðXE þ xÞ ðXE þ xÞeð0Þ
¼ eð0Þ þ eð0Þ XE XE eð0Þ þ ðeð0Þ x xeð0Þ Þ
ðA:2Þ
¼ eð0ÞE þ ðeð0Þ x xeð0Þ Þ where XLog is the logarithmic spin, i.e. XLog ¼ XE þ x
ðA:3Þ
It follows from Eq. (A.2), noting ðeð0Þ Þab ja6¼b ¼ 0 and Eqs. (4.54) and (4.142), that dab ¼ ððxeð0Þ eð0Þ xÞÞab ¼ xar erb þ ear xrb ¼ xab ln kb þ ln ka xab ¼ xab lnðka =kb Þ for a 6¼ b
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
711
712
Appendix B: Logarithmic Spin
from which one has xab ¼
dab lnðka =kb Þ
for a 6¼ b
ðA:4Þ
Taking account of Eqs. (4.107) and (A.3) for the Eulerian spin into Eq. (A.3), we have XLog ¼
XX a
b
ðaÞ XLog nðbÞ ¼ ab n
a
ðXEab þ xab ÞnðaÞ nðbÞ
b
# ) 1 ¼ wab þ þ dab nðaÞ nðbÞ 1 ðka =kb Þ2 lnðka =kb Þ a b " # X X 1 þ ðka =kb Þ2 1 ¼ wþ þ Pa dPb 1 ðka =kb Þ2 lnðka =kb Þ a b XX
(
XX
"
1 þ ðka =kb Þ
2
ðA:5Þ
where Pr is defined by the Eulerian triad fnðrÞ g as Pr ¼ nðrÞ nðrÞ ðno sumÞ
ðA:6Þ
noting Pa dPb ¼ nðaÞ nðaÞ dsc nðsÞ nðcÞ nðbÞ nðbÞ ¼ nðaÞ das dsc dkb nðbÞ ¼ dab nðaÞ nðbÞ The logarithmic spin is applied to the stress and internal variables in the hypoelastic-based plastic constitutive equation. The hypoelastic equation by the corotational rates of the Cauchy stress is described as follows: Log
r
¼ E : de ¼ E :ðd dp Þ
ðA:7Þ
In what follows, it will be verified that the hypoelastic elastic equation with the corotational rates due to the logarithmic spin leads to the hyperelastic equation. The hypoelastic equation based on the logarithmic corotational rates in the Hooke’s type is described by Eq. (7.110) as Log
r
¼ E: d ¼
2 K G ðtr dÞI þ 2Gd 3
ðA:8Þ
Appendix B: Logarithmic Spin
713
Here, we have the following expression in terms of ln V, noting Eqs. (4.42) and (A.2). ð0ÞLog
d¼e
¼ ðln VÞLog
ðA:9Þ
by which Eq. (A.8) is rewritten as Log
r
2 ¼ K G ½trðln VÞLog I þ 2Gðln VÞLog 3
ðA:10Þ
If the elastic parameters K and G are constants, Eq. (A.10) is rewritten as Log
r
Log 2 ¼ K G ½trðln VÞI þ 2Gðln VÞ 3
ðA:11Þ
2 K G ½trðln VÞI þ 2l ln V 3
ðA:12Þ
which results in r¼
Eq. (A.12) is to be the Cauchy elastic equation. The hypoelastic equation with the corotational rate based on the continuum spin causes the oscillation of stress in the simple shear deformation as was shown by Dienes (1979). Further, the hypoelastic-based plasticity with the kinematic hardening and the corotational rate based on the continuum spin causes the oscillation of stress in the simple shear deformation as was shown by Nagtegaal and de Jong (1982). The oscillation is excluded by use of the logarithmic corotational rates as shown by the numerical experiments (Xiao et al. 1997; Brepols et al. 2014; Shutov and Ihlemann 2014). However, it should be noted that the logarithmic corotational rates possesses the following limitations. (1) The hyperelastic equation is obtained only from the Hooke’s type hypoelastic equation with the constant elastic parameters. Therefore, it is applicable to metals but inapplicable to pressure-dependent materials, e.g. soils, rocks and concretes the elastic parameters of which depend on the pressure. (2) The fact that the rotation of the substructure in material is not determined by the geometrical change of the external appearance of material is ignored. The physically meaningful spin is the spin of substructure in material which is suppressed by the plastic dissipative deformation from the rigid-body rotation so that the plastic spin must be incorporated in addition to the spin based on the geometrical change of the outside appearance as described in Sect. 3.4.2. and Chap. 19.
Appendix C: Matrix Representation of Tensor Relations
The representations of the second-order symmetric tensor t and the fourth-order symmetric tensor T by displaying their component is called the Voigt representations. t ¼ tij ¼ ft11 t22 t33 t23 t32 t31 t13 t12 t21 gT ¼ ft1 t2 t3 t4 t5 t6 gT 2
T1111 T1122 T1133 T1123 T1132 T1131 T1113 T1112 T1121
ðA:13Þ 3
7 6T 6 2211 T2222 T2233 T2223 T2232 T2231 T2213 T2212 T2221 7 7 6 6 T3311 T3322 T3333 T3323 T3332 T3331 T3313 T3312 T3321 7 7 6 7 6 6 T2311 T2322 T2333 T2323 T2332 T2231 T2313 T2312 T2321 7 7 6 7 T ¼ ½Tijkl ¼ 6 6 T3211 T3222 T3233 T3223 T3232 T3231 T3213 T3212 T3221 7 7 6 6 T3111 T3122 T3133 T3123 T3132 T3131 T3113 T3112 T3121 7 7 6 6 T1311 T1322 T1333 T1323 T1332 T1331 T1313 T1312 T1321 7 7 6 7 6 4 T1211 T1222 T1233 T1223 T1232 T1231 T1213 T1212 T1221 5 2
T11 6 6 T21 6 6 6 T31 6 ¼6 6 T41 6 6 6 T51 4
T12 T22 T32 T42 T52
T61 T62
T2111 T2122 T2133 T2123 T2132 T2131 T2113 T2112 T2121 3 T13 T14 T15 T16 7 T23 T24 T25 T26 7 7 7 T33 T34 T35 T36 7 7 7 T43 T44 T45 T46 7 7 7 T53 T54 T55 T56 7 5 T63 T64 T65 T66
ðA:14Þ
where the subscript numbers are replaced as 11 ! 1; 22 ! 2; 33 ! 3; 23; 32 ! 4; 31; 13 ! 5; 12; 21 ! 6, 31; 13 ! 5; 12; 21 ! 6.
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
715
716
Appendix C: Matrix Representation of Tensor Relations
If T satisfies the major symmetry Tijkl ¼ Tklij , i.e. Trs ¼ Tsr , it can be represented as 2 6 6 6 T¼6 6 6 4
T11
T12 T22
T13 T23 T33 Sym:
T14 T24 T34 T44
T15 T25 T35 T45 T55
3 T16 T26 7 7 T36 7 7 T46 7 7 T56 5 T66
ðA:15Þ
The fourth-order tracing identity tensor in Eq. (1.191) possesses both of the minor and the major symmetries and thus it is represented as follows: ⎡δ11δ11 δ11δ 22 δ11δ 33 δ11δ 23 δ11δ 31δ11δ12 ⎤ ⎡1 1 1 0 0 0 ⎤ ⎢ δ 22δ 22 δ 22δ 33 δ 22δ 23 δ 22δ 31δ 22δ12 ⎥⎥ ⎢ 1 1 0 0 0 ⎥ ⎢ ⎢ ⎥ ⎢ δ 33δ 33 δ 33δ 23 δ 33δ 31δ 33δ12 ⎥ ⎢ 1 0 0 0 ⎥ = [δ ijδ kl ] = ⎢ ⎥=⎢ ⎥ ðA:16Þ 0 0 0⎥ δ12δ12 δ12δ 31δ12δ12 ⎥ ⎢ ⎢ ⎢ Sym. δ 31δ 31δ 31δ12 ⎥ ⎢ Sym. 0 0 ⎥ ⎢ ⎥ ⎢ ⎥ 0⎦ ⎢⎣ δ12δ12 ⎥⎦ ⎣ The fourth-order tracing identity tensor in Eq. (1.192) possesses both of the minor and the major symmetries and thus it is represented as follows: ⎡δ11δ11 δ12δ12 δ13δ13 δ12δ13 δ13δ11δ11δ12 ⎤ ⎡1 0 0 0 0 0 ⎤ ⎢ δ 22δ 22 δ 23δ 23 δ 22δ 23 δ 23δ 21δ 21δ 22 ⎥⎥ ⎢ 1 0 0 0 0 ⎥ ⎢ ⎢ ⎥ ⎢ δ 33δ 33 δ 32δ 33 δ 33δ 31δ 31δ 32 ⎥ ⎢ 1 0 0 0⎥ ⎥=⎢ = [δ ik δ jl ] = ⎢ ⎥ ðA:17Þ Sym. δ 22δ 33 δ 23δ 31δ 21δ 32 ⎥ ⎢ Sym. 1 0 0 ⎥ ⎢ ⎢ 1 0⎥ δ11δ 22 δ12δ 23 ⎥ ⎢ ⎢ ⎥ ⎢ ⎥ 1⎦ δ11δ 22 ⎦⎥ ⎣ ⎣⎢ The following matrix must be adopted to transform between the infinitesimal scientific strain e and the infinitesimal engineering strain ε%, i.e. ε% = % :ε and ε = I ε% . %
⎡1 0 0 0 0 0 ⎤ ⎡1 0 0 0 0 0 ⎤ ⎢ 1 0 0 0 0⎥ ⎢ 1 0 0 0 0⎥ ⎢ ⎥ ⎢ ⎥ ⎢ ⎢ 1 0 0 0⎥ 1 0 0 0⎥ % =⎢ ⎥, = ⎢ ⎥ 2 0 0⎥ % ⎢ 1/2 0 0 ⎥ ⎢ ⎢ Sym. 2 0 ⎥ ⎢ Sym. 1/2 0 ⎥ ⎢ ⎥ ⎢ ⎥ 2⎦ 1/2 ⎦ ⎣ ⎣
ðA:18Þ
Appendix C: Matrix Representation of Tensor Relations
717
The deviatoric projection tensor ' in Eq. (1.194) possesses both of the minor and the major symmetries and thus it is represented as follows: ⎡1 0 0 0 0 0 ⎤ ⎡1 1 1 0 0 0 ⎤ ⎡ 2 / 3 − 1/ 3 − 1/ 3 0 ⎢ 1 0 0 0 0⎥ ⎢ 1 1 0 0 0⎥ ⎢ 2 / 3 − 1/ 3 0 ⎢ ⎥ ⎢ ⎥ ⎢ ⎢ ⎥ ⎢ ⎢ ⎥ 1 0 0 0 1 0 0 0 2/3 0 1 ' = − 1 = ⎢ ⎥−3⎢ ⎥ =⎢ 3 0 0 1 0 0 0 1 Sym. ⎢ ⎥ ⎢ Sym. ⎥ ⎢ ⎢ Sym. 1 0 ⎥ ⎢ ⎢ ⎥ 0 0 ⎢ ⎥ ⎢ ⎥ ⎢ 1 0⎦ ⎣ ⎣ ⎦ ⎣
0 0 0 0 1
0⎤ 0 ⎥⎥ 0⎥ ⎥ 0⎥ 0⎥ ⎥ 1⎦
ðA:19Þ Here, note that the following matrix must be adopted to transform the engineering strain to the deviatoric scientific strain, i.e. ε' = %' :ε% .. ⎡1 1 1 0 0 0 ⎤ ⎡ 2 / 3 − 1/ 3 − 1/ 3 0 0 ⎡1 0 0 0 0 0 ⎤ ⎢ 1 1 0 0 0⎥ ⎢ ⎢ 1 0 0 0 0 ⎥ 2 / 3 − 1/ 3 0 0 ⎢ ⎥ ⎢ ⎢ ⎥ ⎢ ⎥ ⎢ ⎢ ⎥ 1 0 0 0 1 0 0 0 2/3 0 0 1 %' = [% ] − 1 [ ] = ⎢ ⎥=⎢ ⎥−3⎢ 3 1/2 1/2 0 0 ⎥ 0 0 0⎥ ⎢ ⎢ ⎢ Sym. ⎢ Sym. 0 0 ⎥ ⎢ ⎢ Sym. 1/2 0 ⎥ 1/2 ⎢ ⎥ ⎢ ⎢ ⎥ 1/2 ⎦ 0⎦ ⎣ ⎣ ⎣
The elastic stress-strain relation is represented as follows: 9 8 9 2 38 E11 E12 E13 E14 E15 E16 > e1 > r1 > > > > > > > > > > > > e2 > r > E22 E23 E24 E25 E26 7 > 6 > > > > > 7> = 6 = < 2> < 6 7 r3 E E E E e 33 34 35 36 3 7 ¼6 E44 E45 E46 7 2e > > 6 > r4 > > > 6 7> > > > > 4> > > 4 r5 > Sym: E55 E56 5> 2e > > > > > > > ; : ; : 5> r6 E66 2e6
0 0 0 0 0 1/2
⎤ ⎥ ⎥ ⎥ ⎥ 0⎥ ⎥ ⎥ ⎦
ðA:20Þ
which is represented in terms of the engineering strain ~ei as follows: 8 9 2 E11 r1 > > > > > > 6 > r2 > > > > = 6 < > 6 r3 ¼6 6 r > > 4 > > 6 > > > > 4 r5 > > > > : ; r6
E12 E22
E13 E23 E33 Sym:
E14 E24 E34 E44
E15 E25 E35 E45 E55
38 9 ~e1 > E16 > > > > > > ~e2 > E26 7 > > 7> = < > E36 7 e 3 7 7 ~e > E46 7> > > > 4> ~e > E56 5> > > > ; : 5> ~e6 E66
ðA:21Þ
718
Appendix C: Matrix Representation of Tensor Relations
E: n is calculated as follows: ðE: nÞij ¼ Eij11 n11 þ Eij22 n22 þ Eij33 n33 þ Eij23 n23 þ Eij32 n32 þ Eij31 n31 þ Eij13 n13 þ Eij12 n12 þ Eij21 n21 which is represented using N ij ¼ nij ði ¼ jÞ;
N ij ¼ 2nij ði 6¼ jÞ as follows:
ðE: nÞij ¼ Eij11 N 11 þ Eij22 N 22 þ Eij33 N 33 þ Eij23 N 23 þ Eij31 N 31 þ Eij12 N 12 ðA:22Þ which is further described in the Voigt representation as ðE: nÞi ¼ Ei1 N 1 þ Ei2 N 2 þ Ei3 N 33 þ Ei4 N 4 þ Ei5 N 5 þ Ei6 N 6 ¼
6 X
Eij N j ðA:23Þ
j¼1
The scalar product n : E : n is represented as n : E : n ¼ Eijkl nkl nij ¼ ðE: nÞij nij ¼
3 X
½ðEij11 n11 þ Eij22 n22 þ Eij33 n33
i;j¼1
þ Eij23 n23 þ Eij32 n32 þ Eij31 n31 þ Eij13 n13 þ Eij12 n12 þ Eij21 n21 Þnij ¼
3 X
½Eij11 n11 þ Eij22 n22 þ Eij33 n33
i;j¼1
þ 2ðEij23 n23 þ Eij31 n31 þ Eij12 n12 Þnij ¼
3 X
½ðEij11 N 11 þ Eij22 N 22 þ Eij33 N 33 þ Eij23 N 23 þ Eij31 N 31 þ Eij12 N 12 Þnij
i;j¼1
¼ ðE1111 N 11 þ E1122 N 22 þ E1133 N 33 þ E1123 N 23 þ E1131 N 31 þ E1112 N 12 Þn11 þ ðE2211 N 11 þ E2222 N 22 þ E2233 N 33 þ E2223 N 23 þ E2231 N 31 þ E2212 N 12 Þn22 þ ðE3311 N 11 þ E3322 N 22 þ E3333 N 33 þ E3323 N 23 þ E3331 N 31 þ E3312 N 12 Þn33 þ 2ðE2311 N 11 þ E2322 N 22 þ E2333 N 33 þ E2323 N 23 þ E2331 N 31 þ E2312 N 12 Þn23 þ 2ðE3111 N 11 þ E3122 N 22 þ E3133 N 33 þ E3123 N 23 þ E3131 N 31 þ E3112 N 12 Þn31 þ 2ðE1211 N 11 þ E1222 N 22 þ E1233 N 33 þ E1223 N 23 þ E1231 N 31 þ E1212 N 12 Þn12 ¼ ðE1111 N 11 þ E1122 N 22 þ E1133 N 33 þ E1123 N 23 þ E1131 N 31 þ E1112 N 12 ÞN 11 þ ðE2211 N 11 þ E2222 N 22 þ E2233 N 33 þ E2223 N 23 þ E2231 N 31 þ E2212 N 12 ÞN 22 þ ðE3311 N 11 þ E3322 N 22 þ E3333 N 33 þ E3323 N 23 þ E3331 N 31 þ E3312 N 12 ÞN 33 þ ðE2311 N 11 þ E2322 N 22 þ E2333 N 33 þ E2323 N 23 þ E2331 N 31 þ E2312 N 12 ÞN 23 þ ðE3111 N 11 þ E3122 N 22 þ E3133 N 33 þ E3123 N 23 þ E3131 N 31 þ E3112 N 12 ÞN 31 þ ðE1211 N 11 þ E1222 N 22 þ E1233 N 33 þ E1223 N 23 þ E1231 N 31 þ E1212 N 12 ÞN 12
Appendix C: Matrix Representation of Tensor Relations
719
which is represented in the Voigt representation as follows: n : E : n ¼ EN 1 N 1 þ EN 2 N 2 þ EN 3 N 3 þ EN 4 N 4 þ EN 5 N 5 þ EN 6 N 6 ¼
6 X
EN i N i
i¼1
ðA:24Þ Analogously, the normal part of the strain rate to the subloading surface is given by ðdn Þij ¼ ðn n : dÞij ¼ nij nrs drs ¼ nij ½n11 d11 þ n22 d22 þ n33 d33 þ 2ðn23 d23 þ n31 d31 þ n12 d12 Þ which is described in terms of the engineering strain rate as ðdn Þij ¼ nij ðn11 d~11 þ n22 d~22 þ n33 d~33 þ n23 d~23 þ n31 d~31 þ n12 d~12 Þ which is further represented in the Voigt form as ðdn Þi ¼ ðn n : dÞi ¼ ni ðn1 d~1 þ n2 d~2 þ n3 d~3 þ n4 d~4 þ n5 d~5 þ n6 d~6 Þ ¼ ni
6 X
nr d~r
r¼1
ðA:25Þ
Appendix D: Euler’s Theorem for Homogeneous Function
The function f ðxÞ of the arbitrary tensor x, which satisfies the following relation for an arbitrary constant c, is called the n-degree homogeneous function. f ðcxÞ ¼ cn f ðxÞ
ðA:26Þ
Differentiating the left and the right sides of this equation separately by c ,we have
@f ðcxÞ @f ðcxÞ @ðcxÞ @f ðcxÞ n1 ¼ ¼ x ¼ nc f ðxÞ @c @cx @c @ðcxÞ
ðA:27Þ
Then, the following Euler’s theorem for the homogeneous function is obtained by setting c ¼ 1 in Eq. (A.27). @f ðxÞ x ¼ nf ðxÞ @x
ðA:28Þ
For the yield function f ðrÞ of the stress tensor r in the homogeneous degree-one, one has @f ðrÞ : r ¼ nf ðrÞ @r
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
ðA:29Þ
721
Appendix E: Outward-Normal Tensor of Surface
Consider the surface f ðTÞ ¼ Fð [ 0Þ (T: arbitrary second-order tensor). The quantity ð@f ðTÞ=@TÞ:dT is regarded as the scalar product of @f ðTÞ=@T and dT in the nine-dimensional space ðT11 ; T22 ; ; T31 ; T13 Þ. Here, it holds that @f ðTÞ : dT ¼ rf ðTÞ : dT @T 8 > < [ 0 for dT directed outward normal to surface ¼ dF ¼ 0 for dT directed tangential to surface > : \0 for dT directed inward normal to surface
df ðTÞ ¼
ðA:30Þ
Therefore, @f ðTÞ=@T designates the outward-normal direction of the surface f ðTÞ ¼ F. Actually, @f ðrÞ=@r designates the outward-normal of the yield surface f ðrÞ ¼ F.
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
723
Appendix F: Relationships of Material Constants in ln v ln p and e ln p Linear Relations
The following relation holds from Eqs. (13.10)1 and (13.22)2 for pe ¼ 0, provided that Eq. (4.141) holds for the elastic volumetric strain, i.e. eev ¼ lnð1 þ eev Þ. ~ j ln
p j p ¼ ln 1 ln p0 1 þ e 0 p0
ðA:31Þ
from which one has ln 1 ~¼ j
j p ln 1 þ e 0 p0 p ln p0
ðA:32Þ
It follows from Eq. (A.32) for infinitesimal deformation under p ffi p0 that j 1 1 þ e0 p j p j p ln 1 ln 1 ln j 1 þ e 0 p0 1 þ e 0 p0 ~ ¼ lim lim j ¼ lim ¼ p 1 p!p0 p!p0 p!p0 1 þ e0 ln p0 p
ðA:33Þ
resulting in ~ffi j
j for p ffi p0 1 þ e0
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
ðA:34Þ
725
Appendix F: Relationships of Material Constants …
726
Further, substituting Eqs. (13.10) 2 and (13.22)3 for pe ¼ 0 into epv ¼ lnð1 þ epv Þ due to Eq. (4.141), it follows that py k j py ~ ~Þ ln ¼ ln 1 ln ðk j py0 1 þ e0 py0 which is reduced to the following relation for py ffi py0 . kj 1 1 þ e 0 py k j py k j py ln 1 ln 1 ln kj 1 þ e0 py0 1 þ e0 py0 ~ ~Þ ¼ ¼ lim ¼ ðk j py 1 py !py0 1 þ e0 ln py0 py
leading to kj ~ ~¼ for py ! py0 kj 1 þ e0
ðA:35Þ
~ and ~ Based on Eqs. (A.33) and (A.35), j k are related to j and k for the infinitesimal deformation as follows: ~ffi j
j ; 1 þ e0
~ kffi
k for p ffi p0 1 þ e0
and
py ffi py0
ðA:36Þ
~ can be calculated from a plenty of test data ~ and k The approximate values of j on k and j accumulated in the past. ~ directly Needless to say, one has to determine the material parameters ~k and j from test data for soils without the data of k and j for the case that an accurate formulation is required. Here, note that the curve fitting of ln v ln p linear relation to test data is easier than the fitting of the e ln p linear relation to test data because real soil behavior is far nearer to the former than the latter.
Appendix G: Derivative in Critical State
Differentiation of Eq. (13.27) under the condition f ðrÞ ¼ const: leads to @gm @gm 0 dp þ d r k k @p @ kr0 k 1 kr0 k d k r0 k ¼ 0 ¼ gðgm Þdp þ pg0 ðgm Þ 2 dp þ pM pM
gðgm Þdp þ pg0 ðgm Þ
from which it follows that k r0 k 0 d kr0 k gðgm Þ þ pM g ðgm Þ gðgm Þ ¼ þ gm ¼M 0 1 0 dp g ðgm Þ M g ðgm Þ
ðA:37Þ
Taking account of d kr0 k=dp ¼ 0 at gm ¼ 1 in Eq. (A.37), one has Eq. (13.30).
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
727
Appendix H: Convexity of Two-Dimensional Curve
When the curve is described by the polar coordinates ðr; hÞ as shown in Fig. A.2, the following relation holds. tan a ¼
rdh dr
ðA:38Þ
where a is the angle measured from the radius vector to the tangent line in the anti-clockwise direction. Eq. (A.38) is rewritten as cot a ¼ where ð
r0 ; r
ðA:39Þ
Þ0 designates the first order differentiation with respect to h.
R(Θ )
y
α α
r (Θ )
P
r (θ ) ϑ Θ
θ
0
T
Θ − [θ − (π / 2 − α )] θ − (π / 2 − α ) x
Fig. A.2 Curve in the polar coordinate (r, q) © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
729
730
Appendix H: Convexity of Two-Dimensional Curve
The equation of the tangent line at ðr; hÞ of the curve r ¼ rðhÞ is described by the following equation by using the current coordinates ðR; HÞ on the tangent line. ð0T ¼ÞRðHÞ cosfH ½h ðp=2 aÞg ¼ rðhÞ cosðp=2 aÞ which is rewritten as RðHÞ sinðH h aÞ ¼ rðhÞ sin a ! !
1 1 ¼ RðHÞ sinðH h aÞ rðhÞ sin a
1 1 1 ¼ cosðH hÞ cot a sinðH hÞ RðHÞ rðhÞ rðhÞ
Substituting Eq. (A.39) to this equation and noting ð1=rÞ cot a ¼ ð1=rÞr0 =r ¼ r 0 =r 2 ¼ ð1=rÞ0 , one has the relation 1 1 1 0 sin # ¼ cos # þ RðHÞ rðhÞ rðhÞ
ðA:40Þ
#Hh
ðA:41Þ
where
Equation (A.40) is rewritten by applying the Taylor expansion to cos # and sin # as 1 1 1 1 0 1 1 2 ¼ ¼ cos # þ 1 # þ sin # ¼ RðHÞ Rðh þ #Þ rðhÞ rðhÞ rðhÞ 2 0 1 1 þ # #3 þ rðhÞ 6 0 1 1 1 1 2 ¼ # ðA:42Þ þ # þ rðhÞ rðhÞ 2 rðhÞ On the other hand, the radius rðHÞ of the curve is described by the Taylor expansion as follows: 1 1 1 1 0 1 1 00 2 ¼ ¼ þ #þ # þ rðHÞ rðh þ #Þ rðhÞ rðhÞ 2 rðhÞ
ðA:43Þ
Appendix H: Convexity of Two-Dimensional Curve
731
Eqs. (A.42) and (A.43) lead to 1 1 1 1 1 00 2 # þ ¼ þ rðHÞ RðHÞ 2 rðhÞ rðhÞ
ðA:44Þ
In order that the curve is convex ðrð#Þ Rð#ÞÞ, the following inequality must hold from Eq. (A.44). 1 1 1 00 2 # þ 0 þ 2 rðhÞ rðhÞ
ðA:45Þ
00 1 1 þ
0 r r
ðA:46Þ
leading to
which is called the convexity condition.
Appendix I: Normal Tensor to Subloading Surface with Anisotropic Damage
The proof of Eq. (15.128) given by Prof. Yuki Yamakawa (Tohoku university) is shown below. First, the following equation holds.
0
@ ðH r0 HÞ
@ðH r0 HÞ
0
0
ðH r0 HÞ 1 1
¼
½H r0 H trðH r0 HÞI ¼
0
0
3
ðH r0 HÞ ðH r0 HÞ
ðA:47Þ
Next, one has @½ðH r0 HÞ0 ij 0
@rkl
@ 1 ½Hip r0pq Hqj ðHrp r0pq Hqr Þdij @r0kl 3 1 ¼ Hip dpk dql Hqj Hrp dpk dql Hqr dij 3 1 ¼ Hik Hlj dij Hrk Hlr 3 1 ¼ ½HH I ðHHÞijkl 3 ¼
ðA:48Þ
and further the following equation holds combining Eqs. (A.47) and (A.48), noting Hij ¼ Hji .
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
733
734
Appendix I: Normal Tensor to Subloading Surface with Anisotropic Damage
!
@ ðH r0 HÞ0 @ðH r0 HÞ0 : @r0 @ðH r0 HÞ0 kl i i h h 1 1 1 0 0
trðH r I ðHHÞ ¼
H HÞI H r HH
ðH r0 HÞ0
ij ijkl 3 3 1 1 1 0 0
¼
ðH r0 HÞ0 ðHip rpq Hqj 3 Hrp rpq Hqr dij ÞðHik Hlj 3 dij Hrk Hlr Þ 1 1 0 0
¼
ðH r0 HÞ0 ðHip rpq Hqj Hik Hlj Hip rpq Hqj 3 dij Hrk Hlr 1 1 Hrp r0pq Hqr dij Hik Hlj þ Hrp r0pq Hqr dij dij Hrk Hlr Þ 3 9 1 1 0 0
¼
ðH r0 HÞ0 ðHip rpq Hqj Hik Hlj 3 Hrp rpq Hqr dij Hik Hlj Þ 1 1 0 0
¼
ðH r0 HÞ0 ðHki Hip rpq Hqj Hjl 3 Hki Hrp rpq Hqr dij Hjl Þ 1 1 0 0
¼
ðH r0 HÞ0 ½Hki ðHip rpq Hqj 3 ðHrp rpq Hqr Þdij Hjl 1 1 0 0
¼
ðH r0 HÞ0 fH½H r H 3 trðH r HÞIHgkl 1 0 0
¼
ðH r0 HÞ0 ½HðH r HÞ Hkl leading to
@ ðH r0 HÞ0 @ðH r0 HÞ0 HðH r0 HÞ0 H
: ¼
ðH r0 HÞ0
@r0 @ðH r0 HÞ0
ðA:49Þ
In addition, one has @r0 @ 1 1 ¼ ½r ðtrrÞI ¼ II I I @r @r 3 3
ðA:50Þ
It follows from Eqs. (A.49) and (A.50).
@ ðH r0 HÞ0 @ ðH r0 HÞ0 @ðH r0 HÞ0 @r0 ¼ : : @r0 @r @r @ðH r0 HÞ0 HðH r0 HÞ0 H 1 ðHðH r0 HÞ0 HÞ0
:ðII I IÞ ¼
¼
0
ðH r0 HÞ
ðH r0 HÞ0
3
ðA:51Þ
Appendix I: Normal Tensor to Subloading Surface with Anisotropic Damage
735
Substituting Eq. (A.51) into pffiffiffiffiffiffiffiffi
pffiffiffiffiffiffiffiffi
0
0 3=2 ðH r0 HÞ0
@ 3=2 ðH r HÞ
=
n¼
@r @r
ðA:52Þ
½HðHr0 HÞ0 H0
n¼
½HðHr0 HÞ0 H0
ðA:53Þ
@
leading to
Appendix J: Tensor Exponential Map for Time-Integration of First-Order Linear Differential Equation
Consider the following relation in terms of an arbitrary second-order tensors T and Z.
TðtÞ ¼ ZTðtÞ
ðA:54Þ
Then, examine the validity of the following candidate of solution for the time-interval Dt ¼ ½tn ; tn þ 1 . Tn þ 1 ¼ expðZDtÞTn
ðA:55Þ
provided that Z and Tn are constant during the time-interval. The time-differentiation of Eq. (A.54) leads noting Eq. (1.344) to
dTn þ 1 ðtÞ d½expðZDtÞTn d expðZDtÞ ¼ ¼ Tn dt dt dt d 1 1 1 ¼ ½I þ ZDt þ ðZDtÞ2 þ ðZDtÞ3 þ Tn ¼ ½Z þ ðZDtÞZ þ ðZDtÞ2 Z þ Tn dt 2! 3! 2! 1 ¼ Z½I þ ðZDtÞ þ ðZDtÞ2 þ Tn ¼ Z expðZDtÞTn ¼ ZTn þ 1 ðtÞ 2!
Tn þ 1 ðtÞ ¼
ðA:56Þ coinciding to Eq. (A.54) and thus the rightness of the candidate in Eq. (A.55) was proven. Equation (A.55) is represented by the generalized mid-point rule as Tn þ 1 ¼ expðZn þ h DtÞTn
ð0 h 1Þ
ðA:57Þ
and is given by the backward implicit scheme for h ¼ 1 as Tn þ 1 ¼ expðZn þ 1 DtÞTn
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
ðA:58Þ
737
738
Appendix J: Tensor Exponential Map for Time-Integration …
Numerical computation of the exponential function of a tensor, as well as its derivative, can be referred to Ortiz et al. (2001), de Souza Neto (2001), Itskov and Aksel (2002), Itskov (2003, 2019), etc.
Appendix K: Eyring Equation
The activation rate at which the segment overcomes the potential barrier by the thermal oscillation was given by following Eyring equation (Eyring 1935). DE v ¼ v0 exp RT
ðA:59Þ
where DE is the activation energy, R is the gas constant, T is the absolute temperature, v0 is the value of v at DE ¼ 0. The activation rate vf in the state that the shear stress rs is applied is given as follows:
DE V rs vf ¼ v0 exp RT
ðA:60Þ
where V is the activation volume. On the other hand, The activation rate vr in the state that the shear stress is applied in the opposite direction is given as
DE þ V rs vr ¼ v0 exp RT
ðA:61Þ
The net activation rate is given from Eqs. (A.60) and (A.61) is given by DE V rs DE þ V rs v0 exp v ¼ vf vr ¼ v0 exp RT RT DE V rs V rs exp exp ¼ v0 exp RT RT RT DE V rs sinh ¼ 2v0 exp RT RT
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
ðA:62Þ
739
740
Appendix K: Eyring Equation
Equation (A.62) leads to Eq. (18.47) by the following replacements, noting Eq. (14.34). v ! C;
DE 1 2v0 exp ! ; RT lv expðuc Rc Cn Þ½1 R=ðcm Rs Þ
V rs ! hR Rs in ðA:63Þ
Appendix L: Computer Programs of Subloading Surface Models
The computer programs for analyses of the elastoplastic deformation of metals and soils based on the subloading surface model and of the friction phenomena based on the subloading-friction model in the forward-Euler method are provided in this appendix. These programs have been composed by the author (Technical Adviser of MSC Software Ltd.; Emeritus Prof., Kyushu Univ.), Masami UENO (Emeritus Prof., Univ. Ryukyus), Takashi OKAYASU (Kyushu Univ.), Toshiyuki OZAKI (Kyushu Electric Eng. Consult. Inc.), Shingo OZAKI (Yokohama National Univ.) and Tatsuya MASE (Tezukayama University). In addition, they are installed as the standard uploaded (ready-made) programs in the commercial FEM software Marc of MSC Software Ltd. after the 2017 version by the leading support of Dr. Motoharu TATEISHI (MSC Software, Ltd.). The readers will understand the formulation of these models clearly by pursuing each line in the programs. In addition, the readers will be able to perform readily the FE analyses of the boundary-value problems involving the frictional boundary by installing these programs into the FE program (through user-subroutine in case of commercial software).
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
741
742
Appendix L: Computer Programs of Subloading Surface Models
(a) Subloading surface model for metals (i) Isotropic and kinematic hardening version composed by Koichi Hashiguchi and Masami Ueno.
Appendix L: Computer Programs of Subloading Surface Models
743
744
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
745
746
Appendix L: Computer Programs of Subloading Surface Models
K
Appendix L: Computer Programs of Subloading Surface Models
747
748
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
749
750
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
751
752
Appendix L: Computer Programs of Subloading Surface Models
(ii) Simplified version without tangential-inelastic strain rate and isotropic hardening stagnation composed by Koichi Hashiguchi and Masami Ueno.
Appendix L: Computer Programs of Subloading Surface Models
753
754
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
755
) ) ) σ' : cˆ ' =σ 1' cˆ1' + σ2' cˆ'2 +σ)3' cˆ'3 ) ) ) + 2 (σ 4' cˆ'4 + σ5' cˆ'5 + σ6' cˆ'6 )
756
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
757
758
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
759
)
K
760
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
761
762
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
763
c
764
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
765
766
Appendix L: Computer Programs of Subloading Surface Models
(iii) Full version composed by Koichi Hashiguchi, Masami Ueno and Toshiyuki Ozaki.
Appendix L: Computer Programs of Subloading Surface Models
767
768
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
769
) ) ) σ' : cˆ ' =σ 1' cˆ1' +σ 2' cˆ'2 + σ)3' cˆ'3 ) ) ) + 2 (σ 4' cˆ'4 + σ5' cˆ'5 + σ6' cˆ'6 )
770
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
771
772
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
773
774
Appendix L: Computer Programs of Subloading Surface Models
[1−exp(−n R)]e
ITCOM = TR / (2.0d0 * SG) do i = 1, 6 NN(i) = SALBD_M(i) / SALBD end do
n = σ'
|| σ' ||
Appendix L: Computer Programs of Subloading Surface Models
775
776
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
777
778
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
779
780
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
781
782
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
(b) Initial subloading surface model for soils composed by Koichi Hashiguchi, Masami Ueno and Tatsuya Mase.
783
784
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
785
786
Appendix L: Computer Programs of Subloading Surface Models
(nm / 3)
Appendix L: Computer Programs of Subloading Surface Models
787
Lam Kap d
K
788
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
789
Ed
790
Appendix L: Computer Programs of Subloading Surface Models
Ed ) E
d
d n
Lam Kap
Appendix L: Computer Programs of Subloading Surface Models
791
792
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
(c) Subloading-overstress model with full version for metals composed by Koichi Hashiguchi, Masami Ueno and Takuya Anjiki
793
794
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
795
796
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
797
798
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
799
800
Appendix L: Computer Programs of Subloading Surface Models
~
~ ~
~
~
~ ~
~ ~
~
Appendix L: Computer Programs of Subloading Surface Models
801
802
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
803
804
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
805
806
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
807
808
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
809
810
Appendix L: Computer Programs of Subloading Surface Models
Appendix L: Computer Programs of Subloading Surface Models
(d) Subloading-friction model with saturation of tangential contact stress composed by Koichi Hashiguchi and Masami Ueno
811
812
Appendix L: Computer Programs of Subloading Surface Models
MYU MYU MYU MYU MYU
Appendix L: Computer Programs of Subloading Surface Models
813
814
Appendix L: Computer Programs of Subloading Surface Models
DMYU
MYU MYU MYU
MYU
MYU
DMYU
MYU
Appendix L: Computer Programs of Subloading Surface Models
MYUk MYUs
MYUk MYUs
815
816
Appendix L: Computer Programs of Subloading Surface Models
References
Books on Solid Mechanics and Tensor Analysis Asaro RJ, Lubarda V (2006) Mechanics of Solids and Materials. Cambridge Univ. Press, Cambridge, UK Basar Y, Weichert D (2000) Nonlinear Continuum Mechanics of Solids: Fundamental Mathematical and Physical Concept. Springer Bazant ZP, Cedolin L (1991) Stability of Structures: Elastic, Inelastic, Fracture and Damage Theories. Oxford Univ. Press, New York Becker, E. and Burger, W. (1975): Kontinuumsmechanik, B. G. Teubner, Stuttgart. Belytschko T, Liu WK, Moran B, Elkhodary KI (2014) Nonlinear Finite Elements for Continua and Structures, 2nd edn. John-Wiley, New York Bergstrom J (2015) Mechanics of Solid Polymers. Elsevier Bertram A (2008) Elasticity and Plasticity of Large Deformations. Springer-Verlag, Berlin Bingham EC (1922) Fluidity and Plasticity. McGraw-Hill, New York Besseling JF, Van der Giessen E (1994) Mathematica Modelling of Inelastic Deformation. Chapman & Hall, London Biot MA (1956) Mechanics of Incremental Deformations. John-Wiley, New York Bonet J, Wood RD (1997) Nonlinear Continuum Mechanics for Finite Element Analysis. Cambridge Univ. Press, Cambridge Borja RI (2013) Plasticity: Modeling & Computation. Springer-Verlag, Heidelberg Bowden FP, Tabor D (1958) The Friction and Lubrication of Solids. Clarendon Press, Oxford, UK Bowen RM, Wang C-C (2008) Introduction to Vectors & Tensors, 2nd edn. Dover Publ. Inc., New York Chadwick P (1999) Continuum Mechanics: Concise Theory and Problems. Dover Publ. Inc., New York Chakrabarty J (1987) Theory of Plasticity. McGraw-Hill, New York Ciarlet PG (1988) Mathematical elasticity I: Three-dimensional Elasticity. North-Holland Clayton JD (2011) Nonlinear Mechanics of Crystals. Springer Conway JB (1967) Numerical Method for Creep and Rupture Analysis. Publ, Gordon and Breach Sci Cristescu N (1967) Dynamic Plasticity. North-Holland, Amsterdam Chung TJ (2007) General Continuum Mechanics. Cambridge Univ. Press, Cambridge, UK De Borst R, Crisfield MA, Remmers JJC, Verhoosed CV (2012) Nonlinear Finite Element Analysis of Solids and Structures, 2nd edn. John-Wiley, Chichester, UK Desai C, Siriwardane HJ (1984) Constitutive Laws for Engineering Materials with Emphasis on Geomaterials. Prentice-Hall Inc., New York
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
817
818
References
de Souza Neto EA, Perić D, Owen DJR (2008) Computational Methods for Plasticity, John-Wiley, Chichester, UK Dimitrienko YI (2010) Tensor Analysis and Nonlinear Tensor Functions. Kluwer Academic Publ, Dordrecht, Netherland Dimitrienko YI (2011) Nonlinear Continuum Mechanics and Large Inelastic Deformations. Springer-Verlag, Dordrecht, Netherland Doghri I (2000) Mechanics of Deformable Solids. Springer-Verlag, Berlin Duvaut G, Lions JL (1972) Les Inequations en Mechanique et en Physique, Dunod, Paris (Duvaut G, Lions JL (1976) Inequalities in Mechanics and Physics. Springer-Verlag, New York) Dvorkin EN, Goldschmit MB (2006) Nonlinear Continua. Springer-Verlag, Berlin Ellyin F (1997) Fracture Damage, Crack Growth and Life Prediction. Chapman & Hall, London Ericksen JL (1991) Introduction to the Thermodynamics of Solids. Chapman and Hall, London Eringen AC (1962) Nonlinear Theory of Continuous Media. McGraw-Hill, New York Eringen AC (1967) Mechanics of Continua. John-Wiley, New York Fischer-Cripps AC (2000) Introduction to Contact Mechanics. Springer-Verlag, New York Flugge W (1972) Tensor Analysis and Continuum Mechanics. Springer-Verlag, Berlin Fung YC (1965) Foundations of Solid Mechanics. Prentice-Hall, Englewood Cliffs, New Jersey Fung YC (1969) A First Course in Continuum Mechanics. Prentice-Hall, Englewood Cliffs, New Jersey Gambin W (2001) Plasticity and Textures. Kluwer Academic Publ, Dordrecht, Netherland Golub GH, Loan CFV (2013) Matrix Computations, 4th edn. John Hopkins University Press, Baltimore, Maryland Gurtin ME (1981) An Introduction to Continuum Mechanics. Academic Press, New York Gurtin ME, Fried E, Anand L (2010) The Mechanics and Thermodynamics of Continua. Cambridge University Press, New York Han W, Reddy BD (1999) Plasticity – Mathematical Theory and Numerical Analysis. Springer-Verlag, New York Hashiguchi K (2009) Elastoplasticity theory, 1st edn. Lecture Note in Applied and Computational Mechanics. Springer, Heidelberg Hashiguchi K (2013) Elastoplasticity Theory, Second edition, Lecture Note in Appl. Compt. Mech., Springer-Verlag, Heidelberg Hashiguchi K (2017) Foundations of Elastoplasticity: Subloading Surface Model. Springer Hashiguchi K (2020) Nonlinear Continuum Mechanics for Finite Elastoplasticity: Multiplicative Decomposition with Subloading Surface Model. Elsevier Hashiguchi K, Yamakawa Y (2012) Introduction to Finite Strain Theory for Continuum Elasto-plasticity. Wiley series in computational mechanics. Wiley, Chichester, UK Haupt P (2000) Continuum Mechanics and Theory of Materials. Springer-Verlag, Wien Havner KS (1992) Finite Plastic Deformation of Crystalline Solids. Cambridge University Press, Cambridge Hill R (1950) The Mathematical Theory of Plasticity. Oxford Univ. Press, London Hinton E, Owen DRJ (1980) Finite Elements in Plasticity: Theory and Practice. Pineridge Press, Swansea, UK Hisada T (1992) Tensor Analysis for Nonlinear Finite Element Method. Maruzen Publ, Inc, Tokyo (in Japanese) Hisada T, Noguchi H (1995) Foundations and Applications in Nonlinear Finite Element Method. Maruzen Publ, Tokyo (in Japanese) Holzapfel GA (2000) Nonlinear Solid Mechanics: A Continuum Approach for Engineering. John Wiley & Sons, Ltd Hosford WF (2009) Mechanical Behavior of Solids. Cambridge University Press Houlsby GT, Puzrin AM (2006) Principles of Hyperelasticity. Springer-Verlag, Heidelberg, An Approach to Plasticity Theory Based on Thermodynamic Principles
References
819
Huang Y (1991) A User-material Subroutine Incorporating Single Crystal Plasticity in the Abaqus Finite Element Program. Harvard univ, Division Appl. Sci. Hunter SC (1983) Mechanics of Continuous Media. Ellis Horwood, Chichester, UK Ibrahimbegovic A (2009) Nonlinear Solid Mechanics: Theoretical Formulations and Finite Element Solution Methods. Springer-Verlag, Dordrecht, Netherland Ilyushin AA (1963) Plasticity—Foundation of the General Mathematical Theory, izdatielistbo Akademii Nauk CCCR (Publisher of the Russian Academy of Sciences), Moscow Ishikawa S (2015) How to Learn Finite Element Method for Rubbers: Basic Theories and Practice for Hyperelasticity. Nikkan Kogyo Shinbun, Ltd. (in Japanese) Itskov M (2015) Tensor Algebra and Tensor Analysis for Engineers with Application to Continuum Mechanics, 5th edn. Springer-Verlag, Dordrecht, Netherland Jaunzemis W (1967) Continuum Mechanics. The Macmillan, New York Jeffreys H (1931) Cartesian Tensors. Cambridge Univ. Press, London Kachanov LM (1971) Foundation of Theory of Plasticity. North-Holland, Amsterdam Khan AS, Huang S (1995) Continuum Theory of Plasticity. John-Wiley, New York Kyoya T (2008) Note on Continuum Mechanics, Japan. Assoc. for Nonlinear CAE, Tokyo. (in Japanese) Leigh DC (1968) Nonlinear Continuum Mechanics: An Introduction to the Continuum Physics and Mechanical Theory of the Nonlinear Mechanical Behavior of Materials. McGraw-Hill, New York Lemaitre JA (1996) A course on Damage Mechanics, 2nd edn. Springer, Heidelberg Lemaitre JA (2001) Handbook of Materials Behavior Models (Ed.). Academic Press, San Diego Lemaitre JA, Chaboche J-L (1990) Mechanics of Solid Materials. Cambridge University Press, Cambridge Lemaitre JA, Desmoral R (2005) Engineering Damage Mechanics. Springer, Heidelberg Lubarda VA (2002) Elastoplasticity Theory. CRC Press, Boca Ranton, Florida Lubliner J (1990) Plasticity theory. Macmillan, New York Malvern LE (1969) Introduction to the Mechanics of a Continuous Medium. Prentice-Hall, Englewood Cliffs, New Jersey Martin JB (1975) Plasticity: Foundation and General Results. MIT Press, Cambridge, Massachusetts Marsden JE, Hughes TJR (1983) Mathematical Foundation of Elasticity. Prentice-Hall, Englewood Cliffs, New Jersey Maugin GA (1992) The Thermomechanics of Plasticity and Fracture. Cambridge Univ. Press, Cambridge Murakami S (2012) Continuum Damage Mechanics: A Continuum Mechanics Approach to the Analysis of Damage and Fracture. Springer, Dordrecht, Netherland Nadai A (1963) Theory of Flow and Fracture of Solids. McGraw-Hill, New York Negahban M (2012) The Mechanical and Thermodynamical Theory of Plasticity. CRC Press, Boca Raton, Florida Nemat-Nasser S (2004) Plasticity: A Treatise on Finite Deformation of Heterogeneous Inelastic Materials. Cambridge Univ. Press, New York, NY Norton FH (1929) Creep of Steel at High Temperature. McGraw-Hill, New York, NY Oden JT (2011) An Introduction to Mathematical Modeling: A Course in Mechanics. Wiley Series in Continuum Mechanics, John-Wiley, New York. Odqvist FKG, Hult JAH (1962) Kriechfestigkeit Metallischer Werkstoffe. Springer-Verlag, Berlin Odqvist FKG (1966) Mathematical Theory of Creep and Creep Rupture. Oxford Univ. Press, London Ogden RW (1984) Non-linear Elastic Deformations. Ellis-Horwood, Chichester, UK Ottosen NS, Ristinmaa M (2005) The Mechanics of Constitutive Modeling. Elsevier, Amsterdam Penny RK, Marriott DL (1971) Design for Creep. McGraw-Hill
820
References
Prager W (1961) Introduction to Mechanics of Continua. Ginn & Comp, Boston, MA Press WH, Teukolsky SA, Vetterling WT, Flannery BP (1992) Numerical Recipies, The Art of Scientific Computing. Cambridge Univ. Press, New York Rabotnov YN (1969) Creep Problems in Structural Members. North-Holland, Amsterdam Roscoe KH, Burland JB (1968) On the Generalized Stress-Strain Behaviour of ‘wet’ Clay. Engineering plasticity Cambridge University Press, Cambridge, pp 535–608 Saada AS, Bianchini G (1989) Proceedings of International Workshop on Constitutive Equations for Granular Non-Cohesive Soils. Balkema, Amsterdam Sawczuk, A. (ed.) (1974) Foundations of Plasticity and Problems of Plasticity (Proc. Int. Symp. Foundation of Plasticity), Noordhoff Int. Publ., Leyden, Netherland. Sawczuk A, Bianchi G (eds) (1985) Plasticity Today: Modelling, Methods and Applications. Elsevier Appl. Sci. Publ, London Schofield AN, Wroth CP (1968) Critical State Soil Mechanics. McGraw-Hill, London Sedov LI (1966) Foundation of the Non-Linear Mechanics of Continua. Pergamon, Oxford, UK Simo JC (1998) Numerical Analysis and Simulation of Plasticity. In: Ciarlet PG, Lions JL (eds) Handbook of Numerical Analysis, vol 6, Part 3 (Numerical Methods for Solids), Elsevier, Amsterdam Simo JC, Hughes TJR (1998) Computational Inelasticity. Springer, New York Skrzypek JJ, Hetnarski RB (1993) Plasticity and Creep: Theory, Example and Problems. CRC Press, London Sokolnikoff IS (1964) Tensor Analysis: Theory and Applications to Geometry and Mechanics of Continua. John-Wiley, New York Spencer AJM (1980) Continuum Mechanics. Longman Sci. Tech. Publ, London Steinmann P (2015) Geometrical Foundation of Continuum Mechanics. Springer Tanahashi T (1985) Mechanics of Continua. Rikoh Tosho Publ, Tokyo (in Japanese) Taylor DW (1948) Fundamentals of Soil Mechanics. Wiley, Chichester, UK Truesdell C (1977) A First Course in Rational Continuum Mechanics, vol 1. Academic Press, New York, General Concepts Truesdell C, Noll W (1965) The Nonlinear Field Theories of Mechanics, encyclopedia of physics. In: Flugge S (ed), vol III/3. Springer-Verlag, Berlin Truesdell C, Toupin R (1960) The classical field theories. In: Flugge S (ed) Encyclopedia of physics, vol III/1. Springer-Verlag, Berlin Valanis KC (1972) Irreversible thermodynamics of continuous media, internal variable theory. CISM Courses and Lectures No. 77, International Center for Mechanical Sciences, Springer, Wien Vardoulakis I, Sulem J (1995) Bifurcation Analysis in Geomechanics. Blackie Academic & Profess, London Voyiadjis GZ, Kattan PI (2005) Damage Mechanic (Mechanical Engineering). CRC Press, New York Ward JM, Sweeney J (2013) Mechanical Properties of Solid Polymers. Wiley, NewYork Wriggers P (2003) Computational Contact Mechanics. Wiley, Hoboken, NJ Wu H-C (2004) Continuum Mechanics and Plasticity. Chapman & Hall/CRC, New York Zhong Z-H (1993) Finite Element Procedures for Contact-Impact Problems. Oxford University Press, London Ziegler H (1983) An Introduction to Thermomechanics, 2nd edn. North-Holland, Amsterdam Zienkiewicz OC (1977) The Finite Element Method, 3rd edn. McGraw-Hill, London Zyczkowski M (1981) Combined Loading in the Theory of Plasticity, PWN. Polish Sci. Publ, Warsaw
References
821
Research Articles Abdel-Karim M, Ohno N (2000) Kinematic hardening model suitable for ratchettig with steady-state. Int J Plast 16:225–240 Aub-Rub RKA, Kim S-M (2010) Computational applications of a coupled plasticity-damage constitutive model for simulating plan concrete fracture. Eng Fract Mech 77:1577–1603 Aifantis EC (1984) On the microstructural origin of certain inelastic models. J Eng Mater Tech (ASME) 106:326–330 Alonso EE, Gens A, Josa A (1990) A constitutive model for partially saturated soils. Geotechnique 40:405–430 Anand L (1993) A constitutive model for interface friction. Comput Mech 12:197–213 Anand L, Gurtin ME (2003) A theory of amorphous solids undergoing large deformations, with application to polymathic glass. Int J Solids Struct 40:1465–1487 Anand L, Kothari M (1996) A computational procedure for rate-independent crystal plasticity. J Mech Phys Solids 44:525–558 Andrade JE, Borja RI (2006) Capturing strain localization in dense sands with random density. Int J Numer Meth Eng 67:1531–1564 Anjiki T, Hashiguchi K (2020) Viscoplastic deformation analysis for spherical graphite cast iron by subloading-overstress model. Trans Jpn Soc Mech Eng 86(980) No. 20–000158. https://doi. org/10.1299/transjsme (in Japanese) Anjiki T, Hashiguchi K (2021) Extended overstress model and its implicit stress integration algorithm: formulations, experiments and simulations. Int J Numer Meth Eng 123:1–13 Anjiki T, Oka M, Hashiguchi K (2016) Elastoplastic analysis by complete implicit stress-update algorithm based on the extended subloading surface model. Trans Japan Soc Mech Eng 82 (839), No 16-00029. https://doi.org/10.1299/transjsme.16-00029 Anjiki T, Oka M, Hashiguchi K (2019) Complete implicit stress integration algorithm with extended subloading surface model for elastoplastic deformation analysis. Int J Numer Meth Eng 121:945–966 Argyris JH (1965) Elasto-plastic matrix analysis of three dimensional continua. J. Roy. Aeronaut. Soc. 69:231–262 Argyris JH, Faust G, Szimma J, Warnke EP, William KJ (1973) Recent developments in the finite element analysis of PCRV. In: Proceedings of 2nd international conference SMIRT. Berlin Armstrong PJ, Frederick CO (1966) A mathematical representation of the multiaxial Bauschinger effect, CEGB Report RD/B/N 731 (or in Materials at High Temperature, 24, 1–26 (2007)) Asaoka A, Nakano M, Noda T (1997) Soil-water coupled behaviour of heavily overconsolidated clay near/at critical state. Soils Found 37(1):13–28 Asaoka A, Nakano M, Noda T (2000) Superloading yield surface concept for highly structured soil behavior. Soils Found 40(2):99–110 Asaoka A, Noda T, Yamada E, Kaneda I, Nakano M (2002) An elasto-plastic description of two distinct volume change mechanics of soils. Soils Found 42(5):99–110 Asaro RJ (1983) Micromechanics of crystals and polycrystals. In: Advances in applied mechanics, vol 23. Academic Press Asaro RJ, Needleman A (1985) Texture development and strain hardening in rate dependent polycrystals. Acta Metall 33:923–953 Asaro RJ, Rice JR (1977) Strain localization in ductile single crystals. J Mech Phys Solids 25:309– 338 Avrami M (1939) Kinetics of phase change I: general theory. J Chem Phys 7:1103–1112 Bardet JP (1990) Lode dependences for isotropic pressure-sensitive elastoplastic materials. J Appl Mech (ASME) 57:498–506 Barlat F, Yoon JW, Cazacu O (2007) On linear transformations of stress tensors for the description of plastic anisotropy. Int J Plasticity 23:876–896
822
References
Bassani JL, Wu TY (1991) Latent hardening in single crystals II: Theory analytical characterization and predictions. Proc. Royal Soc. London A 435:21–41 Batdorf SB, Budiansky B (1949) A mathematical theory of plasticity based on the concept of slip, NACA TC1871, 1–31 Baumberger T, Heslot F, Perrin B (1994) Crossover from creep to inertial motion in friction dynamics. Nature 30:544–546 Bay N, Wanheim T (1976) Real area of contact and friction stresses at high pressure sliding contact. Wear 38:201–209 Betton J (1986) Application of tensor functions to the formulation of constitutive equations involving damage and initial anisotropy. Eng Fract Mech 25:573–584 Bishop AW, Webb DL, Lewin PI (1965) Undisturbed samples of London clay from the Ashford common shaft: strength-effective stress relationships. Geotechnique 15:1–31 Bland DR (1957) The associated flow rule of plasticity. J Mech Phys Solids 6:71–78 Bodner SR, Partom Y (1975) Constitutive equations for elastic viscoplastic strain hardening. J Appl Mech (ASME) 42:385–389 Borja RI (2004) Cam-Clay plasticity. Part V: a mathematical framework for three-phase deformation and strain localization analyses of partially saturated porous media. Comput Meth Appl Mech Eng 193:5301–5338 Borja RI, Sama KM, Sanz PF (2003) On the numerical integration of three-invariant elastoplastic constitutive models. Comput Meth Appl Mech Eng 192:1227–1258 Boyce MC, Parks DM, Argon AS (1988) Large inelastic deformation of glass polymers. Part I: rate dependent constitutive model. Mech Mater 7:15–33 Brepols T, Vladimirov IN, Reese S (2014) Numerical comparison of isotropic hypo- and hyperelastic-based plasticity models with application to industrial forming processes. Int. J. Plasticity 63:18–48 Brockley CA, Davis HR (1968) The time-dependence of static friction. J Lubr Tech (ASME) 90:35–41 Budiansky B (1959) A reassessment of deformation theories of plasticity. J. Appl. Mech. (ASME) 20:259–264 Bureau L, Baumberger T, Caroli C, Ronsin O (2001) Low-velocity friction between macroscopic solids. C.R. Acad Sci Paris Ser IV Diff Faces Tribol 2:699–707 Burland JB (1965) The yielding and dilatation of clay, Correspondence. Geotechnique 15:211–214 Butterfield R (1979) A natural compression law for soils (an advance on e-log p’). Geotechnique 29:469–480 Callari C, Auricchio F, Sacco E (1998) A finite-strain Cam-clay model in the framework of multiplicative elasto-plasticity. Int J Plasticity 14:1155–1187 Cao P, Lin X, Park HS (2014) Strain-rate and temperature dependence of yield stress of amorphous solids via a self-learning metabasin escape algorithm. J Mech Phys Solids 68:239– 250 Carlson DE, Hoger A (1986) The derivative of a tensor-valued function of a tensor. Quart. Appl. Math. 406:409–423 Casey J, Naghdi PM (1980) Remarks on the use of the decomposition F ¼ F e F p in plasticity. J Appl Mech (ASME) 47:672–675 Castro G (1969) Liquefaction of Sands, Phd Thesis, Harvard Soil Mech. Series 81 Chaboche, JL (1988): Continuum damage mechanics, Pert I General concept; Part II Damage growth, crack initiation, and crack growth, J. Appl. Mech. (ASME), 55, 59-72 Chaboche JL (1989) Constitutive equations for cyclic plasticity and cyclic viscoplasticity. Int J Plast 5:247–302 Chaboche JL (1991) On some modifications of kinematic hardening to improve the description of ratcheting effects. Int J Plast 7:661–678 Chaboche JL (1997) Thermodynamic formulation of constitutive equations and application to the viscoplasticity and viscoelasticity of metals and polymers. Int J Solids Struct 38:2239–2254
References
823
Chaboche JL (2008) A review of some plasticity and viscoplasticity constitutive theories. Int J Plast 24:1642–1693 Chaboche JL, Dang-Van K, Cordier G (1979) Modelization of the strain memory effect on the cyclic hardening of 316 stainless steel. In: Transactions on 5th International Conference on SMiRT, Berlin, Division L., Paper No. L. 11/3 Chaboche JL, Gaubert A, Kanoute P, Longuet A, Azzouz F, Maziere M (2013) Viscoplastic constitutive equations of combustion chamber materials including cyclic hardening and dynamic strain aging. Int J Plast 46:1–22 Chaboche JL, Rousselier G (1983) On the plastic and viscoplastic constitutive equations, Parts I and II. J Pressure Vessel Tech (ASME) 165:153–164 Chen Z, Bong H-J, Li D, Wagoner RH (2016) The elastic-plastic transition of metals. Int J Plast 83:178–201 Cheng J-H, Kikuchi N (1985) An incremental constitutive relation of uniaxial contact friction for large deformation analysis. J Appl Mech (ASME) 52:639–648 Christoffersen J, Hutchinson JW (1979) A class of phenomenological corner theories of plasticity. J Mech Phys Solids 27:465–487 Chowdhury EQ, Nakai T, Tawada M, Yamada S (1999) A model for clay using modified stress under various loading conditions with the application of subloading concept. Soils Found 39 (6):103–116 Chu CC, Needleman A (1980) Void nucleation effects in biaxially stretched sheets. J Eng Technol ASME 102:249–256 Cicekli U, Voyiadjis GZ, Al-Rub RKA (2007) A plasticity and anisotropic model for plain concrete. Int J Plasticity 23:1874–1900 Cleja-Tigoiu S, Soos E (1990) Elastoviscoplastic models with relaxed configurations and internal state variables. Appl Mech Reviews (ASME) 43:131-151 p e Clifton RJ (1972) On the equivalence of F ¼ Fe Fp and F ¼ F F . J Appl Mech (ASMS) 14:703– 717 Collins IF, Hilder T (2002) A theoretical framework for constructing elastic/plastic constitutive models of triaxial tests. Int J Nemer Anal Meth Geomech 26:1313–1347 Conrad H (1964) Thermally activated deformation of metals. J Metals 16:582 Coombs WM, Crouch RS (2011) Algorithmic issues for three-invariant hyperplastic critical state models. Comput Meth Appl Mech Eng 200:2297–2318 Coombs WM, Crouch RS, Augarde CE (2013) A unique Critical State two-surface hyperplasticity model for fine-grained particulate media. J Mech Phys Solis 61:175–189 Cordebois JP, Sidoroff F (1982a) Damage induced elastic anisotropy. In: Boehler JP (ed) Mechanical behavior of anisotropic solids. Martinuus Nijhoff Publ., pp 761–774 Cordebois JP, Sidoroff F (1982b) Endommagement anisotrope en elasticite et plasticite. J de Mech Theor et Appl Numero Spec, pp 45–60 Cosserat E, Cosserat F (1909) Theorie des corps deformation, Traite de Physique, transl Davaux E. In: Chwolson OD (ed), 2nd edn, vol 2. Paris, pp 953–1173 Cotter BA, Rivlin RS (1955) Tensors associated with time-dependent stresses. Quart. Appl. Math. 13:177–182 Covindjee S, Simo JC (1992) Mullins effect and the strain amplitude dependence of the storage modulus. Int J Solids Struct 29:1737–1751 Cundall P, Board M (1988) A microcomputer program for modeling large-strain plasticity problems. In: Prepare for the 6th International conference on numerical methods in geomechanics. Innsbruck, Austria, pp 2101–2108 Curnier A (1984) A theory of friction. Int J Solids Struct 20:637–647 Dafalias YF (1983) Corotational rates for kinematic hardening at large plastic deformations. J Appl Mech (ASME) 50:561–565 Dafalias YF (1984) The plastic spin concept and a simple illustration of its role in finite plastic transformation. Mech Mater 3:223–233 Dafalias YF (1985a) The plastic spin. J Appl Mech (ASME) 52:865–871
824
References
Dafalias YF (1985b) A missing link in the macroscopic constitutive formulation of large plastic deformations. In: Sawczuk A, Bianchi G (eds) Plasticity today, international symposium recent trends and results in plasticity. Elsevier Publ., pp 135–151 Dafalias YF (1986) Bounding surface plasticity. I: Mathematical foundation and hypoplasticity. J Eng Mech (ASCE) 112:966–987 Dafalias YF (1998) Plastic spin: Necessity or redundancy ? Int. J. Plasticity 14:909–931 Dafalias YF (2011) Finite elastic-plastic deformations: beyond the plastic spin. Theor Appl Mech 38:321–345 Dafalias YF, Feigenbaum HP (2011) Biaxial ratchetting with novel variations of kinematic hardening. Int. J. Plasticity 27:479–491 Dafalias YF, Herrmann LR (1980) A bounding surface soil plasticity model, Proc. Int. Symp. Soils Cyclic Trans. Load., Swansea, pp. 335–345 Dafalias YF, Kourousis KI, Saridis GJ (2008) Multiplicative AF kinematic hardening in plasticity. Int J Solids Struct 45:2861–2880 Dafalias YF, Popov EP (1975) A model of nonlinearly hardening materials for complex loading. Acta Mech 23:173–192 Dafalias YF, Popov EP (1976) Plastic internal variables formalism of cyclic plasticity. J Appl Mech (ASME) 43:645–651 Dafalias YF, Popov EP (1977) Cyclic loading for materials with a vanishing elastic domain. Nucl. Eng. Design 41:293–302 Dafalias YF, Talebat M (2016) SANISAND-Z: zero elastic range plasticity model. Geotechnique 66:999–1013 Darabi MK, Abu Al-Rub RK, Masad EA., Huang CW, Little DA. (2012) A modified viscoplastic model to predict the permanent deformation of asphaltic materials under cyclic-compression loading at high temperatures. Int J Plasticity 35:100–134 Darrieulat M, Piot D (1996) A method of generalized analytical yield surfaces of crystalline materials. Int J Plasticity 12:575–610 Dashner PA (1986) Invariance considerations in large strain elasto-plasticity. J Appl Mech (ASME) 53:55–60 de Borst R, Groen AE (1999) Towards efficient and robust elements for 3-D plasticity. Comput Struct 70:23–34 de Souza Neto EA (2001) The exact derivative of the exponential of an unsymmetric tensor. Comp Meth Appl Mech Eng 190:2377–2383 de Souza Neto EA, Peric D, Owen DR (1998) Continuum modelling and numerical simulation of material damage at finite strains. Comp Meth Appl Mech Eng. 5:341–384 del Peiro G (1979) Some properties of the set of fourth-order tensors, with application to elasticity. J. Elasticity 9:245–261 Denis S, Gautier e, Simon A, Beck G (1985) Stress-phase-transformation basic principles, modelling, and calculation of internal stresses. Mater Sci Tech 1:805–814 Derjaguin BV, Push VE, Tolstoi DM (1957) A theory of stick-slipping of solids. In: Proceedings of the Conference on Lubrication and Wear, Institute Mechanical Engineering, London, pp 257–268 Delobelle P, Robinet P, Bocher L (1995) Experimental study and phenomenological modelization of ratchet under uniaxial and biaxial loading on austenitic stainless steel. Int J Plasticity 11:295–330 Denis S, Gautier e, Simon A, Beck G (1985) Stress-phase-transformation basic principles, modelling, and calculation of internal stresses. Mater Sci Tech 1:805–814 Dettmer W, Reese S (2004) On the theoretical and numerical modelling of Armstrong-Frederic kinematic hardening in the finite strain regime. Comput Meth Appl Mech Eng 193:87–116 Dienes JK (1979) On the analysis of rotation and stress rate in deforming bodies. Acta Mech 32:217–232 Dieterich JH (1978) Time-dependent friction and the mechanism of stick-slip, Pure Appl. Geophys 116:790–806
References
825
Diteterih JH (1979) Modeling of rock friction 1. experimental results and constitutive equations. J Geophys Res 84:2161–2168 Dokos SJ (1946) Sliding friction under extreme pressure—I. Trans ASME 68:A148-156 Drucker DC (1951) A more fundamental approach to plastic stress-strain relations. Proc 1st U.S. National Congr Appl Mech (ASME) 1:487–491 Drucker DC (1988) Conventional and unconventional plastic response and representation. Appl Mech Rev (ASME) 41:151–167 Drucker DC, Gibson RE, Henkel DJ (1957) Soil mechanics and workhardening theories of plasticity. Trans Am Soc Civil Eng 122:338–346 Drucker DC, Prager W (1952) Soil mechanics and plastic analysis or limit design. Quart Appl Math 10:157–165 Dunkin JE, Kim DE (1996) Measurement of static friction coefficient between flat surfaces. Wear 193:186–192 Duszek MK, Perzyna P (1991) On combined isotropic and kinematic hardening effects in plastic flow process. Int. J. Plasticity 9:351–363 Eckart C (1948) The thermodynamics of irreversible process, IV: the theory of elasticity and anelasticity. Phys Rev 73:373–380 Eidel B, Gruttmann F (2003) Elastoplastic orthotropy at finite strains: multiplicative formulation and numerical implementation. Compt Materials Sci 28:732–742 Ellyin F (1989) An anisotropic hardening rule for elastoplastic solids based on experimental observations. J Appl Mech (ASME) 56:499–507 Ellyin F, Xia Z (1989) A rate-independent constitutive model for transient non-proportional loading. J Mech Phys Solids 37:71–91 Eshelby JD (1957) The determination of the elastic field of an ellipsoidal inclusion and related problems. Proc R Soc Lond A 241:376–396 Eyring H (1935) The activated complex in chemical reactions. J Chem Phys 3:107–115 Eyring H (1936) Viscocity, plasticity and diffusion as examples of absolute reaction rate. J Chem Phys 4:283–291 Fardshisheh F, Onat ET (1974) Representations of elastoplastic behavior by means of state variables. In: Sawczuk A (ed) Problems of plasticity, pp 89–115 Feigenbaum HP, Dafalias YF (2007) Directional distortional hardening in metal plasticity within thermodynamics. Int J Solids Struct 44:7526–7542 Feigenbaum HP, Dafalias YF (2014) Directional distortional hardening at large plastic deformation. Int J Solids Struct 51:3904–3918 Ferrero JF, Barrau JJ (1997) Study of dry friction under small displacements and near-zero sliding velocity. Wear 209:322–327 Fincato R, Tsutsumi S (2017) Closest-point projection method for the extended subloading surface model. Acta Mech 228:4213–4233 Fincato R, Tsutsumi S (2018) A return mapping algorithm for elastoplastic and ductile damage constitutive equations using the subloading surface method. Int J Numer Meth Eng 113:1729– 1754 Fincato R, Tsutsumi S (2019a) Numerical modeling of the evolution of ductile damage under proportional and non-proportional loading. Int J Solids Struct 160:247–264 Fincato R, Tsutsumi S (2019b) Closest projection method for extended subloading surface model. Acta Mech 228:4213–4233 Fincato R, Tsutsumi S (2020a) An overstress elasto-viscoplasticity model for high/low cyclic strain rates loading conditions: Part I—Formulation and computational aspects. Int J Solids Struct 207:279–294 Fincato R, Tsutsumi S (2020b) An overstress elasto-viscoplasticity model for high/low cyclic strain rates loading conditions: Part II—Numerical analyses. Int J Solids Struct 207:247–261 Fincato R, Tsutsumi S (2021) Coupled elasto-viscoplastic and damage model accounting for plastic anisotropy and damage evolution dependent on loading conditions. Comput Meth Appl Mech Eng 387:114165
826
References
Fincato R, Tsutsumi S (2022) Numerical implementation of the multiplicative hyperelastic-based extended subloading surface plasticity model. Comput Meth Appl. Mech Eng 401:115612 Flanagan DP, Taylor LM (1987) An accurate numerical algorithm for stress integration with finite rotations. Comput. Meth. Appl. Mech. Eng. 62:305–320 Franciosi P, Zaoui A (1991) Crystal hardening and the issue of uniqueness. Int J Plasticity 7:295– 311 Fredriksson B (1976) Finite element solution of surface nonlinearities in structural mechanics with special emphasis to contact and fracture mechanics problems. Comput Struct 6:281–290 Gambin W, Barlat F (1997) Modeling of deformation texture development based on rate independent crystal plasticity. Int J Plasticity 13:75–85 Gearing BP, Moon HS, Anand L (2001) A plasticity model for interface friction: application to sheet metal forming. Int J Plast 17:237–271 Germain P, Nguyen QS, Suquet P (1983) Continuum thermodynamics. J Appl Mech 50:1010– 1020 Gotoh M (1985) A class of plastic constitutive equations with vertex effect. Int. J. Solids Structures 21:1101–1163 Goya M, Ito K (1991) An expression of elastic-plastic constitutive laws incorporating vertex formulation and kinematic hardening. J. Appl. Mech. (ASME) 58:617–622 Green AE, Naghdi PM (1965) A general theory of an elastic-plastic continuum. Arch Ration Mech Anal 18:251–281 Gudehus G (1973) Elastoplastische Stoffgleichungen fur trockenen Sand. Ing Arch 42:151–169 (in German) Gudehus G (1979): A comparison of some constitutive laws for soils under radially symmetric loading and unloading, Proc. 3rd Int. Conf. Numer. Meth. Geomech. (Wittke, W. ed.), Aachen, Balkema Publ. (Rotterdam), pp. 1309- 1323. Guo S, Kang G, Zhang J (2013) A cyclic visco-plastic constitutive model for time-dependent ratchetting of particle-reinforced metal matrix composites. Int J Plast 40:101–125 Gurson AL (1977) Continuum theory of ductile rupture by void nucleation and growth: part I – yield criteria and flow rules for porous media. J Eng Mater Technol (ASME) 99:2–15 Han C-S, Chung K, Wagoner RH, Oh S-I (2003) A multiplicative finite elasto-plastic formulation with anisotropic yield functions. Int J Plasticity 19:197–211 Harder J (1999) A crystallographic model for the study of local deformation processes in polycrystals. Int J Plasticity 15:605–624 Harrysson M, Ristinmaa M (2007) Description of evolving anisotropy at large strains. Mech Mater 39:267–282 Hashiguchi K (1972) Theories of yielding for frictional materials. Proc JSCE 199:57–66 (in Japanese) Hashiguchi K (1974) Isotropic hardening theory of granular media Proc Jpn Soc Civil Eng 227:45–60 (in Japanese) Hashiguchi K (1977) An expression of anisotropy in a plastic constitutive equation of soils. In: Murayama S, Schofield AN (eds) Constitutive equations of soils (Proc. 9th Int. Conf. Soil Mech. Found. Eng., Spec. Ses.), vol 9. Tokyo, JSSMFE, pp 302–305 Hashiguchi K (1978) Plastic constitutive equations of granular materials. In: Cowin SC, Satake M (eds) Proceedings of US-Japan seminar on continuum mechanical and statistical approaches in the mechanics of granular materials. Sendai, pp 321–329 Hashiguchi K (1980) Constitutive equations of elastoplastic materials with elastic-plastic transition. J. Appl. Mech. (ASME) 47:266–272 Hashiguchi K (1981) Constitutive equations of elastoplastic materials with anisotropic hardening and elastic-plastic transition. J Appl Mech (ASME) 48:297–301 Hashiguchi K (1985a) Macrometric approaches-static-intrinsically time-independent. In: Constitutive laws of soils (Proc. Dsicuss. Ses. 1A, 11th Int. Conf. Soil Mech. Found. Eng.). San Francisco, pp 25–65
References
827
Hashiguchi K (1985b) Subloading surface model of plasticity. In: Constitutive Laws of Soils (Proc Dsicuss Ses 1A, 11th International Conference on Soil Mechanics and Foundation Engineering), San Francisco, pp 127–130 Hashiguchi K (1986) Elastoplastic constitutive model with a subloading surface. In: Proceedings of International Conference on Computing Mechanics, pp IV65–70 Hashiguchi K (1988) A mathematical modification of two surface model formulation in plasticity. Int J Solids Struct 24:987–1001 Hashiguchi K (1989) Subloading surface model in unconventional plasticity. Int J Solids Struct 25:917–945 Hashiguchi K (1991) Inexpedience of the non-associated flow rule. Int J Numer Anal Meth Geomech 15:753–756 Hashiguchi K (1993a) Fundamental requirements and formulation of elastoplastic constitutive equations with tangential plasticity. Int. J. Plasticity 9:525–549 Hashiguchi K (1993b) Mechanical requirements and structures of cyclic plasticity models. Int. J. Plasticity 9:721–748 Hashiguchi K (1994) On the loading criterion. Int J Plast 9:721–748 Hashiguchi K (1995) On the linear relations of V-lnp and lnv-lnp for isotropic consolidation of soils. Int J Numer Anal Meth Geomech 19:367–376 Hashiguchi K (1997) The extended flow rule in plasticity. Int J Plasticity 13:37–58 Hashiguchi K (1998) The tangential plasticity. Met Mater 4:652–656 Hashiguchi K (2000) Fundamentals in constitutive equation: continuity and smoothness conditions and loading criterion. Soils Found 40(3):155–161 Hashiguchi K (2001a) Description of inherent/induced anisotropy of soils: rotational hardening rule with objectivity. Soils Found 41(6):139–145 Hashiguchi K (2001b) On the thermomechanical approach to the formulation of plastic constitutive equations. Soils Found 41(4):89–94 Hashiguchi K (2002) A proposal of the simplest convex-conical surface for soils. Soils Found 42 (3):107–113 Hashiguchi K (2003) On the replacement of material-time derivative to corotational rate of yield function: mathematical proof. Soils Found 43(5):189–194 Hashiguchi K (2006) Constitutive model of friction with transition from static-to kinetic-friction-time-dependent subloading-friction model. In: Proceedings of the international symposium plasticity 2006, pp 178–180 Hashiguchi K (2007a) General corotational rate tensor and replacement of material-time derivative to corotational derivative of yield function. Comput Model Eng Sci 17:55–62 Hashiguchi K (2007b) Anisotropic constitutive equation of friction with rotational hardening. In: Proceedings of 13th International symposium on plasticity & its current applications, pp 34–36 Hashiguchi K (2007c) Extended overstress model for general rate of deformation including impact load. Proc 13th Int Symp Plasticity Its Current Appl, 37–39 Hashiguchi K (2007d) Yield condition of soils with tensile yield strength and rotational hardening. Proc Int Conf Comput Exp Eng Sci 07:1441–1446 Hashiguchi K (2008) Verification of compatibility of isotropic consolidation characteristics of soils to multiplicative decomposition of deformation gradient. Soils Found 48:597–602 Hashiguchi K (2011) General interpretations and tensor symbols for pull-back, push-forward and convected derivative. Proc JSME 24th Comp Mech Conf, 669–671 Hashiguchi K (2013) General description of elastoplastic deformation/sliding phenomena of solids in high accuracy and numerical efficiency: Subloading Surface Concept. Arch. Compt. Meth. Eng. 20:361–417 Hashiguchi K (2015) Cyclic stagnation of isotropic hardening in metals. Proc 2nd Sci Meeting of Kyushu Branch of Soc Mater Sci Japan, B18 Hashiguchi K (2016a) Exact formulation of subloading surface model: unified constitutive law for irreversible mechanical phenomena in solids. Arch Comput Methods Eng 23:417–447
828
References
Hashiguchi K (2016b) Loading criterion in return-mapping for subloading surface model. Proc Comput Mech Division JSME, 03–6 Hashiguchi K (2018a) Extension of loading criterion in return-mapping from infinitesimal to multiplicative hyperelastic-bsed plasticity for subloading surface model. Proc Comput Eng Conf (JSCES) A-08-01 Hashiguchi K (2018b) Evolution rule of elastic-core in subloading surface model in current and multiplicative hyperelastic-based plasticity. Proc Comput Eng Conf(JSCES), A-09-04 Hashiguchi K (2018c) Hypo-elastic and hyper-elastic equations of soils. Int J Numer Anal Meth Geomech 42:1–11 Hashiguchi K (2018d) Multiplicative hyperelastic-based plasticity for finite elastoplastic deformation/sliding: A comprehensive review. Arch Compt Meth Eng 23:1–41 Hashiguchi K (2021a) Complete implicit stress-integration for the subloading surface model with tangential-inelasticity and hardening stagnation. Proc Comput Mech Division, JSME, 006 Hashiguchi K (2021b) Subloading-friction model with limitation of tangential contact stress. Proc. MDT2021, JSME Hashiguchi K (2022a) Rigorous formulation of viscoplastic strain rate in subloading-overstress model. Proc Comput Eng Conf (JSCES) 27:C000021 Hashiguchi K (2022b) Rigorous evolution rule of tangential-inelastic strain rate in subloading surface model. Proc Comput Eng Conf (JSCES) 27:C000022 Hashiguchi K (2022c) Isotropic hardening stagnation in infinitesimal and multiplicative-infinite elastoplasticity with subloading surface concept and implicit integration. Proc Comput Eng Conf (JSCES) 27:000037 Hashiguchi K, Chen Z-P (1998) Elastoplastic constitutive equations of soils with the subloading surface and the rotational hardening. Int J Numer Anal Meth Geomech 22:197–227 Hashiguchi K, Mase T (2007) Extended yield condition of soils with tensile strength and rotational hardening. Int J Plast 23:1939–1956 Hashiguchi K, Mase T, Yamakawa Y (2021) Elaborated subloading surface model for accurate description of cyclic mobility in granular materials. Acta Geotech 17:699–719 Hashiguchi K, Okamura K (2014) Subloading-phase transformation model. In: Proceedings of 27th JSME computational mechanics division conferences, pp OS17–1707 Hashiguchi K (2007) Anisotropic constitutive equation of friction with rotational hardening. In: Proceedings of the 13th international symposium plasticity & its current applied, pp 34–36 Hashiguchi K, Ozaki S (2008a) Constitutive equation for friction with transition from static to kinetic friction and recovery of static friction. Int J Plast 24:2102–2124 Hashiguchi K, Ozaki S (2008b) Anisotropic constitutive equation for friction with transition from static to kinetic friction and vice versa. J Appl Mech (JSCE) 11:271–282 Hashiguchi K, Ozaki S, Okayasu T (2005) Unconventional friction theory based on the subloading surface concept. Int J Solids Struct 42:1705–1727 Hashiguchi K, Protasov A (2004) Localized necking analysis by the subloading surface model with tangential-strain rate and anisotropy. Int. J. Plasticity 20:1909–1930 Hashiguchi K, Saitoh K, Okayasu T, Tsutsumi S (2002) Evaluation of typical conventional and unconventional plasticity models for prediction of softening behavior of soils. Geotechnique 52:561–573 Hashiguchi K, Tsutsumi S (2001) Elastoplastic constitutive equation with tangential stress rate effect. Int J Plast 17:117–145 Hashiguchi K, Tsutsumi S (2003) Shear band formation analysis in soils by the subloading surface model with tangential stress rate effect. Int J Plasticity 19:1651–1677 Hashiguchi K, Tsutsumi S (2005) General non-proportional loading behavior of soils. Int J Plasticity 21:767–797 Hashiguchi K, Tsutsumi S (2007) Gradient plasticity with the tangential subloading surface model and the prediction of shear band thickness of granular materials. Int J Plast 22:767–797
References
829
Hashiguchi K, Ueno M (1977) Elastoplastic constitutive laws of granular materials. In: Murayama S, Schofield AN (eds) Constitutive Equations of Soils (Proceedings of 9th International Conference on Soil Mechanics Foundation Engineering, Special Session 9, Tokyo, JSSMFE, pp 73–82 Hashiguchi K, Ueno M (2017) Elastoplastic constitutive equation of metals under cyclic loading. Int J Eng Sci 111:86–112 Hashiguchi K, Ueno M (2022) Subloading-friction model with saturation of tangential contact stress. Friction, 10. https://doi.org/10.1007/s40544-022-0656-z Hashiguchi K, Ueno M, Anjiki T (2023) Subloading-overstress model: Unified constitutive equation for elasto-plastic and elasto-viscoplastic deformations under monotonic and cyclic loadings -Research with Systematic Review. Arch Compt Meth Eng 30. https://link.springer. com/article/10.1007/s11831-22-09880-y Hashiguchi K, Ueno M, Kuwayama T, Suzuki N, Yonemura S, Yoshikawa N (2016) Constitutive equation of friction based on the subloading-surface concept. Proc Royal Soc London A472 472:20160212. https://doi.org/10.1098/rspa.2016.0212 Hashiguchi K, Ueno M, Ozaki T (2012) Elastoplastic model of metals with smooth elastic-plastic transition. Acta Mech 223:985–1013 Hashiguchi K, Yoshimaru T (1995) A generalized formulation of the concept of nonhardening region. Int J Plast 11:347–365 Hassan T, Taleb T, Krishna S (2008) Influence of non-proportional loading on ratcheting responses and simulations by two recent cyclic plasticity models. Int J Plasticity 24:1863–1889 Huang Y (1991) A user-material subroutine incorporating single crystal plasticity in the Abaqus finite element program. Harvard univ, Division Appl. Sci. Havner KS (1982) The theory of finite plastic deformation of crystalline solids. In: Mechanics of solids—Rodney Hill 60th Anniversary Volume, Pergamon, pp 265–302 Hecker SS (1972) Experimental investigation of corners in yield surface. Acta Mech 13:69–86 Hencky H (1924) Zur Theorie plastischer Deformationen und der hierdurch im Material herforgerufenen Nachspannungen. Z. a. m. m. 4:323–334 Henshell GA, Miller AK, Tanaka JG (1987) Modeling cyclic deformation with MATMOD-Bossor. Unified constitutive equations, 2nd Int. Conf. on Low-cycle fatigue and elasto-plastic behavior of materials, Munich Higuchi R, Okamura K (2016) Estimation of residual stress change due to cyclic loading by classical elastoplasticity model and subloading surface model. Key Eng Mater 75:281–286 Hill R (1948a) Theory of yielding and plastic flow of anisotropic metals. Proc R Soc Lond A193:281–297 Hill R (1948b) A variational principle of maximum plastic work in classical plasticity. Q J Mech Appl Math 1:18–28 Hill R (1959) Some basic principles in the mechanics of solids without a natural time. J Mech Phys Solids 7:225–229 Hill R (1966) Generalized constitutive relations for incremental deformation of metal crystals. J Mech Phys Solids 14:95–102 Hill R (1968) On the constitutive inequalities for simple materials—1. J Mech Phys Solids 16:229–242 Hill R (1978) Aspects of invariance in solid mechanics. Adv Appl Mech 18:1–75 Hill R (1983) On the intrinsic eigenstates in plasticity with generalized variables. Math Proc Cambridge Phil Soc 93:177–189 Hill R (1990) Constitutive modeling of orthotropic plasticity in sheet metals. J Mech Phys Solids 38:241–249 Hill R, Rice JR (1972) Constitutive analysis of elastic-plastic crystals at arbitrary strain. J Mech Phys Solids 20:401–413 Hoger A, Carlson DE (1984) On the derivative of the square root of a tensor and Guo’s theorem. J. Elasticity 14:329–336 Hohenemser K, Prager W (1932) Uber die Ansatze der Mechanik isotroper Kontinua. Z A M M 12:216–226
830
References
Holzapfel GA (1996) On large strain viscoelasticity: Continuum formulation and finite element applications to elastomeric structures. Int J Numer Meth Eng 39:3903–3926 Holzapfel GA, Gasser TC (2001) A viscoelastic model for fiber-reinforced composite at finite strains: continuum basis, computational aspects and applications. Comput Methods Appl Mech Eng 190:4379–4430 Holzapfel GA, Simo JC (1996) Entropy elasticity of isotropic rubber-like solids at finite strains. Comput Methods Appl Mech Eng 132:17–44 Horowitz F, Ruina A (1989) Slip patterns in a spatially homogeneous fault model. J Geophys Res 94:10279–10298 Hosford WF (1974) A generalized isotropic yield criterion. J Appl Mech (ASME) 41:607–609 Houlsby GT (1985) The use of a variable shear modulus in elastic-plastic models for clays. Comput Geotech 1:3–13 Houlsby GT, Amorosi A, Rojas E (2005) Elastic moduli of soils dependent on pressure: a hyperelastic formulation. Geotechnique 55(5):383–392 Howe PG, Benson DP, Puddington IE (1955) London-Van der Waals’ attractive forces between glass surface. Can J Chem 33:1375–1383 Hughes TJR, Shakib F (1986) Pseudo-corner theory: A simple enhancement of J2-flow theory for applications involving non-proportional loading. Eng Comput 3:116–120 Hughes TJR, Winget J (1980) Finite rotation effects in numerical integration of rate consistent equations arising in large-deformation analysis. Int J Numer Meth Eng 15:1862–1867 Hutchinson JW (1976) Bounds and self-consistent estimates for creep of polycrystalline materials. Proc R Soc Lond A 348:101–127 Iguchi T, Sato T, Luo, Yamakawa Y (2019) Fully-implicit stress-update algorithm for extended subloading surface model incorporating nonhardening strain region. J Appl Mech (JAME) 121 (5):945–966 (in Japanese) Ikegami K (1979) Experimental plasticity on the anisotropy of metals. Proc. Euromech. Colloquium 115:201–242 Ilyushin AA (1961) On the postulate of plasticity. Appl.Math Meek 25:746–752 (Translation of O postulate plastichnosti. Prikladnaya Mathematika i Mekkanika 25:503–507 Itskov M (2003) Computation of the exponential and other isotropic tensor functions and their derivatives. Comput Methods Appl Mech Eng 192:3985–3999 Itskov M, Aksel N (2002) A closed-form representation for the derivative of non-symmetric tensor power series. Int J Solids Struct 39:5963–5978 Inoue T, Raniecki B (1978) Determination of thermal-hardening stress in steels by use of thermoplasticity theory. J Mech Phys Solids 26:187–212 Inoue T, Watanabe Y, Okamura K, Narazaki M, Shichino H, Ju D-Y, Kanamori H, Ichitani K (2007) Metallo-thermo-mechanical simulation of carburized quenching process by several codes – a benchmark project. Key Eng Mater 340–341:1061–1066 Itasca Consulting Group (2006) FLAC3D, “Fast Lagrangian Analysis of Continua in 3 Dimensions”, Minneapolis, Minnesota, USA Ito K (1979) New flow rule for elastic-plastic solids based on KBW model with a view to lowering the buckling stress of plates and shells. Tech Report Tohoku Univ 44: 199–232 Iwan WD (1967) On a class of models for the yielding behavior of continuous and composite systems. J Appl Mech (ASME) 34:612–617 Jaumann G (1911) Geschlossenes system physicalisher und chemischer differentialgesetze. Sitzber Akad 120:385–530 Jiang Y, Zhang J (2008) Benchmark experiments and characteristic cyclic plasticity deformation. Int J Plasticity 24:1481–1515 Johnson GR, Cook WH (1983) A constitutive model and data for metals subjected to large strain, high strain rates and high temperatures. In: Proceedings of 7th international sympoisum ballistics. The Hague, pp 541–547 Johnson WA, Mehl RF (1939) Reaction kinetics in processes of nucleation and growth. Trans ASME 135:416–458
References
831
Kachanov LM (1958) On rupture time under condition of creep. Izvestia Akademi Nauk SSSR. Otd Tekhn Nauk 8(1958):26–31 (in Russian) Kame N, Fujita S, Nakatani M, Kusakabe T (2013) Effects of a revised rate- and state-dependent friction law on aftershock triggering model. Tectonophys 600:187–195 Kato S, Sato N, Matsubayashi T (1972) Some considerations on characteristics of static friction of machine tool sideway. J Lubr Tech (ASME) 94:234–247 Khojastehpour M, Hashiguchi K (2004a) The plane strain bifurcation analysis of soils by the tangential-subloading surface model. Int J Solids Struct 41:5541–5563 Khojastehpour M, Hashiguchi K (2004b) Axisymmetric bifurcation analysis in soils by the tangential-subloading surface model. J Mech Phys Solids 52:2235–2262 Khojastehpour M, Murakami Y, Hashiguchi K (2006) Antisymmetric bifurcation in a circular cylinder with tangential plasticity. Mech. Materials 38:1061–1071 Kikuchi N, Oden JT (1988) Contact problem in elasticity: a study of variational inequalities and finite element methods. SIAM, Philadelphia Klisinski M, Abifadel N, Runesson K, Sture S (1991) Modelling of the behavior of dry sand by an elasto-plastic “fuzzy set” model. Comput Geotech 11:229–261 Knockaert R, Chastel Y, Massoni YCE (2000) Rate-independent crystalline plasticity, application to FCC materials. Int J Plasticity 16:179–198 Kobayashi M, Ohno N (2002) Implementation of cyclic plasticity models based on a general form of kinematic hardening. Int J Numer Meth Eng 53:2217–2238 Kohgo Y, Nakano M, Miyazaki T (1993) Verification of the generalized elastoplastic model for unsaturated soils. Soil Found 33(4):64–73 Koiter WT (1953) Stress-strain relations, uniqueness and variational theories for elastic-plastic materials with a singular yield surface. Quart Appl Math 11:350–354 Kojic M, Bathe K-J (1987) Studies of finite element procedures – Stress solution of a closed elastic strain path with stretching and shearing using the update Lagrangian Jaumann formulation. Compt Struct 26:175–179 Kolymbas D, Wu W (1993) Introduction to plasticity. Elsevier, Modern Approaches to Plasticity, pp 213–224 Kratochvil J (1971) Finite-strain theory of crystalline elastic-inelastic materials. J Appl Phys 42:1104–1108 Krieg RD (1975) A practical two surface plasticity theory. J Appl Mech (ASME) 42:641–646 Krieg RD, Key SW (1976) Implementation of a time dependent plasticity theory into structural computer programs. Constitutive Equations in Viscoplasticity: Computational and Engineering Aspects (eds. Strickin, J. A. and Saczlski, K. J.), AMD-20, ASME Krieg RD, Krieg DB (1977) Accuracies of numerical solution methods for the elastic-perfectly plastic models, J Pressure Vessel Tech (ASME) 99:510–515 Kroner E (1960) Allgemeine Kontinuumstheoreie der Versetzungen und Eigenspannnungen. Arch Ration Mech Anal 4:273–334 Kuroda M (1997) Interpretation of the behavior of metals under large plastic shear deformations: a macroscopic approach. Int J Plast 13:359–383 Kuroda M, Tvergaard V (2001) A phenomenological plasticity model with non-normality effects representing observations in crystal plasticity. J Mech Phys Solids 49:1239–1263 Kurtyka T, Zyczkowsky M (1996) Evolution equations for distortional plastic hardening. Int. J. Plasticity 12:191–213 Ladevéze P, Lemaitre JA (1984) Damage effective stress in quasi unilateral conditions. 16th Int. Congr. Theor. Appl. Mech., Lyngby, Denmark Leckie EA, Onate ET (1981) Tensorial nature of damage measuring internal variables. In: Proceedings of IUTAM symposium on physics nonlinearities in structures, pp 140–155 Lee EH, Liu DT (1967) Finite-strain elastic-plastic theory with application to plane-wave analysis. J Appl Phys 38:19–27 Lee EH (1969) Elastic-plastic deformation at finite strain. J Appl Mech (ASME) 36:1–6
832
References
Lee J, Bong HJ, Kim S-J, Lee M-G, Kim D (2020) An enhanced distortional-hardening-based constitutive model for hexagonal close-packed metals: application to AZ31B magnesium alloy sheets at elevated temperatures. Int J Pasticity 126:102618 Lemaitre JA (1971) Evaluation of dissipation and damage in metals subjected to dynamic loading. In: Proceedings of international congress of mechanical behavior of materials Lemaitre JA (1990) Micro-mechanics of crack initiation. Int J Frac 42:87–99 Lemaitre JA, Doghri I (1994) Damage 90: a post-processor for crack initiation. Comput Methods Appl Mech Eng 115:197–232 Lemaitre JA, Dosmorat R, Sauzay M (2000) Anisotropic damage law of evolution. Eur J Mech A/Solids 19:182–208 Lemaitre J, Sermage JP, Desmorat R (1999) A two scale damage concept applied to fatigue. Int J Fract 97:67–81 Li W, Ma J, Kou H, Shao J, Zhang X, Deng Y, Tao Y, Fang D (2019) Modeling the effect of temperature on the yield strength of precipitation strength Ni-base superalloys. Int J Plasticity 116:143–158 Lin FB, Bažant ZP, Chern JC, Marchertas AH (1987) Concrete model with normality and sequential identification. Comput Struct 26:1011–1025 Liu D, Yang H, Elkhodary KL, Tang S, Liu WK, Guo X (2022) Mechanistically informed data-driven modeling of cyclic plasticity via artificial neural networks. Comput Meth Appl Mech Eng 393:114766 Lion A (2000) Constitutive modeling in finite thermoviscoplasticity: a physical approach based on nonlinear rheological models. Int J Plasticity 16:469–494 Loret B (1983) On the effects of plastic rotation in the finite deformation of anisotropic elastoplastic materials. Mech Mater 2:287–304 Lu Y, Jiang Y, Zhu W, Bao Y, Ye G, Zhang F (2022) Unified description of different soils based on the superloading and subloading surface concept. J Rock Mech Geotech Eng. https://doi. org/10.1016/j.jrmge.2022.02.015 Lubarda VA (2004) Constitutive theories based on the multiplicative decomposition of deformation gradient: thermoplasticity, elastoplasticity, and biomechanics. Appl Mech Rev 57:95–108 Lubarda VA, Krajcinovic D (1993) Damage tensors and the crack density distribution. Int J Solids Struct 30:2859–2877 Lubarda VA, Krajcinovic D, Mastilovic S (1994) Damage model for brittle elastic solids with unequal tensile and compressible strength. Eng Fract Mech 49:681–697 Lubarda VA, Lee EH (1981) A correct definition of elastic and plastic deformation and its computational significance. J Appl Mech (ASME) 48:35–40 Lubliner J (1984) A maximum-dissipation principle in generalized plasticity. Acta Mech 52:225– 237 Macvean DB (1968) Die elementararbeit in einem kontinuum und die zuordnung von spannungsund verzerrungstensoren. Z Angew Math Phys 19:157–185 Mandel J (1964) Contribution theorique a l’eude de l’ecrouissage et des lois de l’ecoulement plastique. Proc. 11th Int. Congr. Appl. Mech. pp. 502–509 Mandel J (1965) Generalisation de la theorie de plasticite de W.T. Koiter. Int J Solids Struct 1:273–295 Mandel J (1971) Plastidite classique et viscoplasticite. Course & lectures, No. 97, international centre for mechanical sciences, Udine, Springer-Verlag, Heidelberg Mandel J (1973) Equations constitutives directeurs dans les milieux plastiques at viscoplastiques. Int J Solids Struct 9:725–740 Mandel J (1974) Director vectors and constitutive equations for plastic and viscoplastic media. In: Sawczuk A (ed) Problems of plasticity (Proceedings of the international symposium foundation of plasticity). Noordhoff Int. Publ., Leyden, Netherland, pp 135–141 Maranha JR, Pereira C, Vieira A (2016) A viscoplastic subloading soil model for rat-dependent cyclic anisotropic structured behavior. Int J Nemer Anal Meth Geomech 40:1531–1555
References
833
Mase T, Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model. Soils Found 49:207–220 Masing G (1926) Eigenspannungen und Verfestigung beim Messing. In: Proceedings of 2nd International Congress on Applied Mechanics, Zurich, pp 332–335 Matsubara S, Terada K, Maeda R, Kobayashi T, Murata M, Sumiyama T, Furuichi K, Nonomura C (2015) Viscoelastic-viscoplastic combined constitutive model for glassy amorphous polymers under loading/unloading/no-load states. Eng Comput 37:1703–1735 Matsuoka H, Nakai T (1974) Stress-deformation and strength characteristics of soil under three different principal stress. Proc Jpn Soc Civil Eng 232:59–70 Matsuoka H, Yao YP, Sun DA (1999) The Cam-clay model revised by SMP criterion. Soils Found 39(1):81–95 McDowell DL (1985) An experimental study of the structure of constitutive equations for nonproportional cyclic plasticity. J Eng Mater Tech (ASME) 107:307–315 McDowell DL (1989) Evaluation of intersection conditions for two-surface plasticity theory. Int J Plasti 5:29–50 Mengoni M, Ponthot JP (2015) A generic anisotropic continuum damage model integration scheme adaptable to both ductile damage and biological damage-like situations. Int. J. Plasticity 66:46–70 Michalowski R, Mroz Z (1978) Associated and non-associated sliding rules in contact friction problems. Archiv Mech 30:259–276 Miehe C (1995) Discontinuous and continuous damage evolution in Ogden-type large-strain elastic materials. European J Mech A/Solids 14:697–720 Miehe C (1996) Numerical computation of algorithmic (consistent) tangent moduli in large-strain Computational inelasticity. Comput Methods Appl Mech Eng 134:223–240 Miehe C, Schroder J (2001) A comparative study of stress update algorithms for rate-independent and rate-dependent crystal plasticity. Int J Num Meth Eng 50:273–298 Mindlin RD (1963) Influence of couple-stresses on stress concentrations. Experiment. Mech. 3:1–7 Mooney M (1940) A theory of large elastic deformation. J Appl Phys 11(9):582–592 Maranha JR, Pereira C, Viera A (2016) A viscoplastic subloading soil model for rate-dependent cyclic anisotropic structured behaviour. Int J Numer Anal Meth Geomech 40:1531–1555 Mroz Z (1966) On forms of constitutive laws for elastic-plastic solids. Arch. Mech. Stos. 18:3–35 Mroz Z (1967) On the description of anisotropic workhardening. J Mech Phys Solids 15:163–175 Mroz Z (1976) A non-linear hardening model and its application to cyclic plasticity. Acta Mech 25:51–61 Mroz Z, Stupkiewicz S (1994) An anisotropic friction and wear model. Int J Solids Struct 31:1113–1131 Muhlhaus HB, Vardoulakis I (1987) The thickness of shear bands in granular materials. Geotechnique 37:271–283 Mullins L (1947) Effect of stretching on the properties of rubber. J. Rubber Research 16:275–289 Mullins L (1969) Softening of rubber by deformation. Rubber Chem Tech 42:339–362 Murakami S (1988) Mechanical modelling of material damage. J Appl Mech (ASME) 55:280–286 Murakami S, Ohno N (1981) A continuum theory of creep and creep damage. In: Creep in structure, proceedings 3rd IUTAM symposium, pp 422–444 Nagtegaal JC, De Jong JE (1982) Some aspects of non-isotropic workhardening in finite strain plasticity. In: Lee EH, Mallett RL (eds) Plasticity of metals and finite strain: theory, experiment and computation, division of applied mechanics, Stanford's department of mechanical engineering, Rensselaer Polytechnic institute, pp 65–102 Nakada Y, Keh AS (1966) Latent hardening in iron single crystals. Acta Metall 14:961–973 Nakai T, Hinokio M (2004) A simple elastoplastic model for normally and over consolidated soils with unified material parameters. Soils Found 44(2):53–70 Nakai T, Mihara Y (1984) A new mechanical quantity for soils and its application to elastoplastic constitutive models. Soils Found 24(2):82–941
834
References
Needleman A, Rice JR (1978) Limits to ductility set by plastic flow localization. In: Koistinen DP, Wang N-M (eds) Mechanics of sheet metal forming. Plenum press, New York, pp 237–265 Needleman A, Tvergaard V (1985) Material strain-rate sensitivity in round tensile bar. In: Proceedings of international symposium plastic instability. Pressure de l’cole nationale des Ponts et Shausseses, Paris, pp 251–262 Niemunis A, Cudny M (1998) On hyperelasticity for clays. Comput Geotech 23:221–236 Nova R (1977) On the hardening of soils. Arch Mech Stos. 29:445–458 Oden JT, Martines JAC (1986) Models and computational methods for dynamic friction phenomena. Comput Meth Appl Mech Eng 52:527–634 Oden JT, Pires EB (1983) Algorithms and numerical results for finite element approximations of contact problems with non-classical friction laws. Comput Struct 19:137–147 Ohno N (1982) A constitutive model of cyclic plasticity with a non-hardening strain region. J Appl Mech (ASME) 49:721–727 Ohno N, Kachi Y (1986) A constitutive model of cyclic plasticity for nonlinearly hardening materials. J Appl Mech(ASME) 53:395–403 Ohno T, Nakamoto H, Morimatsu Y, Okjumura D (2021) Modeling od cyclic hardening and evaluation of plastic strain range in the presence of pre-loading and ratchetting. Int J plasticity 145:103074 Ohno N, Wang JD (1993) Kinematic hardening rules with critical state of dynamic recovery, Part I: formulation and basic features for ratcheting behavior. Part II: application to experiments of ratcheting behavior. Int J Plast 9:375–403 Ohno T, Yamamoto R, Okumura D (2018) Thermo-mechanical cyclic hardening behavior of 304 stainless steel at large temperature ranges: experiments and simulations. Int J Mech Sci 146– 147:517–526 Okahara M, Takagi S, Mori H, Koike S, Tatsuda M, Tatsuoka F, Morimoto H (1989) Largescale plane strain bearing capacity tests of shallow foundation on sand (part 1). In: Proceedings of 24th annual meeting Japanese geotechnical society engineering, pp 1239–1242 (in Japanese) Okamura K (2006a) Reviews and perspective on hardening simulation. In: Proc. Japan of and the and Japan, Kyushu-branch, pp 1–12 (in Japanese) Okamura K (2006b) Actuarity and scope on simulation of heat treatment: I: material properties and database. J Soc Mater Sci Jpn 55:529–535 (in Japanese) Okamura K, Kawashima H (1988a) Finite element analysis of thermal stress in heat treatment. Netu-Shori (heat Treatment) 28:141–148 (in Japanese) Okamura K, Kawashima H (1988b) Analysis of residual deformation of a gear during quenching. In: Proceedings of , pp 323–329 (in Japanese) Okamura K, Yamamoto K, Fukumoto M (2005) Material properties for quenching simulation and assessment on computational results. In: Proceedings of 3rd Asian on material, pp 353–355 Okorokov V, Gorash M, D. and van Rijswick, R. (2019) Now formulation of nonlinear kinematic hardening model, Part I: A Dirac delta function approach. Int J Plast 122:89–114 Oldroyd JG (1950) On the formulation of rheological equations of state. Proc R Soc Lond A200:523–541 Ortiz M, Radovitzky RA, Repetto EA (2001) The computation of the exponential and logarithmic mappings and their first and second linearizations. Int J Numer Meth Eng 52:1431–1441 Ozaki S, Hashiguchi K (2010) Numerical analysis of stick-slip instability by a rate-dependent elastoplastic formulation for friction. Tribology Int 43:2120–2133 Ozaki S, Hikida K, Hashiguchi K (2012) Elastoplastic formulation for friction with orthotropic anisotropy and rotational hardening. Int J Solids Struct 49:648–657 Ozaki S, Matsuura T, Maegawa S (2020) Rate-, state-, and pressure-dependent friction model based on the elastoplasticity theory. Friction 8:768–783 Ozaki T, Yamakawa Y, Ueno M, Hashiguchi K (2022) Description of sand–metal friction behavior by subloading-friction model. Friction (in printing) Pan J, Rice JR (1983) Rate sensitivity of plastic flow and implications for yield surface vertices. J Mech Phys Solids 19:973–987
References
835
Peirce D, Asaro JR, Needleman A (1982) Overview 21: an analysis of nonuniform and localized deformation in ductile single crystals. Act Metall 30:1087–1119 Peirce D, Asaro JR, Needleman A (1983) Overview 32: material rate dependence and localized deformation in crystal solids. Act Metall 31:1951–1976 Perić D (1992) On consistent stress rates in solids mechanics: computational implications. Int J Numer Meth Eng 33:799–817 Peric D, Owen RJ (1992) Computational model for 3-D contact problems with friction based on the penalty method. Int J Numer Meth Eng 35:1289–1309 Perzyna P (1963) The constitutive equations for rate sensitive plastic materials. Q Appl Math 20:321–332 Perzyna P (1966) Fundamental problems in viscoplasticity. Adv Appl Mech 9:243–377 Petryk H (1991) On the second-order work in plasticity. Arch Mech 43:377–397 Petryk H (1997) Plastic instability: Criteria and computational approaches. Arch Comput Approach Meth Eng 4:111–151 Petryk H, Thermann K (1997) A yield-vertex modification of two-surface models of metal plasticity. Arch Mech 49:847–863 Pietruszczak St, Mroz Z (1981) Finit element analysis of deformation of strain-softening materials. Int J Numer Anal Meth Geomech 17:327–334 Pietruszczak ST, Niu X (1993) On the description of localized deformation. Int J Numer Anal Meth Geomech 17:791–805 Pinsky PM, Ortiz M, Pister KS (1983) Numerical integration of rate constitutive equations in finite deformation analysis. Comput Meth Appl Mech Eng 193:5223–5256 Poh KW (2001) Stress-strain-temperature relationship for structural steel. J Mater Civil Eng 371:21953 Prager W (1945) Strain hardening under combined stress. J Appl Phys 16:837–840 Prager W (1949) Recent development in the mathematical theory of plasticity. J Appl Mech (ASME) 20:235–241 Prager W (1956) A new methods of analyzing stresses and strains in work hardening plastic solids. J Appl Mech (ASME) 23:493–496 Prager W (1961) Linearization in visco-plasticity. Ing Archiv 15:152–157 Rabinowicz E (1951) The nature of the static and kinetic coefficients of friction. J Appl Phys 22:1373–1379 Rabinowicz E (1958) The intrinsic variables affecting the stick-slip process. Proc Phys Soc 71:668–675 Rice JR (1970a) On the structure of stress-strain relations for time dependent plastic deformation in metals. J Appl Mech (ASME) 37:728–737 Rice JR (1970b) Inelastic constitutive relations for solids: an internal variable theory and its applications to metal plasticity. J Mech Phys Solids 19:433–455 Rice JR, Lapusta N, Ranjith K (2001) Rate and state dependent friction and the stability of sliding between elastically deformable solids. J Mech Phys Solids 49:1865–1898 Rice JR, Ruina AL (1983) Stability of steady frictional slipping. J Appl Mech (ASME) 50:343– 349 Richeton J, Ahzi S, Daridon I, Remond Y (2005) A formulation of the cooperative model for the yield stress of amorphous polymers for a wide range of strain rates and temperatures. Polymer 46:6035–6043 Rivlin RS (1948) Large elastic deformations of isotropic materials. IV. Further developments of the general theory. Phil Trans R Soc Lond Ser A Math Phys Sci 241(835):379–397 Rivlin RS (1955) Further remarks on the stress-deformation relations for isotropic materials. J Rational Mech Aanl 4:681–702 Rohde J, Jeppsson A (2000) Literature review of heat treatment simulations with respect to phase transformation, residual stresses and distortion. Scandinavian J Metallurgy 29:47–62 Rubinstein R, Atluri SN (1983) Objectivity of incremental constitutive relations over finite time steps in computational finite deformation analysis. Compt Meth Appl Meth Eng 36:277–290
836
References
Rudnicki JW, Rice JR (1975) Conditions for the localization of deformation in pressure-sensitive dilatant materials. J Mech Phys Solids 23:371–394 Ruina AL (1980) Friction laws and instabilities: quasistatic analysis of some dry frictional behavior, Ph.D. thesis. Brown University, Providence Ruina AL (1983) Slip instability and state variable friction laws. J Geophys Res 88:10359–10370 Rusinek A, Zaera R, Klepaczko JR (2007) Constitutive relations in 3D for a wide range of strain rates and temperature—application to mild steels. Int J Solids Struct 44:5611–5634 Sasaki T, Sogab K, Elshafiec MZEB (2018) Simulation of wellbore construction in offshore unconsolidated methane hydrate-bearing formation. J Natural Gas Sci Eng 60:312–326 Satake M (1972) A proposal of new yield criterion for soils. Proc Jpn Soc Civil Eng 189:79–88 (in Japanese) Scholz CH (1998) Rate-and state-variable friction law. Nature 391:37–41 Seguchi Y, Shindo A, Tomita Y, Sunohara M (1974) Sliding rule of friction in plastic forming of metal. Compt Meth Nonlinear Mech. Univ Texas Austin, pp 683–692 Sekiguchi H, Ohta H (1977) Induced anisotropy and its time dependence in clays. Constitutive equations of soils (Proc. Spec. Session 9, 9th ICSFME). Tokyo, pp 229–238 Seth BR (1964) Generalized strain measure with applications to physical problems Second-order effects inelasticity, plasticity, and fluid dynamics. Pergamon, Oxford Sewell MJ (1973) A yield-surface comer lowers the buckling stress of an elastic-plastic plate under compression. J Mech Phys Solids 21:19–45 Sewell MJ (1974) A plastic flow at a yield vertex. J Mech Phys Solids 22:469–490 Sheng D, Sloan SW, Yu HS (2000) Aspects of finite element implementation of critical state models. Comput Mech 26:185–196 Shi B, Peng Y, Yang C, Pan F, Cheng R, Peng Q (2017) Loading path dependent distortional hardening of Mg alloys: Experimental investigation and constitutive modeling. Int J Plasticity 90:76–95 Shield RT, Ziegler H (1958) On Prager’s hardening rule. Z Ang Math Mech 9:260–276 Siddiquee MSA, Tanaka T, Tatsuoka F, Tani K, Morimoto T (1999) Numerical simulation of bearing capacity characteristics of strip footing on sand. Soil Found 39(4):93–109 Simo JC (1987a) On a fully three-dimensional finite-stain viscoelastic damage model: formulation and computational aspects. Comput Meth Appl Mech Eng 60:153–173 Simo JC (1987b) A J2-flow theory exhibiting a corner-like effect and suitable for large-scale computation. Comput Meth Appl Mech Eng 62:169–194 Simo JC, Kennedy JG, Govindjee S (1988) Non-smooth multisurface plasticity and viscoplasticity –Loading unloading conditions and numerical algorithms. Int J Numer Meth Eng 26:2161– 2185 Simo JC, Meschke G (1993) A new class of algorithms for classical plasticity extended to finite strains. Appl Geomater Comput Mech 11:253–278 Simo JC, Ortiz M (1985) A unified approach to finite deformation elastoplasticity based on the use of hyperelastic constitutive equations. Comput Meth Appl Mech Eng 49:221–245 Simo JC, Pister KS (1984) Remarks on rate constitutive equations for finite deformation problems: Computational implications. Compt Meth Appl Mech Eng 46:201–215 Simo JC, Taylor RL (1985) Consistent tangent operators for rate-independent elastoplasticity. Comput Meth Appl Mech Eng 48:101–118 Simo JC, Taylor RL (1986) A return mapping algorithm for plane stress elastoplasticity. Int J Numer Meth Eng 22:649–670 Skempton AW, Brown JD (1961) A landslide in boulder clay at Selset, Yorkshire. Geotechnique 11:280–293 Sloan SW, Randolph MF (1982) Numerical prediction of collapse loads using finite element methods. Int J Numer Anal Meth Geomech 6:47–76 Stallebrass SE, Taylor RN (1997) The development and evaluation of a constitutive model for the prediction of ground movements in overconsolidated clay. Geotechnique 47:235–253
References
837
Stark TD, Ebeling RM, Vettel JJ (1994) Hyperbolic stress-strain parameters for silts. J Geotech Eng (ASCE) 120:420–441 Stribeck R (1902) Die Wesentlichen Eigenschaften der Gleit- und Rollenlager. Z Verein Deut Ing 46:1341–1348 (in German) Stupkiewicz S, Mroz Z (1999) A model of third body abrasive and wear in hot metal forming. Wear 231:124–138 Sun L, Wagoner RH (2011) Complex unloading behavior: Nature of the deformation and its consistent constitutive representation. Int J Past 27:1126–1144 Tamagnini C, Castellanza R, Nova R (2002) A generalized backward Euler algorithm for the numerical integration of an isotropic hardening elastoplastic model for mechanical and chemical degradation of bonded geomaterials. Int J Numer Anal Meth Geomech 26:963–1004 Tanaka T, Kawamoto O (1988) Three dimensional finite element collapse analysis for foundations and slopes using dynamic relaxation. In: Proc. of Numerical Methods in Geomechanics. Innsbruck, pp 1213–1218 Tanaka T, Sakai T (1993) Progressive failure effect of trap-door problems with granular materials. Soils Found 33(1):11–22 Tani K (1986) Mechanism of bearing capacity of shallow foundation, Master Thesis. Univ, Tokyo (in Japanese) Tatsuoka F, Ikuhara O, Fukushima S, Kawamura T (1984) On the relation of bearing capacity of shallow footing on model sand ground and element test strength. In: Proceedings of Symp. Asses. Deform. Fail. Strength of Sandy Soils and Sand Grounds, Japan. Soc. Geotech. Eng., pp 141–148 (in Japanese) Tatsuoka F, Iwasaki and T., Takagi, Y. (1978) Hysteretic damping of sands under cyclic loading and its relation to shear modulus. Soils Found 18(2):25–40 Taylor GI (1938) Plastic strain in metals. J Inst Metals 62:307–324 Topolnicki M (1990) An elasto-plastic suboading surface model for clay with isotropic and kinematic mixed hardening parameters. Soil Found 30(2):103–113 Truedell C (1952) The mechanical foundations of elasticity and fluid mechanics. J Rational Mech Anal 1:125–300 Truesdell C (1955) Hypo-elasticity. J Rational Mech Anal 4:83–133 Tsutstumi S, Fincato R, Momii H (2019) Effect of tangential plasticity on structural response under non-proportional cyclic loading. Acta Mech 230:2425–2446 Tsutsumi S, Hashiguchi K (2005) General non-proportional loading behavior of soils. Int J Plasticity 21:1941–1969 Tvergaard V (1982) On localization in ductile materials containing spherical voids. Int J Fracture 18:237–252 Tvergaard V, Needleman A (1984) Analysis of the cup-cone fracture in a round tensile bar. Acta Metall 32:157–169 Ulz MH, Celigoj CC (2021) A uniquely defined multiplicative elasto-plasticity model with orthotropic yield function and plastic spin. Compt Meth Appl Mech Eng 374:1–24 Uesugi M, Kishida H (1986a) Influential factors of friction between steel and dry sands. Soil Found 26(2):33–46 Uesugi M, Kishida H (1986b) Frictional resistance at yield between dry sand and mild steel. Soil Found 26(4):139–149 Vermeer PA (1982) A simple shear band analysis using compliances. In: Proceedings of the IUTAM symposium deformation and failure of granular materials, Balkema, Amsterdam, pp 493–499 Vladimirov IN, Pietryga MP, Reese S (2008) On the modeling of nonlinear kinematic hardening at finite strains with application to springback comparison of time integration algorithm. Int J Numer Meth Eng 75:1–28 Vladimirov IN, Pietryga MP, Reese S (2010) Anisotroipc finite elastoplasticity with nonlinear kinematic and isotropic hardening and application to shear metal forming. Int J Plasticity 26:659–687
838
References
von Mises R (1923) Mechanik der plastischen Formanderung von Kristallen. Z Angew Math Mech 8:161–185 Voyiadjis and Taqieddin, ZKattan, P. I. , 2008.Voyiadjis G, Taqieddin Z, Kattan PI (2008) Anisotropic damage-plasticity model for concrete. Int J Plast 24:1946–1965 Vu DK, Staat M (2007) Shakedown analysis of structures made of materials with temperature-dependent yield stress. Int J Soilds Struct 44:4524–4540 Wallin M, Ristinmaa M (2005) Deformation gradient based kinematic hardening model. Int J Plasticity 21:2025–2050 Wallin M, Ristinmaa M, Ottesen NS (2003) Kinematic hardening in large strain plasticity. Eur J Mech A/Solids 22:341–356 Wagoner RH, Lim H, Lee M-G (2013) Advanced issues in springback. Int J Pasticity 45:3–20 Watchtman JB Jr, Tefft WE, Lam FG Jr, Apstein CS (1961) Exponential temperature dependence of Young’s modulus for several oxides. Phys Review 122:1754–1759 Wesley LD (1990) Influence of structure and composition on residual soils. J Geotech Eng (ASCE) 116:589–603 Wilde P (1977) Two invariants depending models of granular media. Arch Mech Stos 29:799–809 White PS (1975) Elastic-plastic solids as simple materials. Q J Mech Appl Math 28:483–496 William KJ, Warnke EP (1974) Constitutive model of the triaxial behaviour of concrete. Seminar on Concrete structures subjected to triaxial stresses, ISMES, Bergamo, Italy, Published in Int. Assoc. Bridge & Struct. Eng. (IABSE) Report 1974;19:III-1 Wongsaroj J, Soga K, Mair RJ (2007) Modeling of long-term ground response to tunneling under St James’ Park. London, Geotechnique 57:75–90 Wriggers P, Van Vu T, Stein E (1990) Finite element formulation of large deformation impact-contact problems with friction. Comput Struct 37:319–331 Wu JY, Li J, Faira R (2006) An energy release rate-based plastic-damage model for concrete. Int J Solid Struct 43:583–612 Wu J-Y, Xu S-L (2013) Reconsideration on the elastic damage/degradation theory for the modeling of microcrack closure-reopening (MCR) effects. Int J Solid Struct 50: 795-805 Wu DL, Xuan F-Z, Guo S-J, Zhao P (2018) Uniaxial mean stress relaxation of 9–12% Cr steel at high temperature: Experiments and viscoplastic constitutive modeling. Int J Plasticity 77: 156–173 Xia Z, Ellyin F (1994) Biaxial ratcheting under strain or stress-controlled axial cycling with constant hoop stress. J Appl Mech (ASME) 61:422–428 Xiao H (1995) Unified explicit basis-free expressions for time rate and conjugate stress of an arbitrary Hill strain. Int Solids Struct 32:3327–3340 Xiao H, Bruhns OT, Meyers A (1997) Logarithmic strain, logarithmic spin and logarithmic rate. Acta Mech 124:89–105 Xiao H, Bruhns OT, Meyers A (1999) Existence and uniqueness of the integrable-exactly
hypoelastic equation s ¼ kðtr DÞI þ 2lD and its significance to finite inelasticity. Acta Mech 138:31–50 Xiong Y-L, Yang Q-L, Zhang S, Ye G-L, Zheng RY, Zhang F (2018) Thermo-elastoplastic model for soft rock considering effects of structure and overconsolidation. Rock Mech Rock Eng 51:3771–3784 Yamakawa Y (2022) Hyperelastic constitutive models for geomaterials: Extension of existing models to include finite strains and their comparison. Comput Geotech 143: 104600 Yamakawa Y, Hashiguchi K, Ikeda K (2010a) Implicit stress-update algorithm for isotropic Cam-clay model based on the subloading surface concept at finite strains. Int J Plast 26: 634–658 Yamakawa Y, Hashiguchi K, Sasaki T, Higuchi M, Sato K, Kawai T, Machishima T, Iguchi T (2021) Anisotropic subloading surface Cam-clay model with rotational hardening: deformation gradient-based formulation for finite strain. Int J Numer Anal Meth Goemech. https://doi.org/ 10.1002/nag.3268
References
839
Yamakawa Y, Yamaguchi Y, Hashiguchi K, Ikeda K (2010) Formulation and implicit stress-update algorithm of the extended subloading surface Cam-clay model with kinematic hardening for finite strains. J Appl Mech (ASCE) 13:411–412 Yoshida F, Amaishi T (2020) Model for description of nonlinear unloading-reloading stress-strain response with special reference to plastic-strain dependent chord modulus. Int J Plasticity 30. https://doi.org/10.1016/j.ijplas.2020.102708 Yoshida F, Hamasaki H, Uemori T (2015) Modeling of anisotropic hardening of sheet metals including description of the Bauschinger effect. Int J Plasticity 75:170–188 Yoshida F, Uemori T (2002a) Elastic-plastic behavior of steel sheets under in-plane cyclic tension-compression at large strain. Int J Plast 18:633–659 Yoshida F, Uemori T (2002b) A model of large-strain cyclic plasticity describing the Bauschinger effect and workhardening stagnation. Int J Plast 18:661–686 Yoshida F, Uemori T (2003) A model of large-strain cyclic plasticity and its application to springback simulation. Int J Mech Sci 45:1687–1702 Yoshida K, Kuroda M (2012) Comparison of bifurcation and imperfection analyses of localized necking in rate-independent polycrystalline sheets. Int J Solids Struct 49:2073–2084 Yuanming L, Liao M, and Hu K (2016) A constitutive model of frozen saline sandy soil based on energy dissipation theor. Int J Plasticity 78:84–113 Yuanming L, Long, J, Xiaoxiao C (2009) Yield criterion and elasto-plastic damage constitutive model for frozen sandy soil. Int J Plasticity 25:1177–1205 Zamiri A, Pourbogharat F (2010) A novel yield function for single crystal based on combined constraints optimization. Int J Plasticity 26:731–746 Zamiri A, Pourbogharat F, Barlat F (2007) An effective computational algorithm for rate-independent crystal plasticity based on a single crystal yield surface with an application to tube hydroforming. Int J Plasticity 23:1126–1147 Zaremba S (1903) Su une forme perfectionnee de la theorie de la relaxation. Bull Int Acad Sci 594–614 (in French) Zbib HM, Aifantis EC (1988) On the concept of relative and plastic spins and its implications to large deformation theories. Part I: Hypoelasticity and Vertex-Type Plasticity. Acta Mech 75:15–33 Zhang S, Leng W, Zhang F, Xiong Y (2012) A simple thermo-elastoplastic model for geomaterials. Int J Plasticity 34:93–113 Zhang F, Ye B, Noda T, Nakano M, Nakai K (2007) Explanation of cyclic mobility of soils: approach by stress-induced anisotropy. Soils Found 47:635–648 Ziegler H (1959) A modification of Prager’s. hardening rule. Quart Appl Phys 17:55–60 Ziegler H, McVean D (1967) Recent progress in applied mechanics. Wiley
Index
A Acoustic tensor, 591 Additive decomposition strain rate, 91, 373, 506 continuum spin, 91 Admissible transformation, 84 Algorithmic tangent modulus, See Consistent tangent modulus Almansi (Eulerian) strain, See Strain Alternating symbol, See Permutation symbol Angular momentum, 145, 152, 155 Anisotropy, 183, 233, 235, 240, 346, 349, 351, 353, 355, 360, 363, 397, 398, 573, 641, 643, 694–696, 699–701 kinematic hardening, See Kinematic hardening orthotropic, See Orthotropic anisotropy rotational hardening, See Rotational hardening Anisotropy for friction orthotropic, 346, 349, 354, 355, 358, 360, 363, 397, 575, 641, 694, 695, 696, 699, 701 rotational, 641, 694, 696, 700, 701 Anti(Skew)-symmetric tensor, See Tensor Armstrong-Frederick kinematic hardening rule, 239, 307, 708 Associated flow rule (Associativity) Clausius-Duhem inequality, 708 Drucker’s interpretation, 228, 233 Ilyushin’s interpretation, 229, 232 Maximum plastic work, 228, 232 second-order plastic relaxation work rate, 229, 232, 233 Prager’s interpretation, 227 Associative law of vector, See Vector
Axial vector, See Vector B Back stress, See Kinematic hardening Bauschinger effect, 234 Bilateral damage, 460, 472, 477, 485 Bingham model, 437, 441 Biot, 123, 158 strain tensor, See Strain stress tensor, See Stress Body force, 151, 152, 156 Bounding surface model, 288 model with radial mapping, 259, 260, 288, 708 Bulk modulus, 268, 369, 372, 373, 429, 557 C Cam-clay model modified, 375, 376, 383, 388, 397, 401 original, 375, 376, 397 Cap model, 384, 387–389 Cartesian summation convention, 1 decomposition, See Tensor Cauchy ‘s first law of motion, See Equilibrium equation ’s fundamental theorem (’s stress principle), 145–147, 645 stress, See Stress elastic material, 208 Cauchy-Green deformation tensor, 119, 126, 159, 175, 190, 191, 505, 512, 517, 534, 543, 551 Cayley-Hamilton theorem, 41, 47, 48, 57, 58
© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 K. Hashiguchi, Foundations of Elastoplasticity: Subloading Surface Model, https://doi.org/10.1007/978-3-030-93138-4
841
842 Chaboche model, 273, 276, 277–279, 282, 283, 306, 334, 335, 344–346, 504, 706–708 Characteristic direction, See Principal direction equation, 36, 40, 42, 46, 48, 119, 568 length, 584 value, See Eigenvalue vector, See Eigenvector Circle of relative velocity, 137, 138 Closest point projection, 327, 527 Coaxial (Coaxiality), 48, 55, 123, 124 Cofactor, 4, 27, 29, 89 Commutative law of vector, See Vector Compliance method, 593 Component of tensor, See Tensor vector, See Vector Computer program (commercial) Abaqus, 109, 283, 345, 442, 454, 557, 641 LS-DYNA, 283, 288, 342, 642 Marc, 283, 319, 343, 454, 641, 705, 706, 741 PAM-STAMP, 342 Configuration 49 current (Eulerian), 83, 87, 121, 157, 597 initial, 83, 87, 95, 121, 157, 167, 233, 597, 598 intermediate, 192, 368, 501–514, 516–519, 522, 523, 525, 529, 531, 532, 538, 539, 543, 551, 596–600, 603, 608, 626 reference (Lagrangian), 83, 85, 88, 95–97, 103, 104, 106, 109, 111, 119, 126, 147, 149, 155, 157, 175, 178, 191, 502, 503, 505, 529, 551, 596 Conservation law of angular momentum, 145, 152, 155 linear momentum, 151 mass, 145, 151 Consistency condition conventional plasticity, 242, 253, 255, 259, 269, 273–275, 277, 281, 282, 309, 332, 342, 365, 387, 493, 605, 615 extended subloading surface model, 282, 288, 289, 290, 292, 293, 306, 309, 315, 319, 326, 327, 330, 393, 401, 403, 413, 419, 433, 444, 445, 457, 474, 499, 504 initial subloading surface model, 270, 272, 282, 288, 293, 309, 314, 326, 329, 383, 387, 393, 403, 413, 423, 424, 499, 623, 783 Consolidation of soils isotropic, 367 linear relation, 366, 370, 373, 403, 430, 726 normal, 367, 368, 370
Index swelling, 367, 368, 370 Contact elastic modulus for friction, 646 Contact traction for friction normal and tangential, 646 Continuity condition Prager’s, 245 Continuity equation, 151 Continuum spin, See Spin Contraction of tensor, See Tensor Contravariant and covariant base vector, 101, 107 Convected base vector, 107 coordinate system, 74, 95, 96, 98 stress rate contravariant, 180, 181 covariant, 150, 181 covariant-contravariant, 181 corotational, 181 tensor, 103, 145–150, 160 time-derivative, 104, 106, 108, 110, 171, 178, 186, 216 vector, 145, 146, 148, 644–647, 662, 695 Convective, See Convected Conventional plasticity model, 253, 259, 269, 273–275, 277, 281, 282, 309, 332, 342, 365 Conventional friction model, 642, 662, 663, 665, 667 Convective term, See Steady term Convexity condition of curve, 379, 731 Convexity of yield surface, 229, 230 Coordinate transformation Cartesian, 82 curvilinear, 74–76 Corner theory, 262 Corotational rate Green-Naghdi (Dienes), 567, 569 Zaremba-Jaumann, 566 Cosserat elastic material, 210 Cotter-Rivlin rate, 110, 112, 181, 182 Coulomb-Mohr critical state surface, 377 yield condition, 377 Coulomb sliding-yield condition, 662 Couple stress, 210 Covariant and contravariant, See Contravariant and covariant Covariant-contravariant convected stress, See Stress Creep model, 433, 438–440, 450, 595, 638, 639, 642, 683, 704, 705, 707, 708 Critical state, 374, 377, 379, 380, 383, 384, 386, 391
Index Cross product, See Vector Crystal plasticity crystal shear yield region, 605 creep-type crystal plasticity model, 637 critical shear stress, 605, 606, 613, 637, 704, 705 limit shear elastic-core region, 626 loading criterion, 638 multiplicative hyperelastic-based crystal plasticity, 595 normal-yield shear ratio, 613, 617 plastic shear strain rate, 604, 605, 613, 614, 616, 617, 621, 625 regularized Schmid law, 636, 637 resolved shear stress, 637, 704, 705 shear elastic-core yield region, 624–626 singular balue decomposition, 633, 635, 636 subloading-overstress-crystal plasticity model, 621 uniqueness of slip rate mode, 629 Curl of tensor field, See Tensor filed Current configuration, See Configuration Cutting plane projection, 327 Cyclic loading elastoplastic deformation, 440, 443, 705 friction, 242 Cyclic mobility, 371, 423 Cyclic plasticity model expansion of loading surface type, 276 extended subloading surface model, 288 subloading surface model, 276 kinematic hardening type, 278, 290 cylindrical yield surface model, 250, 273, 274, 276, 277, 290, 342, 440, 491, 706 –Chaboche model, 273, 276, 277, 278, 279, 282, 283, 306, 334, 335, 344, 345, 346, 504, 706, 707, 708 –Ohno-Wang model, 279 multi-surface model (Mroz model), 274, 276, 277, 283, 284, 286, 290 two(bounding)-surface model, 708 Cyclic stagnation of isotropic hardening, 321, 338, 363 Cylindrical yield surface (Chaboche) model, 273, 274, 276, 277, 281, 290, 342, 491, 706 D Damage anisotropic (orthotropic) damage model, 484
843 associated variable (strain energy density release rate), 472–474 bilateral model, 460, 472, 477, 485 brittle, 458, 473, 490 constitutive model, 486 ductile, 458, 486 Gurson model, 486 hypothesis of strain equivalences, 458 strain energy release rate, 472–474 stress transformation tensor, 462, 473, 479, 480, 485 stress triaxiality, 473 subloading damage model, 484, 491 subloading-overstress damage model, 484, 485 tensor anisotropic, 481 isotropic, 481 two-scale damage model, 491 unilateral model principal stress representation, 469, 473, 480, 485 stress transformation tensor, 462, 473, 479, 485 Deformation gradient elastic polar decomposition, 115, 516 plastic polar decomposition, 115, 516 polar decomposition, 87, 115, 117 relative, 119 Deformation theory Hencky, 263, 264 J2-, 263, 264 Description Eulerian, 85, 121 Lagrangian, 85, 130 material. See Lagrangian relative, 85, 119 spatial. See Eulerian total Lagrangian, 85 updated Lagrangian, 85, 130 Determinant product law, 5 Vandermonde’s, 56 Deviatoric part, 32, 44, 59, 203, 262, 268, 431, 532 plane, 58–60, 238 principal invariant, 42 projection tensor (fourth-order), 34, 204, 266 tangential projection tensor (fourth-order), 266
844 tangential stress rate, 262, 266 Diagonal component. See Tensor Partial differential calculi, 66 Dilatancy locking, 424 Directional distortional hardening, 240 Direction cosine, 8 Direct notation. See Tensor notation Director triad, 598 Discontinuity of velocity gradient, 589–591 Dissipation energy, 232, 241 Distributive law of vector, 7, 8 Divergence of tensor field. See Tensor field Divergence theorem. See Gauss’ s divergence theorem Drucker-Prager yield surface, 387, 388, 391 Drucker’s classification of plasticity model, 243, 273 postulate for stress cycle, 228 Dummy index, 2 Dyad. See Vector Dynamic-loading subloading-overstress model ratio, 447 surface, 442, 449, 544 Dynamic-loading for friction subloading-overstress sliding ratio, 685, 686, 689 sliding surface, 686 Dynamic recovery of kinematic hardening, 279, 281, 282, 476 E Eddington’s epsilon. See Permutation symbol Eigen (principal) direction. See Principal direction projection, 45 value, 36 value analysis, 591 vector, 36 Einstein’s summation convention. See Summation convention Elastic bulk modulus, 203, 369, 372, 373, 429 constitutive equation, 159, 189, 197, 200, 248, 431, 434, 435, 471, 527, 535 deformation gradient, 501, 506–508, 510, 516, 517, 538, 542, 596, 597, 501 hyper-, 501 modulus, 204, 205, 209, 214, 215, 224, 232, 267, 369, 435, 458, 460, 465, 467, 468, 490, 494, 602, 660 predictor step. See trial step shear modulus, 369, 420, 423, 430, 535 spin, 508
Index strain energy function, 192, 195, 203, 207, 301, 458, 459, 551, 558 strain energy function for elastic-core, 531 strain energy function for kinematic hardening, 530 strain energy function for rotational hardening, 400 strain rate, 214, 215, 226, 247, 248, 264, 289, 290, 359, 435, 459, 469, 477, 479, 485, 510, 602, 704 stress rate, 220, 226 tangent modulus, 434 trial step, 309, 311–313, 328 volumetric strain, 386, 725 Elastic-plastic transition, 253, 296, 331, 346, 365, 387, 391, 652, 703, 704 Elastoplastic stiffness modulus tensor, 220, 221, 388 Element test, 583, 584, 589 e – ln p linear relation. See Consolidation Embedded. See Convected Energetic elasticity, 549, 550 Entropic elasticity, 549, 550, 707 Equilibrium equation rate form, 148, 153, 155, 181 moment, 145, 155 Equivalent plastic strain, 221, 222, 237, 342 stress, 222, 349 viscoplastic strain, 437 Euler ’s first law of motion, 151, 153 ’s second law of motion, 152 ’s theorem for homogeneous function, 217, 721 Eulerian configuration. See Configuration description. See Description spin tensor. See Spin strain. See Strain tensor. See Tensor triad. See Triad Eulerian-Lagrangian two-point tensor, 96, 147 Explicit method. See Forward-Euler Extended subloading surface model. See Subloading surface model Eyring equation, 563, 739 F Failure surface, 382, 388 Finite strain theory, 211, 216, 228, 499, 625, 628 First Piola-Kirchhoff stress. See Stress Flow rule
Index associated, 218, 220, 221, 223, 227–229, 232, 235, 236, 251, 261, 269, 276, 288, 295, 362, 387, 388, 390, 391, 488, 495, 524, 587, 605, 613, 625, 636, 637, 654, 708 multiplicative plasticity, 506 non-associated, 220, 221, 223, 227, 387, 388, 391, 708 Flow rule for friction, 227 Footing settlement analysis, 371 Forward-Euler method, 113, 114, 304, 471, 477, 479, 480, 493, 548, 661, 705, 741 Friction coefficient evolution rule, 685 hyperelasticity, 189 kinetic, 641–643, 648, 649, 663–665, 667, 669, 687, 689, 704 negative-rate sensitivity, 683, 688, 704 positive-rate sensitivity, 683, 688, 690, 693 static, 641, 642, 648, 649, 663, 664, 669, 687–689, 704 Functional determinant. See Jacobian G Gauss’s divergence theorem, 72, 73, 92, 152, 153 Generalized Maxwell model, 554 Glass transition, 561 Glassy state, 561–563 Gradient of tensor field. See Tensor field Gradient theory, 584, 588 Green elastic equation. See Hyperelastic equation strain. See Strain Green-Naghdi rate stress rate, 181 Gurson model, 486, 490 H Hamilton operator. See Nabra Hardening isotropic, 217, 221, 255–258, 271, 273, 278, 296, 299, 300, 302, 319–327, 330–339, 344, 362, 363, 365, 368, 369, 383, 394, 403, 438, 447, 448, 453–455, 484, 494, 536–539, 542, 547, 585, 586, 606, 620, 621, 642, 647, 679, 706, 752 linear kinematic, 237, 569, 578 nonlinear-kinematic, 238–240, 278, 279, 281, 320, 326, 607
845 rotational, 235, 240, 271, 282, 397–401, 405, 409, 421, 423, 586, 696, 701 Hardening for friction, 648 Helmholtz free-energy function, 190 Hencky deformation theory, 263, 264 strain. See Strain Hooke’s law, 203, 214, 215, 268, 369, 459, 460, 482 Hyperelastic(-based) plasticity infinitesimal, 213, 214, 216, 522 multiplicative, 150, 215, 216, 302, 501, 502, 505, 522, 525, 529, 601, 603, 604, 608, 609, 611, 618, 705 Hyperelastic equation metals, 529, 530 soils, 430, 531, 532 Hypoplasticity hyperelastic equation, 213, 373, 532 infinitesimal strain for soils, 213, 373 multiplicative finite strain for soils, 215, 216, 373 Hysteresis loop, 260, 272, 284, 286, 288, 289, 290, 293, 296, 332, 335, 336, 338, 504, 623 I Identity tensor fourth-order, 34, 70 second-order, 34 Ill-posedness of solution, 584 Ilyushin’s isotropic stress space, 357, 358 postulate of strain cycle, 229, 232 Impact load, 442 Implicit method. See Backward-Euler method Indicial notation. See Tensor notation Infinitesimal strain. See Strain Infinite surface model. See Cyclic plasticity model Initial configuration. See Configuration Initial subloading surface model. See Subloading surface model Inner product. See Vector Intermediate configuration. See Configuration Internal variable, 104, 106, 109, 172, 177, 182, 183, 212, 214, 216, 220, 229, 233–235, 244, 246, 248, 262, 271, 309, 322, 395, 397, 405, 437, 447, 448, 509, 510, 516, 517, 519, 537, 551, 552, 565, 584, 586, 642, 705, 706, 712
846 Intersection of yield surfaces, 262 Invariant. See Principal invariant Inverse loading, 237, 289, 290, 305–308, 331, 454 Inverse tensor. See Tensor Isoclinic concept, 192, 501, 502, 516, 517, 538, 595, 596, 598 Isotropic hardening (variable), 217, 296, 321, 325, 394, 438, 484, 494, 537, 542, 547, 585, 620, 621 definition of isotropic material, 213 scalar-valued tensor function, 40 tensor-valued tensor function, 54, 57 traverse, 574, 575 J Jacobian, 84, 88, 366, 373 Jaumann. See Zaremba-Jaumann J2-deformation theory, 263 K Kinematically admissible velocity field. See Admissible field Kinematic hardening dynamic recovery, 238, 279–282, 307, 476 linear, 237, 569, 578 nonlinear, 238–240, 279, 281, 320, 326, 607 Prager. See linear rheological model, 549–551, 561, 562 variable (back stress), 212, 234, 235, 326, 362, 510 Ziegler, 237 Kinetic friction. See Friction Kirchhoff stress. See Stress Kronecker’ delta, 2, 5, 14, 23, 98 L Lagrangian configuration. See Configuration description. See Description spin tensor. See Description strain. See Strain tensor. See Tensor triad. See Triad Lame constants, 203 Laplacian (Laplace operator), 72 Lee decomposition. See Multiplicative decomposition Lie derivative, 106 Limit dynamic loading surface, 689 Limit dynamic loading surface for friction, 689
Index Linear kinematic hardening. See Kinematic hardening ln v ln p linear relation. See Consolidation Linear transformation, 20, 36 Liquefaction, 371, 423 Loading criterion plastic sliding velocity, 659 plastic strain rate, 659 Local form, 152, 153, 156, 157 Localization of deformation, 583 Local theory, 584 Local-time derivative. See Spatial-time derivative Local-time derivative term. See Non-steady term Lode angle, 60, 379, 388 Logarithmic corotational rate, 712, 713 spin, 182, 711, 712 strain, 125, 143, 144, 161, 371, 711 volumetric strain, 126, 143, 144, 366, 368 M Macauley bracket, 250, 463 Mandel stress. See Stress Masing rule, 283, 284, 306 Material description. See Lagrangian description frame-indifference. See Objectivity -time-derivative. See Time-derivative of volume integration, 92 Matrix representation of tensor relation, 715 Maxwell model, 435, 436, 439, 549, 552, 554, 557 Mean part of tensor. See Tensor Mechanical ratcheting effect, 332 Mesh size dependence (sensitivity), 584 Metric tensor, 14, 74, 76, 77, 79, 81, 96, 103, 505, 511, 551 Mises ellipse, 358 yield condition, 221, 223, 263, 283, 319, 356–358, 437, 486, 578 plane strain, 359 plane stress, 357 Modified Cam-clay model, 375, 376, 397, 401 Mohr’s circle, 62, 64, 360 Momentum linear, 151 angular, 145, 152, 155 Motion, 15, 18, 19, 83–85, 151, 152, 641, 669, 671 Mroz model, 273, 276, 277, 283, 708
Index Mullins effect, 557, 560 Multikinematic model, 278 Multiplicative decomposition crystal plasticity, 596 deformation gradient, 502, 510, 511 elastic-core, 510, 511 kinematic hardening, 510, 517 plastic deformation gradient, 501, 510 viscoplastic deformation gradient, 542 Multiplicative hyperelastic-based plasticity, 150, 215, 216, 302, 501, 502, 505, 522, 525, 529, 601, 603, 604, 608, 609, 611, 618, 705 Multi surface (Mroz) model. See Cyclic plasticity model N Nabra, 70 Nanson’s formula, 89 Natural strain. See Logarithmic strain Navier’s equation, 208 Negative transformation, 85 Nominal strain. See Strain stress. See Stress stress rate. See Stress rate stress vector. See Stress vector Nonassociated flow rule (Non-associativity), 390 Nonhardening. See Stagnation of isotropic hardening Nonlinear kinematic hardening. See Kinematic hardening Non-local theory, 584 Non-proportional loading, 262, 305, 331, 332, 338, 419, 704 Non-singular tensor. See Tensor Non-steady term, 86 Normal component, 61, 64, 139, 223, 233, 266 Normal isotropic hardening ratio, 322, 323–325, 332–336, 338, 339 surface, 321–325, 332, 335, 336, 338, 339, 363 Normality rule. See Associated flow rule Normalized orthonormal base, 9 Normal-sliding for friction ratio, 644 evolution rule, 648 surface, 643 Normal stress rate, 262, 593 Normal-yield ratio, 648, 651, 652, 661, 685 evolution rule, 651, 665, 666
847 surface, 648, 652, 653, 663, 666, 674, 675, 679 Norton law, 438 O Objective rate of tensor, 172, 177 rate of vector, 177, 183 time-derivative of scalar-valued tensor function, 183 time-integration of rate tensor, 216 stress rate tensor, 180, 182 tensor, 19 transformation, 19, 171, 176, 177, 183 Objectivity, 95, 106, 171, 172, 175–179, 191, 216, 233, 243, 440, 565, 642 Octahedral plane. See deviatoric plane shear stress, 223 Oldroyd rate stress rate, 180, 182 Original Cam-clay model, 375, 376 Orthogonal coordinate system, 8, 15, 635 tensor, 22, 23, 39, 52, 55, 62, 116, 183, 634 Orthotropic anisotropy, 346, 349, 355, 360, 363, 397, 694–696, 699, 701 Orthotropic anisotropy for friction, 694, 696, 699, 701 Over stress, 269, 433–435, 437–443, 445, 447–449, 451, 452, 454, 484, 485, 542, 545, 547–549, 561, 563, 619, 621, 638, 683–686, 688–690, 692, 704, 705, 707, 793 Overstress friction model. See Subloading-overstress friction model Overstress model Bingham model, 437, 440, 441 Perzyna model, 437, 440, 445 Prager model, 437, 440, 445 return-mapping, 688 subloading overstress model, 442, 443, 447–449, 451, 452, 454, 485, 542, 545, 547, 549, 561, 684, 692, 704, 705, 793 P Parallelogram law of vector, 7 Partial differential calculi, 66 Perfectly-plastic material, 244, 253 Permutation symbol, 2, 5 Phase-transformation heat-transformation strain increment, 496 thermal and transformation strain increment, 496
848 Piola-Kirchhoff stress. See Stress p-plane. See Deviatoric plane Plastic corrector step, 309, 312, 537, 186 deformation gradient. See Deformation gradient material spin, 108, 114 flow rule. See Flow rule modulus, 218, 224, 252, 255, 258, 283, 285, 286, 303, 320, 326, 383, 386, 389, 393, 404, 408, 425, 436, 448, 476, 527, 548, 586, 662 multiplier. See Positive proportionality factor potential, 218, 234, 387–390, 647, 708 relaxation modulus tensor, 221 relaxation stress rate, 220, 225 shakedown. See Shakedown spin, 108, 114 strain rate, 211, 213, 215, 218–220, 223, 227–229, 232, 233, 240, 243, 244, 247–250, 252–255, 262–264, 271, 272, 276–278, 280, 282, 284–286, 288–290, 302–304, 306, 309, 321, 323, 325, 332, 342, 344, 351, 359, 361, 362, 383, 387, 388, 399, 401, 403, 407, 408, 434, 436–439, 442–448, 474, 476–479, 483, 484, 486, 491, 509, 516, 524, 527, 544, 547, 549, 579, 580, 587, 588, 637, 659, 684, 685, 703–705, 708 volumetric strain, 386, 401, 423, 490 Plastic for friction corrector step, 309, 312, 537 flow rule, 218, 408, 506, 509, 524, 537, 707 modulus, 218, 224, 246, 252, 255, 258, 283, 285, 286, 303, 320, 326, 383, 386, 389, 393, 404, 408, 425, 436, 448, 476, 527, 548, 586, 662 multiplier, 218, 268, 303, 445, 446, 448, 449, 455, 477, 479, 480, 489, 509, 524, 529, 538, 546, 548, 562, 577, 587, 588, 610, 617, 636, 637, 656 relaxation traction rate, 591, 593 sliding rate, 656, 663, 685, 701 Poisson’s ratio, 204, 205, 331, 369, 393, 455, 459, 460, 481, 494 Polar decomposition, 51, 53, 87, 115–117, 119, 516 spin. See Spin Polymer glassy state, 562 viscoelastic rheology model, 549, 550
Index Positive definite tensor, 51, 224, 633, 660 Positive proportionality factor, 218, 307, 445 Positive transformation, 84, 85 Positive definite tensor, 51, 224, 633, 660 Prager’s continuity condition, 243–245, 296 interpretation of associated flow rule, 227 linear kinematic hardening rule, 237, 569, 578 overstress model, 437 Prandtl plasticity model, 435, 436, 441 Prandtl-Reuss equation, 222 Primary base vector, 14, 74, 76, 97, 101, 598 vector, 12, 74 Principal direction, 38, 39, 48, 53, 63, 116, 118, 125, 137, 138, 142, 143, 186, 483, 485, 567, 573, 695 invariant, 28, 41, 42, 48, 49, 186 time-derivative, 186 space, 58 stretch, 41, 54, 116, 118, 119, 123, 124, 128, 194, 557 value. See Eigenvalue vector. See Eigenvector Principle of objectivity. See Objectivity material-frame indifference. See Objectivity maximum plastic work, 228, 232 Product law of determinant, 5 Projection of area, 709 Projection of tensor deviatoric, 34, 204, 266, 315, 717 deviatoric-tangential, 266 Proper value. See Eigenvalue Pull-back operation, 95, 100, 101, 103, 104, 106, 109, 121 Pulsating loading, 276, 281, 332, 333, 343, 452, 638, 639 Push-forward operation, 95, 99, 100–104, 106, 109, 121, 178, 197 Q Quasi-static deformation, 433, 435, 438, 439, 441, 444, 447, 449, 548, 638, 707 Quotient law, 19, 120, 145 R Ratcheting effect. See Mechanical ratcheting effect Rate of elongation, 137, 139
Index shear strain, 139, 567, 604, 605, 611, 613, 614, 616, 617, 621, 625, 632, 633, 636, 705 surface area, 41 volume, 386 Rate-type equilibrium equation, 145, 148, 152, 153, 155, 157, 181, 424 virtual work principle, 145, 156 Reciprocal base vector, 14, 74, 76, 77, 95, 96, 98, 505, 598 vector, 11, 12, 76 Reference configuration. See Configuration Regularized Schmid law, 636, 637 Relative deformation gradient, 119 description, 85, 119 left and right Cauchy-Green deformation tensor, 119 spin, 107, 132, 175, 567, 569, 571 Reloading behavior in subloading surface model, 282, 306 Rate of elongation, 137, 139 Representation theorem, 57 Residual stress analysis, 342, 344 Return-mapping closet point, 527 cutting plane, 327 elastic trial (predictor) step, 310 plastic corrector step, 537 Return-mapping for friction model, 688 Reynolds’ transportation theorem, 93, 151 Rigid-body rotation, 104, 107–109, 113, 120, 136, 138, 139, 171, 172, 175, 176, 183, 185, 190, 215, 501, 508, 516, 517, 538, 596, 598, 699, 713 Rigid-plastic material, 228 Rotation of tensor field. See Tensor field -free(insensitive) tensor, 178 rate tensor of material, 131 Rotational anisotropy of friction, 641, 695, 700, 701 Rotational hardening evolution rule, 212, 235, 237, 238, 251, 256, 258, 259, 262, 271, 280, 296, 299, 300, 306, 308, 322–325, 327, 362, 389, 399, 400, 403, 405, 407, 421, 438, 444, 445, 472, 476, 486, 495, 499, 525, 545, 558, 559, 587, 607, 614, 620, 626, 648, 649, 651, 652, 665, 666 rheological model, 549, 550, 551, 561, 562
849 Rotational strain tensor, 120, 121, 123–125, 129, 131, 139–142, 175, 191, 192, 201, 208, 213, 430, 535, 706, 711 S Scalar product. See Vector triple product. See Vector Second-order work rate elastic stress, 220, 226, 229, 230, 306, 469, 717 Second Piola-Kirchhoff stress. See Stress Shear band inception, 589 band thickness, 583, 584, 587, 588 -band embedded model, 588 Shear modulus, 369, 420, 423, 430, 535 Shear strain rate, 139, 567, 604, 605, 611, 613, 614, 616, 617, 621, 625, 632, 633, 635, 636, 705 Similarity-center enclosing condition, 297–299, 406 surface, 289, 290, 293, 295, 306, 332, 335, 338, 393, 499, 624 translation rule, 285 yield ratio, 306 Similarity-ratio, 362 Similar tensor, 39, 53, 116 Simple shear, 113, 164, 165, 564, 565, 567, 569, 572, 573, 578–580, 598, 599, 713 Single surface model. See Cyclic plasticity model Skew-symmetric tensor. See Tensor Slidinghardening function, 648–650, 675, 679–681 subloading surface, 648, 649, 654, 696 yield condition, 647, 662, 665 Sliding velocity elastic and plastic, 683 normal and tangential, 265, 646 Small single surface model. See Cyclic plasticity model Smeared crack model. See Shear-band embedded model Smoothness condition, 243, 245–247, 250, 266, 269–271, 276, 290, 296, 703, 704 Spatial description. See Eulerian description Spatial-time derivative, 85, 86 Spectral representation, 39, 43, 66 Spherical part of tensor. See Tensor Spin base vector, 8, 9, 14, 15, 22, 38, 59, 74, 76, 77, 82, 87, 95–103, 106, 107, 146, 148, 149, 178, 504, 510, 598, 617, 622, 701
850 Spin (cont.) continuum, 91, 107, 108, 114, 131, 135, 137, 175, 177, 183, 508, 509, 567, 569, 713 elastic, 508 Eulerian, 133, 135, 711, 712 Lagrangian, 133, 135 plastic, 209, 228, 244, 253, 501, 569, 579, 638 plastic material, 209, 228, 244, 253, 501, 569, 579, 638 relative (polar), 107, 109, 132, 175, 567, 569, 571 substructure, 108, 508, 509, 526, 574 tensor, 33, 91, 108, 114, 129, 131, 133, 175, 177, 180, 506, 507, 526 Springback analysis, 342, 343, 706 Stagnation of isotropic hardening in soils, 321 Statically-admissible stress field. See Admissible field Static friction. See Friction Steady term, 00 Stick-slip phenomenon, 641, 669–671 Strain Almansi (Eulerian), 120, 121, 123, 139, 141, 175, 208 Biot, 123 energy function, 183, 186, 189, 191, 192, 194–196, 198–200, 203, 204, 207, 208, 213–215, 301, 430, 435, 458, 459, 461, 472–474, 518, 524, 530–532, 534, 551, 552, 555, 557, 558 Green (Lagrangian), 121 Hencky, 125, 131, 142, 182 infinitesimal, 121, 122, 142, 163, 202, 203, 208, 211, 213, 215, 216, 230, 232, 369, 373, 525, 532, 534, 535, 538, 552 logarithmic (natural), 125, 143, 144, 161, 371, 711 volume, 126, 143, 144, 366, 368 nominal, 121, 141–144, 161, 371, 372 volume, 144, 371–373 Strain space theory, 226, 321 Strain rate elastic and plastic, 373 intermediate configuration, 507, 509 viscoplastic, 434–439, 442–448, 484, 486, 544, 547, 549, 684, 685 Stress Biot, 158 Cauchy, 113, 145–149, 153, 155, 160, 176, 177, 180–182, 194, 202, 208, 209, 213, 557, 576, 577, 612, 623, 645, 712 covariant-contravariant convected, 181, 182
Index first Piola-Kirchhoff (nominal), 147, 148, 159, 160, 176 Kirchhoff, 146–149, 159–161, 176, 180–182, 190–192, 198, 200, 201, 233, 517–519, 534, 552, 556, 561, 604, 617, 622 Mandel, 149, 150, 159, 192, 197, 198, 200, 508, 516, 518, 524, 526, 533, 538, 553, 603, 612, 623 Nominal, 146, 148, 153, 161, 181, 182 second Piola-Kirchhoff stress, 148, 149, 159, 160, 176, 190, 192, 200, 233, 517, 552, 556, 561 pull-backed to intermediate configuration, 102, 106 Stress-controlling function subloading-friction model, 653 subloading surface model, 254, 256 Stress rate Cotter-Rivlin, 110, 112, 181, 182 Green-Naghdi, 107, 110, 112–114, 181, 182, 567, 569, 571–573, 579 Jaumann, 153 nominal, 146, 148, 153, 181, 182 Oldroyd, 180, 182 Truesdell, 57, 106, 159, 172, 180, 182, 192, 201, 209, 213, 215, 216, 263 Zaremba-Jaumann, 153 Stress space theory, 219, 234, 235, 266, 273, 315, 356–358, 394, 397–400, 425, 529, 639, 675 Stress vector nominal, 146, 148 Stretch left and right, 118 principal, 41, 54, 116, 118, 119, 123, 124, 128, 194, 557 Stretching. See Strain rate Subloading crystal plasticity model, 613, 639 Subloading-damage model, 484, 491 Subloading-friction model normal-sliding ratio, 651, 661, 663 normal-sliding surface, 663, 666 return-mapping, 310, 311, 314 sliding-subloading surface, 648, 649 stress-controlling function, 253, 256, 653 Subloading-overstress friction model, 683, 684, 686 Subloading-overstress model dynamic-loading ratio, 686, 689 dynamic-loading surface, 449, 544, 686, 689, 706 Subloading phase transformation model, 493, 499
Index Subloading surface model extended, 282, 288–290, 292, 293, 306, 309, 315, 319, 326, 327, 330, 393, 401, 403, 413, 419, 433, 444, 445, 457, 474, 499, 504 initial, 270, 272, 282, 288, 293, 309, 314, 326, 329, 383, 387, 393, 403, 413, 423, 424, 499, 623, 783 kinematic hardening, 260, 706 metals, 529 normal-yield ratio, 293, 295 return-mapping, 310, 311, 314, 479, 688 stress-controlling function, 254, 256, 653 soils, 529 subloading surface, 645 tangential-inelastic strain rate, 262–270, 315, 316, 327, 330, 331, 419, 704, 752 Substructure spin. See Spin Subyield state, 249, 250, 331, 614, 615 Summation convention, 1, 2 Superposed kinematic hardening slender Mises model, 296 Superposition of rigid-body rotation, 176 Surface element, 80, 88, 91, 146 Sylvester’s formula, 46 Symbolic notation. See Tensor notation Symmetric tensor. See Tensor Symmetry of Cauchy stress, 155 T Tangential inelastic strain rate, 244, 262–270, 315, 316, 327, 330, 331, 419, 425, 704, 752 stress rate, 262, 265, 266 Tangential for friction associated flow rule of friction, 654 contact traction, 663, 665 sliding velocity, 666, 679, 682 Tangent (stiffness) modulus, 192, 232, 247, 255, 276, 296, 493, 646, 703 Tension and distortion behavior, 167 Tension cut of yield surface, 388 Tensor acoustic, 591 anti-symmetric. See skew-symmetric Cartesian decomposition , 31 characteristic equation, 36, 46 coaxiality, 48, 55 component, 79 contraction, 20, 21 damage, 481, 483
851 definition, 18, 19 deviatoric part, 32, 44, 59 deviatoric projection (fourth-order), 34, 204, 266, 315, 717 diagonal component. See normal component direct notation. See symbolic notation eigenprojection, 45 eigenvalue, 39, 44, 46, 48 eigenvector, 39, 43, 46–48, 54, 55, 633 Eulerian, 95, 99, 101–103, 150, 178 identity (second- and fourth-order), 21, 24, 25, 34, 70, 177, 180, 204, 220, 246, 266, 315, 347, 553, 716 indicial notation, 20 inverse, 29, 30, 36, 77, 116, 165 invertible (non-singular), 30, 31 Lagrangian, 99, 101–103, 148, 178 magnitude, 26, 81 matrix notation, 20, 21 mean part. See spherical part non-singular, 31 normal component, 61, 266 notation, 37 objective (transformation), 19, 171, 176, 177, 183 orthogonal, 22, 23, 39, 52, 55, 62, 116, 183, 634 partial derivative, 66, 127, 192 polar decomposition, 51, 516 positive definite, 51, 224, 633, 660 principal direction, 47, 48, 53, 142, 186, 485 principal invariant. See Principal invariant principal value. See Eigenvalue product. See Vector representation in principal space, 58–61 rotation-free(insensitive), 178 shear component, 22 similar. See Similar tensor skew-symmetric (anti-symmetric), 31, 32, 46, 47, 136 skew-symmetrizing (fourth-order), 34, 70, 553 spectral representation, 39, 43, 66 spherical part, 149 spin, 33, 91, 108, 114, 129, 131, 133, 175, 177, 180, 506, 507, 526 strain. See Strain strain rate. See Strain rate stress. See Stress
852 Tensor (cont.) stress rate. See Stress rate symbolic notation, 21, 177, 179 symmetric, 37, 39, 43, 51, 55, 57, 66, 70, 82, 101, 116, 146, 185, 221, 519, 633, 715 symmetrizing (fourth-order), 34, 70, 553 time-derivative. See Time derivative trace, 25 tracing (fourth-order), 204 transpose, 21, 25, 81, 82 triple decomposition, 32 two-dimensional state, 172 two-point, 96, 147, 148, 173, 176 Tensor field curl. See rotation divergence, 71 gradient, 70 rotation (curl), 71 Tensor notation direct. See symbolic indicial, 20 matrix, 37 symbolic, 20, 21 Time derivative corotational, 180, 183, 185, 186 local (spatial-time), 85, 86 material, 85, 86, 92, 93, 104–106, 108, 139, 140, 144, 171, 172, 176, 177, 183, 185, 186, 211, 217, 235, 260, 297, 302, 404, 475, 524, 565, 579, 585, 586, 602, 605, 613, 627 moment of tensor, 92 non-steady (local time derivative) term, 86 principal invariant, 186 scalar-valued tensor function, 183 steady (convective) term, 104, 106, 110, 171, 178 Total Lagrangian description. See Description Trace, 25, 41, 81, 203, 212, 366, 371 Transformation negative, 85 objective, 19, 171, 176, 177, 183 phase, 293, 493, 496, 497, 499, 704 positive, 84, 85 tensor, 20 Transportation theorem. See Reynolds’ transportation theorem Transpose, 3 Traction. See Stress vector Traverse isotropy, 574, 575 Triad Eulerian, 116, 118, 133, 712 Lagrangian, 116, 118, 133
Index Triple decomposition, 32 Truesdell stress rate. See Stress rate True strain, 121, 125, 161 True stress, 148, 161 Two-dimensional state, 15, 61, 172, 173 Two-point identity tensor. See Tensor Two-point tensor, 96, 147, 148, 173, 176 Two surface model. See Cyclic plasticity model U Unconventional plasticity, 273–275, 282, 332 Uniaxial loading behavior, 256, 257, 278, 286, 289, 300, 450, 451 Uniqueness of solution, 244 Uniqueness of crystal slip rate model, 629–632 Updated Lagrangian description. See Description V Vandermonde’s determinant, 56 Vector associative law, 7, 8 axial, 32, 33, 47, 132, 136 commutative law, 10, 55 component description, 179 cross product. See tensor product definition, 7 distributive law, 7, 8 dyad. See tensor product eigen, 35–39, 43, 45–48, 51, 54, 55, 58, 633, 634 equivalence, 7 inner product. See scalar product magnitude, 7, 8, 16 parallelogram law, 7 principal. See eigen product, 9, 11, 13, 28, 78 scalar product, 7–9, 11, 13, 38, 72, 80, 137, 709 scalar triple product, 9 tensor product, 11, 74, 79, 80 Velocity gradient discontinuity, 589–591 Vertex theory. See Corner theory Virtual work principle, 145, 156 Viscoelastic model rheology model, 549–550 damage model: Mullins effect, 557–561 Viscoplastic friction model, 689 Viscoplastic model creep model, 438, 595 Perzyna model, vi
Index subloading-overstress model, 442, 448, 449, 561 Void volume fraction, 486, 487, 490 Voigt representation subloading-overstress friction model, 719 Volume element, 86, 88, 89, 93, 151, 491 Volumetric strain. See Strain Volumetric strain rate elastic and plastic, 368, 373 Vorticity, 137 W Work conjugacy, 157, 159, 240 conjugate pair, 145, 158–161, 190, 192, 210, 240, 526 hardening, 222 rate (stress power), 157, 159, 160, 192, 222, 227–229, 231–233 Y Yield condition (surface)
853 Cam-clay tensile strength, 393 Cap-model, 387 Coulomb-Mohr, 377 directional distortional, 180, 240 Drucker-Prager, 381 Hill, 349, 354 Mises, 221, 223, 263, 283, 319, 356, 357, 358, 437, 486, 578 orthotropic, 349, 354 orthotropic anisotropy, 349 Yoshida-Uemori model, 343, 344 Young's modulus, 160, 161, 204, 205, 287, 331, 342, 343, 454, 455, 459, 460, 468, 494 Z Zaremba-Jaumann rate Almansi strain, 141 Cauchy stress, 182, 202, 209, 576, 577 Kirchhoff stress, 182, 604, 617, 622 Ziegler's kinematic hardening rule, 237