Superalloys : Production, Properties and Applications [1 ed.] 9781622570355, 9781612095363

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Copyright © 2011. Nova Science Publishers, Incorporated. All rights reserved. Superalloys : Production, Properties and Applications, Nova Science Publishers, Incorporated, 2011. ProQuest Ebook Central,

Copyright © 2011. Nova Science Publishers, Incorporated. All rights reserved. Superalloys : Production, Properties and Applications, Nova Science Publishers, Incorporated, 2011. ProQuest Ebook Central,

MATERIALS SCIENCE AND TECHNOLOGIES

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SUPERALLOYS: PRODUCTION, PROPERTIES AND APPLICATIONS

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SUPERALLOYS: PRODUCTION, PROPERTIES AND APPLICATIONS

JEREMY E. WATSON EDITOR

Nova Science Publishers, Inc. New York Superalloys : Production, Properties and Applications, Nova Science Publishers, Incorporated, 2011. ProQuest Ebook Central,

Copyright © 2011 by Nova Science Publishers, Inc. All rights reserved. No part of this book may be reproduced, stored in a retrieval system or transmitted in any form or by any means: electronic, electrostatic, magnetic, tape, mechanical photocopying, recording or otherwise without the written permission of the Publisher. For permission to use material from this book please contact us: Telephone 631-231-7269; Fax 631-231-8175 Web Site: http://www.novapublishers.com NOTICE TO THE READER The Publisher has taken reasonable care in the preparation of this book, but makes no expressed or implied warranty of any kind and assumes no responsibility for any errors or omissions. No liability is assumed for incidental or consequential damages in connection with or arising out of information contained in this book. The Publisher shall not be liable for any special, consequential, or exemplary damages resulting, in whole or in part, from the readers‟ use of, or reliance upon, this material. Any parts of this book based on government reports are so indicated and copyright is claimed for those parts to the extent applicable to compilations of such works.

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Independent verification should be sought for any data, advice or recommendations contained in this book. In addition, no responsibility is assumed by the publisher for any injury and/or damage to persons or property arising from any methods, products, instructions, ideas or otherwise contained in this publication. This publication is designed to provide accurate and authoritative information with regard to the subject matter covered herein. It is sold with the clear understanding that the Publisher is not engaged in rendering legal or any other professional services. If legal or any other expert assistance is required, the services of a competent person should be sought. FROM A DECLARATION OF PARTICIPANTS JOINTLY ADOPTED BY A COMMITTEE OF THE AMERICAN BAR ASSOCIATION AND A COMMITTEE OF PUBLISHERS. Additional color graphics may be available in the e-book version of this book. Library of Congress Cataloging-in-Publication Data Superalloys : production, properties, and applications / [edited by] Jeremy E. Watson. p. cm. Includes bibliographical references and index. ISBN 978-1-62257-035-5 (E-Book) 1. Heat resistant alloys. I. Watson, Jeremy E. TA485.S94 2011 546'.3--dc22 2011003586 ISBN 978-1-61209-536-3

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CONTENTS Preface Chapter 1

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Chapter 2

Chapter 3

Chapter 4

Chapter 5

Chapter 6

vii Effect of Residual Stresses on Crack Shape of Corner Cracks at Holes in Nickel Base Superalloys R. Branco, F. V. Antunes, J. A. Martins Ferreira and J. M. Silva Wrought Superalloys – Science, Technology and Applications M. Nageswara Rao High-Temperature Oxidation Behavior of the Nickel-Base Single Crystal Superalloy and Its Aluminide Diffusion Coating C. T. Liu and J. Ma Low Cycle Fatigue Characterization of Nickel-Base Aeronautical Superalloys S. Chiozzi, V. Dattoma, M. Di Castri and R. Nobile Mechanical Property and Characterization of a NiCoCrAlY Type Metallic Bond Coat Used in Turbine Blade Made of AE-437A Ni Base Superalloy Ashok Kumar Ray Control of Microstructures by Heat Treatments and High-Temperature Properties in High-Tungsten Cobalt-Base Superalloys Manabu Tanaka and Ryuichi Kato

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1

25

45

61

73

97

vi Chapter 7

Contents Development, Properties, Applications and Machinability of Superalloy Inconel Chandra Nath, Kapil Krishna Baidya and Mustafizur Rahman

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Index

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115

159

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PREFACE This new book presents current research in the study of the production, properties and applications of superalloys. Topics discussed include the effect of residual stresses on crack shape of corner cracks at holes in nickel base superalloys; the application of wrought superalloys; high-temperature oxidation behavior of the nickel-base single crystal superalloy and its aluminide diffusion coating and the properties and machinability of superalloy Inconel. Chapter 1 - The main objective of this paper is to study the crack shape evolution in double-U and central hole specimens with corner cracks, both representative of gas turbine discs in terms of critical zones with stress concentrations. An automatic crack growth technique is employed, consisting of a three iterative steps: 3D finite element analysis, K calculation and crack propagation. Complementary experimental work was developed in RR1000, a nickel base superalloy applied in turbine disks, in order to obtain crack shapes and fatigue crack growth at elevated temperature. The effect of residual stresses on crack shape is investigated. Chapter 2 - Superalloys are heat-resistant alloys based on iron, nickel-iron, nickel or cobalt, exhibiting high creep and rupture strength at elevated temperatures. A number of both wrought (produced through ingot metallurgy route) and cast grades of superalloys have been developed over the years. There is also a group of superalloys produced through powder metallurgy route. This article focuses on the metallurgy, high temperature mechanical properties and applications of conventional wrought superalloys based on iron, nickel and cobalt. A number of wrought superalloy grades have become commercially available over the last seven decades; the emphasis was on the development of alloys with higher and higher creep resistance. A noteworthy feature of

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Jeremy E. Watson

wrought nickel base superalloys available today is their heat-resistance upto 80% of their incipient melting temperature, a fraction that is higher than for any other class of engineering alloys. There have been a number of processing discoveries which helped in development of wrought superalloys with higher and higher performance levels. The strengthening mechanisms that come into play in different groups of wrought superalloys have been well understood. Certain trace elements play a major role in superalloys and there have been a number of studies dealing with underlying mechanisms. Considerable progress has been achieved with reference to understanding the metallurgical stability of wrought superalloys under high temperature loading conditions and its relation to chemical composition. While non-aerospace applications have been there, prominently in the power generation sector, aerospace accounts for the major part of the consumption of the wrought superalloys. Developments in wrought nickel-iron and nickel base superalloys over the decades have importantly contributed to the evolution of modern high performance jet engines. Chapter 3 - Isothermal oxidation tests on the Ni-base single crystal superalloy and its aluminide coating were performed in the temperature range of 900-1000°C to describe the high - temperature oxide scales formed and changed then to put forward their oxidation mechanisms, with morphological observations and chemical analysis of the oxide-scale surface and crosssectional microstructure by scanning-electron microscope (SEM) coupled with energy-dispersive X-ray analyzer (EDX) and analysis of crystal structure by X-ray diffraction (XRD). It was found that the better oxidation behavior and more excellent resistance to spallation were shown by a laboratory version of the aluminide coating than that of the Ni-base single crystal superalloy. Chapter 4 - Superalloys are the most diffused material for aeronautical and aerospace applications, mainly for turbines and compressors production, because of their excellent resistance at high temperature. Among them, polycrystalline superalloys Udimet 720Li and Inconel 718Plus are interesting candidates for turbine disc application, due to their high temperature strength, good corrosion resistance and excellent workability. The aim of the present work is to study the low cycle fatigue (LCF) behaviour of these two nickel-base superalloys for aeronautical applications. The experimental tests plan has been predisposed with two different working temperatures (650 and 700°C). Materials were provided from Avio s.p.a. Torino in the form of manufactured article. This project includes tensile and creep tests, too.

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Preface

ix

Chapter 5 - This work highlights some of the results obtained while studying Ni20Co18Cr12.5Al0.6Y (NiCoCrAlY) type metallic bond-coat properties of a thermal-barrier coated (TBC), AE-437A Ni base superalloy mostly employed for manufacturing compressor and stationary stator blades in aero turbines. Experiments were mainly focused in the area of evaluation of microstructure, residual stress, shear strength, hardness and with special emphasis in establishing the ductile to brittle transition temperature (DBTT) of the bond coat by using acoustic emission technique during room temperature and high temperature tensile tests. Results reveal that the residual stress was tensile in nature in the TBC layer and compressive in the bond coat as well as in the substrate. The DBTT of this bond coat is around 650°C, which is in close proximity to the value reported in literature for CoCrAlY type of bond coat. Finite element technique was used to analyze the equivalent stresses in the bond coat, the result of which revealed the highest order of equivalent stress 800°C, as the bond coat is ductile above 650°C. Shear strength of the bond coat is in close proximity with that of the bond strength reported in literature for CoCrAlY and Ni22Co17Cr12.5Al0.6Y types of bond coat. Chapter 6 - Cobalt-base HS-21 (L-605) alloy has high strength and good oxidation resistance at high temperatures, and is widely used for hightemperature components including blades, vanes and combustor parts in the hot sections of jet engines. In this study, effects of microstructures on the creep-rupture properties were investigated on the heat-treated specimens of the HS-25 (L-605) type heat-resistant alloys containing about 14 to 20% (mass %) tungsten (W) at 1089 and 1311 K. Serrated grain boundaries which were formed by precipitation of W-rich phase and M6C carbide by heat treatment, improved rupture strength without significant loss of creep ductility. Ageing for 1080 ks (300 h) at 1273 K (1000°C) caused similar precipitates on grain boundaries and in grains, and also increased rupture strength in the specimens with normal straight grain boundaries. Improvement of the rupture properties by heat treatments was remarkable in the alloys with the higher W content at 1089 K, while such heat treatments were effective in relatively short-term creep at 1311 K. In the non-aged specimens with straight grain boundaries, the rupture strength increased with increasing W content at 1311 K, although the rupture strength was not improved largely with increasing W content at 1089 K. The principal strengthening mechanism in these alloys was attributed to the strengthening of grain boundaries and grains by precipitates of W-rich phase and carbide phases in addition to solid-solution strengthening by W atoms. Fracture surfaces of specimens with serrated grain boundaries and those of aged specimens were ductile grain-boundary fracture surfaces with small

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dimples and ledges, while the non-aged specimens with straight grain boundaries exhibited brittle grain-boundary facets at 1089 K. Chapter 7 - Ni-based alloys contain Cr, Mo, Fe, Cu, Al, Ti, Co etc. as alloying elements and possess unique and versatile mechanical and chemical properties. So, these alloys are widely used for different impact-resistant and high-temperature applications such as aircraft engines, gas and steam turbines, pumps, rocket motors, nuclear reactors and petrochemical industries, etc. However, conventional machining of these high-temperature superalloys are cumbersome as they cause tool blunting, unusual and rapid tool wear, workhardening, chatter vibration and high cutting temperatures and so on, which result in cutting instability and poor machinability. This study provides an overview on development, properties and applications of Ni-based high temperature alloys (HTAs). Then, it presents a review of research findings on machinability of these superalloys while applied conventional turning (CT), conventional milling (CM), high speed machining (HSM), hybrid machining and ultrasonic vibration cutting (UVC) techniques using mainly coated and uncoated carbide, cermets, ceramics, Sialons, SiC WRA (whisker reinforced alumina) and CBN (cubic boron nitride) tools. It has been observed that the properties of Ni-based superalloys not only depend on various alloying elements but also on their percentages and combinations, microstructures, strengthening mechanisms and heat treatment procedures during formation of metallic structure or phase(s). Among the Ni-based HTAs, Inconel 718 is most widely applied and thereby it has drawn much attention by the researchers and manufacturers. It has also been observed that carbide tools are frequently used to machine the Ni-based HTAs as they are comparatively inexpensive. However, the CBN tools perform always better than the carbide tools in all aspects. The HSM method offers higher productivity than the other methods but it requires special machining arrangements, while the UVC technique offers better surface finish (e.g. < 1 μm Ra) and longer tool life but lower productivity. At higher cutting speeds, the cutting tools chipped off or catastrophically failed in the UVC method. It has been observed through this study that the CT, CM, HSM, hybrid machining and UVC techniques can be applied for machining superalloy Inconel; however, cutting performances and productivity in each method depend on specific cutting conditions.

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In: Superalloys: Production, Properties… ISBN: 978-1-61209-536-3 Editor: Jeremy E. Watson, pp. 1-24 © 2011 Nova Science Publishers, Inc.

Chapter 1

EFFECT OF RESIDUAL STRESSES ON CRACK SHAPE OF CORNER CRACKS AT HOLES IN NICKEL BASE SUPERALLOYS R. Branco1*, F.V. Antunes2, J.A. Martins Ferreira2 and J.M. Silva3 Copyright © 2011. Nova Science Publishers, Incorporated. All rights reserved.

1

Department of Mechanical Engineering, ISEC, Polytechnic Institute of Coimbra, Portugal. 2 CEMUC, Department of Mechanical Engineering, University of Coimbra, Portugal. 3 Department of Aerospace Sciences, University of Beira Interior, Portugal.

ABSTRACT The main objective of this paper is to study the crack shape evolution in double-U and central hole specimens with corner cracks, both representative of gas turbine discs in terms of critical zones with stress concentrations. An automatic crack growth technique is employed, consisting of a three iterative steps: 3D finite element analysis, K calculation and crack propagation. Complementary experimental work was developed in RR1000, a nickel base superalloy applied in

*

Corresponding author. Tel: +351 239 790 200; Fax: +351 239 790 201. E-mail: [email protected]

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turbine disks, in order to obtain crack shapes and fatigue crack growth at elevated temperature. The effect of residual stresses on crack shape is investigated.

Keywords: Corner cracks at holes, stress intensity factor, crack shape, propagation stages, residual stresses.

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1. INTRODUCTION Aeronautic industry is highly demanding in terms of safety and economy. Turbine disks made of nickel base superalloys are critical components subjected to severe fatigue loading which cannot fail in service. Fracture and fatigue properties of materials are normally obtained from standard test pieces with through thickness cracks, such as the CT or the MT specimens. However, the application of results from CT specimens to corner cracks gives conservative life predictions, i.e., the fatigue life is underestimated (Tong et al, 1999; Brown et al, 1982). Therefore, considering that corner and surface cracks are quite frequent, alternative specimen geometries have been developed, as illustrated in figure 1. The fatigue propagation in some recent nickel base superalloys, such as Udimet 720 or RR1000, has been studied using specimens representative of gas turbine discs in terms of notch geometry and bulk stress, such as the corner crack specimen (Antunes et al, 2001) presented in figure 1c, the double-U specimen showed in figure 1d (Evans et al, 2005; Silva et al, 2010), or the washer specimen (Claudio, 2005). The double-U specimen was developed to reproduce the geometry of the discs at the connection with the blades, since this is a high stress concentration region prone to fatigue failures. Stress intensity factor solutions have been proposed by different authors for corner crack geometries (Raju et al, 1979; Pickard et al, 1986), however, quarter-circular or quarter-elliptical ideal shapes are usually assumed. In practice, the shapes of corner cracks are different from these ideal shapes, namely under conditions of high temperature fatigue. In this case, depending on test parameters, an important tunnelling effect can be observed, which is mainly attributed to the change of crack propagation mechanism with the stress state (Antunes et al, 2001). The crack shape has a great influence on the distribution of the stress intensity factor along the crack front (Branco et al, 2008a; Branco et al, 2008b). Additional surface effects influencing crack shapes are residual stresses and crack closure. Shot peening is one of the many techniques available to mechanically improve the surface properties of

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Effect of Residual Stresses on Crack Shape of Corner Cracks…

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components. The process creates strain hardening and a layer of compressive residual stress at the surface, by plastic deformation. This compressive layer offsets the applied stress, resulting in a benefit in terms of fatigue, corrosionfatigue and fretting fatigue (Byrne et al, 2002). However, little information is published about the possible effect of shot peening in long crack propagation. The fatigue life, the crack shape evolution and the K solutions for the specimens mentioned in figure 1 can be studied using an automatic numerical procedure (Lin et al, 1999). A 3D finite element model is developed to calculate the displacement field, which is used to obtain the stress intensity factors along the crack front. Finally, by applying an adequate crack growth model, taking into consideration experimental da/dN-K curves, it is possible to define the new crack front position. Repetition of this procedure up to the final fracture enables the characterization of crack shape evolution and fatigue life. Surface effects, such as crack closure, residual stresses or change of propagation mechanism can be taken into consideration in this propagation model. A wide range of planar cracks, fastener holes, notched and unnotched round bars, under tension, bending and combined load have been simulated (Lin et al, 1997; Lin et al, 1998; Lin et al, 1998a). Other geometries have been studied, namely MT specimens (Branco et al, 2008a), CT specimens (Branco et al, 2008b), edge flaws in a round bar (Couroneau et al, 1998), short deep and long shallow semi-elliptical surface cracks (Gilchrist el al, 1991) and composite-repaired aluminum plates (Lee et al, 2004). The automatic crack growth technique has been also applied to more complex situations involving realistic components and mixed mode loading (Richard et al, 2007; Sander et al; 2005; Sander et al 2006). The aim of the present article is to study the crack shape evolution in double U and central hole specimens using numerical and experimental approaches. The effect of residual stresses on crack shape is investigated. The experimental work was developed in RR1000, a powder metallurgy nickel base superalloy applied in turbine disks, to obtain crack shapes and fatigue crack growth at elevated temperature.

2. NUMERICAL AND EXPERIMENTAL PROCEDURES 2.1. Numerical Figures 1a and 1d exhibit the geometries of the central hole and the double-U specimens, respectively. Considering the symmetries of geometry

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and loading, only one quarter of the specimens was studied assuming adequate boundary conditions (figure 2). The restrictions at the head of the specimen avoid rotation and bending, and intend to simulate the boundary conditions imposed by the rigid grips of the testing machine. The corner crack is plane, normal to the axis of the specimen and exists in its middle-section, therefore mode-I loading occurs along the whole crack front. The material was assumed to be continuous, homogeneous, isotropic and with linear elastic behaviour.

Figure 1 a) Central hole specimen; b) single edge notch tension specimen; c) corner crack specimen; d) double-U specimen (dimensions in millimeters).

Figure 2. Physical model (R = 5 mm; H/2 = 90 mm, W/2 = 22.5 mm, t = 5.025 mm).

The non-commercial finite element software ModuleF was used to develop the mesh (figure 3a). This consisted of a spider web pattern with three concentric rings centred on the crack tip disposed along eighteen layers (figure 3b); a large regular mesh was used away from the crack front to reduce the number of elements and consequently the computational effort. Isoparametric pentahedric singular elements (with mid-side nodes positioned at quarter point positions) were considered around the crack front. 3D isoparametric elements

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were considered elsewhere: 20-node hexahedral elements and 15-node pentahedric elements. A full Gauss integration was used for these elements, i.e., 333 integration points for the hexahedric elements and 21 integration points for the pentahedric elements. The model contains 1142 nodes and 948 elements (294 pentahedric elements and 654 hexahedral elements). The mesh was refined near the free surfaces to account for surface effects. The nodes along the crack front were positioned on a cubic spline, which provides a good simulation of the real shape (better than a polygonal line).

Figure 3. a) Assembled model; b) spider web mesh; c) layers of spider web mesh.

The automatic crack growth technique employed in this article is schematically presented in figure 4. The automatic procedure generates automatically the three-dimensional finite element mesh since several variables such as geometry, boundary conditions, loading, initial crack shape, elastic properties and fatigue crack growth rate, are previously defined. Then, the displacement field is obtained (figure 4a) which enables the calculation of the stress intensity factor along the crack front (figure 4b). After that, the crack increments are defined using experimental da/dN-K results (figure 4c) and

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consequently the new crack front is obtained (figure 4d) connecting all nodes. Finally, cubic spline functions are used to redefine the positions of mid-side nodes (figure 4e), which gives more realistic crack shapes and contributes to improve the accuracy of the procedure. This crack front is considered as the initial crack shape of the next increment. The entire procedure can be repeated until the final fracture or as long as necessary. The stress intensity factor along the crack front was calculated using the two point extrapolation method (Zhu et al, 1995). At a generic point P, located on the crack surface, in mode-I, it can be written as follows,

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KI 

 8r

 E '  vP

(1)

where vp is the crack displacement, r is the radial distance from the crack tip, E’ is the modified Young modulus with E’ = E/(1-2) in plane strain state and E’ = E in plane stress state and  is the Poisson‟s ratio. Using the K values calculated earlier, the local increments are then calculated using the Paris law for a finite number of load cycles. The propagation at each crack front node (under remote mode-I loading) occurs in a normal direction of the tangent to the crack front at that position (see figure 4c). The crack increment at an arbitrary node along the crack front can be calculated by the following expression,

 K ( j )  ( j ) ai( j )   i( j )  amax  K max 

(2)

( j) where amax is the maximum crack growth increment for the jth iteration.

Once K varies with crack growth, Euler algorithm can be used to calculate the number of load cycles according to the equation,

N ( j 1)  N ( j )  N ( j )  N ( j 1)  N ( j ) 

( j) amax



( j) C K max



m

(3)

where C and m are the Paris law constants.

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Effect of Residual Stresses on Crack Shape of Corner Cracks…

Figure 4. Automatic crack growth: a) definition of initial crack front; b) K calculation along the crack front; c) nodal advances; d) new crack front definition; e) final adjustments.

2.2. Experimental Procedure

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Experimental work on double-U specimens was carried out to validate the numerical predictions. This research was carried out using the RR1000 nickel base superalloy, developed by Rolls-Royce for specific usage in turbine discs of aeroengines. The specimens were tested at 650ºC in a servo hydraulic testing machine with a 100 kN load capacity. Testing temperature was obtained via an electrical furnace. The superalloy‟s chemical composition and mechanical properties at room temperature and at 650ºC are presented in Tables 1 and 2, respectively (Claudio, 2005). Table 1. Chemical composition of the RR1000 nickel base superalloy (mass percentage) Ni

Co

Cr

Mo

Ta

Ti

Al

B

C

Zr

Hf

O2

52.4

18.5

15.0

5.0

2.0

3.6

3.0

0.015

0.027

0.06

0.07

---

Table 2. Mechanical properties of the RR1000 nickel base superalloy Room temp.

650ºC

E [GPa]

214

188.6

0.2% Proof. MPa

1086

1034

UTS MPa

1602

1448

Poisson‟s ratio, 

0.255

Fatigue tests were performed with a 0.1 load ratio taking into consideration two waveforms: sinusoidal with 5Hz frequency and trapezoidal with 30s dwell time (1-30-1-1s). The potential drop technique was used for crack propagation monitoring purposes using a DCPD pulsed system coupled

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with the controller of the servo hydraulic machine. During the tests the loading conditions (either stress ratio or loading frequency) were changed to record visible marks on the fracture surface which enables the identification of crack shapes. Figure 5a exhibits the fracture surface of a particular specimen. At least three corner cracks are distinctly observed. The two smaller visible crack shapes were measured for validation purposes. Figure 5b compares these crack shapes with the ones predicted by the application of the automatic crack growth technique described earlier.

Figure 5. a) Fracture surface of LF001 specimen tested with a 5Hz sinusoidal load and R=0.1; b) Numerical versus experimental results.

The crack lengths of each mesh layer for the smallest and the following visible crack shapes were measured. The former was used to define the initial crack shape in the numerical procedure whilst the latter was compared with the corresponding numerical prediction (dashed line). Both crack shapes (dashed and full lines) are almost similar which indicates a good validation. Differences between experimental loading and the loads considered in the numerical analysis can explain this mismatch. Eight test specimens were shot peened on all the surfaces with peening parameter 110H 6-8A 100%. This parameter was based on previous optimization studies made by Rolls-Royce plc. They also provided the residual stress profiles in the conditions “as-machined” and “shot peened” which are plotted in figure 6 (Cláudio, 2005). The residuals were measured at room temperature, before any thermal exposure, by X-Ray diffraction combined with electrolytic polishing to assess in depth values. According to Cláudio

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(2005), and as can be seen in figure 6, the residual stress level introduced by the shot peening is very high. At a 20 m depth, the stress reaches 1520 MPa which is close to the tensile limit of the material. However, the effect of shot peening almost disappears at a depth greater than 100 m.

Figure 6. Residual stress measurements made in RR1000 washer specimens at room temperature and estimated profile after 100h at 650ºC (Cláudio, 2005).

Figure 7a presents predictions of the crack opening level (open/max) obtained by fixing the stress range and increasing the minimum and maximum stresses, as it is schematised in figure 7b. Each cycle in figure 7b corresponds to one numerical prediction () in figure 7a. The first node behind the crack tip was used to quantify the crack opening level. The ratio between maximum stress and yield stress (max/ys) is also presented, showing that it increases up to about 0.78 for R=0.64. The monotonic plastic zone is expected to increase with Kmax, while the cyclic plastic zone is expected to be constant. From the figure it is possible to conclude that the increase of the mean stress produces a clear increase of U. For R>0.64 no closure is observed, i.e., U=1. The predictions obtained according to these models are plotted in figure 7a. Despite the discrepancies in loading conditions, materials and closure definition, a good agreement can be found between the predictions and the numerical simulation results up to R=0.

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Figure 7. Effect of R for a fixed K (a/W=0.24 mm, =36 MPa, L1=16 m, a=15×16=240 m).

Schijve (1981), based on the work of Newman (1976), proposed the equation,

 op  max

 0.45  0.22R  0.21R 2  0.12R 3

(4)

to describe the variation of the opening level (open/max) with R. In the same year, Koning (1981) proposed the following model for the 7075-T6 aluminium alloy: Superalloys : Production, Properties and Applications, Nova Science Publishers, Incorporated, 2011. ProQuest Ebook Central,

Effect of Residual Stresses on Crack Shape of Corner Cracks…

 op  max

  3  2 3 4 1  0.25(1  R) max (0.45  0.2 R  0.15 R  0.9 R  0.4 R )  ys     1  0.25(1  R ) 3  max (0.45  0.2 R )   ys  

11

R 0 R0

(5)

3. PRESENTATION AND ANALYSIS OF RESULTS

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3.1. Stable Crack Shapes Figures 8 and 9 illustrate the crack shape development from four initial crack shapes (one quarter-circular, two quarter-elliptical and one irregular) for both double-U and central notch specimens, respectively. Many profiles were suppressed in order to provide a better illustration. The effect of the initial crack shape is clearly shown during the early propagation stage. Different crack profiles were obtained in each case during this period. However, with further propagation the crack tends towards similar profiles. By comparing both figures, for double-U and central hole specimens, there are no significant differences. However, the crack front is usually not balanced which can be explained by a non-uniform stress distribution along the crack in both free surfaces (hole and front directions), as can be seen in figure 10. Figure 10 shows the evolution of the stress concentration factors at the hole edge direction (Kt,z) and at the front surface direction (Kt,x) for both double U and central hole specimens. The stress concentration factors were defined as the ratio between local and remote tension stresses. The normal stress at the hole edge caused by remote uniform stress tension is about three times the remote stress whilst the normal stress decreases continuously along the plate surface direction towards the remote stress. Comparing both specimen geometries, it can be concluded that the central hole specimen has greater stress concentration factors at the hole edge, whereas in the normal direction to the hole edge (frontal surface) the values are quite similar and homogeneous.

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R. Branco, F.V. Antunes, J.A. Martins Ferreira et al.

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Figure 8. Shape development for double-U specimen considering a: a) quarter-circular initial crack shape (a0/c0 = 1); b) irregular initial crack shape (a0/c0 = 0.90); c) quarterelliptical initial crack shape (a0/c0 = 1.53); d) quarter-elliptical initial crack shape (a0/c0 = 0.66).

Figure 9. Shape development for the central hole specimen considering a: a) quartercircular initial crack shape (a0/c0 = 1); b) irregular initial crack shape (a0/c0 = 0.90); c) quarter-elliptical initial crack shape (a0/c0 = 1.53); d) quarter-elliptical initial crack shape (a0/c0 = 0.66).

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Effect of Residual Stresses on Crack Shape of Corner Cracks… 3,5 3,0 Kt,z 2,5

Kt

2,0

x

z

1,5 1,0 central hole

0,5

Kt,x

double U 0,0 Copyright © 2011. Nova Science Publishers, Incorporated. All rights reserved.

0,00

0,01

0,01

0,02

0,02

x or z coordinate

Figure 10. Stress concentration factors (Kt) for double U and central hole specimens at the hole direction and at the front surface direction.

In truth, observing the crack shape development, it seems that a quasi quarter-elliptical crack shape is reached and maintained after the early propagation stage. This is a consistent behaviour which is independent on the initial crack shape. In order to investigate how much close to the quarter-elliptical crack shape are those profiles, a quantitative study of crack shape variation was carried out. Two effective characterising parameters were defined. The residual difference (hi) can be written as follows,

hi 

di ri  d i

(6)

being the variables defined in figure 11 (ri and  are the polar coordinates of the ith node along the predicted crack front). The standard residual deviation

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R. Branco, F.V. Antunes, J.A. Martins Ferreira et al.

(st) which gives a general appreciation of whole crack front, is defined by the expression,

st 

1 n 2  di n i 1

(7)

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where di is the difference between the radius of ith and i‟th nodes, and n is the number of corner nodes of the crack front. The average crack length (r‟) for a given crack front was calculated by the mean of the nodal crack length (ri) of each mesh layer (see figure 3c).

Figure 11. Definition of the dependent parameters.

Figure 12a shows the distribution of the residual difference around the crack front during the crack growth for several crack fronts in both double U and central hole specimens with r/t = 1 and initial crack shapes with a0/c0 = 1.0. The crack fronts analysed are exhibited at the bottom of this figure. As can be seen, the residual difference is always zero at both end points because the perfect quarter elliptical shape is defined through these two points. Besides, this variable is always negative in the remaining points, which means that the predicted crack shape lays inside the corresponding quarter-ellipse. On the other hand, the residual difference increases continuously during the crack propagation, reaching absolute values higher than 8%. It is also clearly observed that the influence of the specimen geometry is basically small once the residual differences for both double U and central hole are very close and present the same type of evolution. Additionally, this figure shows a comparison between the results obtained in this specific study and the ones found by Lin et al (Lin et al, 1998) in their studies concerned with corner cracks at fastener holes with similar ratios of a0/c0 and r/t. As observed, the

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evolutions of di with the angle  are quite similar, even though in this situation smaller residual differences are obtained. Figure 12b presents the evolution of the standard residual deviation with the crack growth for both double-U and central hole specimens with r0‟/t=0.2. Two initial crack shapes with a0/c0=1.53 (dashed lines) and a0/c0 = 1.0 (full lines) are analysed. Basically, it can be seen that the st parameter increases continuously as the crack propagates and is about 7% as the crack nearly approaches the back surface of the specimen. Lin et al (1998) studied exhaustively the evolution of st for different values of a/t and a0/c0 and have concluded that the maximum value of st is about 6%. An important effect of the initial crack shape is clearly observed in figure 12b. Nevertheless, the crack propagation tends asymptotically to preferred crack paths independently on the value of a0/c0. There is a more intense convergence in the early propagation regime, specially exhibited in the case with a0/c0=1.53, which may occur due to a higher driving force of crack in this period which forces more rapidly the crack towards the equilibrium. Figure 12b also shows that the influence of specimen geometry on the standard residual deviation is not relevant since the evolutions of st are always quite similar for both specimens. The previous conclusions demonstrate that the numerical predictions are close to the quarter-elliptical crack shape. Therefore, accurate analyses of fatigue crack growth based on this presupposition can be successfully carried out. 2 Double U Lin et al

0

Central hole

Sé

Sé

a/t=0.33

-2

hi [%]

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Effect of Residual Stresses on Crack Shape of Corner Cracks…

Sé -4 -6

Sé

a/t=0.51

r/t=1

Sé

a0 /c0 =1

Sé

Sé

-8

3,0

a/t=0.77 -10 0

15

30

45

60

75

90

 [º]

Figure 12 (continued).

a

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Sé

16

R. Branco, F.V. Antunes, J.A. Martins Ferreira et al. 8 Double U

Central hole

st [%]

6

a/c = 0.66 (central hole) a/c = 1.53 (double-U) a/c = 0.66 (double-U) a/c = 1.53 (central hole)

4 a0/c0=1.53 2

a0/c0=0.66 r0' /t=0.2

0 00

0.2 1

20.4

b

30.6

40.8

r'/t

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Figure 12. Evolution of: a) di with  during the crack growth; b) st with the r‟ during the crack growth for both double-U and central hole specimens.

3.2. Crack Propagation Stages The crack shape change during the propagation can be characterised by studying the variation of a/c against the ratio r0’/t. Figure 13a exhibits the variation of both variables employing the present simulation technique for several initial quarter-elliptical crack shapes with different values of a0/c0 taking into consideration the double U specimen. The dashed line presents the evolution of an initial quarter-elliptical crack shape with a0/c0=1 for the central hole geometry. Two propagation stages are perfectly distinguished in figure 13a. During the early propagation (I) stage the crack path depends significantly on the initial crack shape and for each case a different trajectory is followed by the crack while in the remaining propagation (II) the crack follows a preferential path independent on the initial crack shape. These concepts have already been quoted in literature by several authors. Lazarus (1999) analysed an embedded crack in an infinite body under uniform and cyclic tension and observed that, after certain propagation time the crack reaches a circular shape independently on the initial crack shape analysed. Lin et al (Lin et al, 1997) studied three semi-elliptical surface cracks with different initial aspect ratios, subjected to

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17

tension and bending and concluded that all these cracks propagate towards a preferred aspect ratio. Similar results were found in their studies with cracks emanating from fastener holes and cracks in notched an unnotched fatigued round bars (Lin et al, 1998a). It is also observed that the cracks which are initially more distant from the second propagation stage require prolonged growth to reach it, but for each a0/c0 only a single path can be followed by the crack. Besides, comparing the evolutions of a0/c0=1 for both geometry it can be found again minimal differences between those curves which is in accordance with the general conclusions obtained before. A different analysis is carried out in figure 13b. Three distinct initial crack shapes (plotted at the bottom of the figure) with the same average crack length (r0‟) and unitary a0/c0 ratios are studied for a central hole specimen. As can be seen, in this case the curves are similar during the whole crack propagation. Therefore, crack path depends only on the initial state which is a good indication of the robustness of a/c ratio as a dependent parameter to quantify variations of crack shape. A mathematical model suitable to predict both stages has already been enunciated by Branco et al (2007) in their studies involving double U and central hole geometries. According to the authors the early propagation model has an exponential behaviour which only depends on its initial remoteness while the subsequent propagation is quite stable and can be fitted by a fourth order polynomial function. 2,6 2.6 a0/c0

2.40 1.00 0.65

2.0 2,0

1.53

Série8

1.36

Série3

1.00 (central hole) 0.42

I a/c

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Effect of Residual Stresses on Crack Shape of Corner Cracks…

Série1 Série9

II

Série10

1.4 1,4

Série2 Série4

0,8 0.8

t

c

0.2 0,2

0.15 0,8

Figure 13. (continued).

Série5 a

0.28 1,4

0.41 2,1

a

0.54 2,7

0.67 3,4

0.80 4,0

r'/t

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R. Branco, F.V. Antunes, J.A. Martins Ferreira et al. 0,0012

1.4 1,4

0,001

0,0008

0,0006

1.3 1,3

0,0004

0,0002

1,2 1.2

r

a/c

0 0

1,1 1.1

1

r/t=1

t

a0 /c0 =1 1.0 1,0 1 0.9 0,9

0.15 0,001 0.28 0.41 0,001 0,002 0.54 0,003 0.67 0,003 0.80 0,004

b

r'/t

Figure 13. Evolution of a/c with r’/t for a) several initial crack shapes; b) for three different initial crack shapes with the same average crack length.

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3.3. Effect of Residual Stress As earlier mentioned (see figures 6 and 7), the residual stress level introduced by shot peening causes mean stress changes. This effect is also responsible by a variation of the crack closure level. Thus, in order to study the effect of residual stresses on crack shape and fatigue crack growth, an effective stress intensity range was considered, given as follows:

K ( j )  U i . K ( j ) i ,eff

i

(8)

where Ui is the fraction of the load cycle for which the crack remains fully open. In this study, a value of U=0.48 was defined on the surface layers. This value is representative of the residual stress field under simulation. Figure 14 shows the crack shape development for four initial crack shapes (one quarter-circular, two quarter-elliptical and one irregular) for the double-U specimen considering crack closure on the surface layers. Similar results were obtained for the central hole specimen. Many profiles were suppressed in order to provide a better illustration.

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0,0002

0,0004

0,0006

0,

Effect of Residual Stresses on Crack Shape of Corner Cracks…

19

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Figure 14. Shape development for double-U specimen, with U=0.48, considering: a) quarter-circular initial crack shape (a0/c0 = 1); b) irregular initial crack shape (a0/c0 = 0.90); c) quarter-elliptical initial crack shape (a0/c0 = 1.53); d) quarter-elliptical initial crack shape (a0/c0 = 0.66).

In this case, there is no doubt that the crack grows towards a non quarterelliptical crack shape. A superficial delay at the hole edge and at the front surface can be clearly distinguished. This effect is more evident especially near the front surface whilst at the hole edge it is also observed but with a smoother intensity. These conclusions are illustrated in figure 15a which exhibits the distribution of the residual difference around the crack front during the crack growth for several crack fronts in a double U specimen with r/t = 1 and an initial quarter-circular crack shape (a0/c0 = 1.0). As enunciated, near the front surface (angle amplitudes about 0º) the residual differences are always greater than near the hole edge (amplitudes about 90º). Naturally, for =0º and =90º the values of hi are null because both crack lengths are used to define the quarter-elliptical crack shape. Figure 15b presents the evolution of the standard residual deviation with the crack growth for a double-U specimen with r0‟/t=0.1 and for an initial crack shapes with a0/c0=1.0. Two situations with and without crack closure are compared. In the latter, as presented in figure 12b, this parameter increases rapidly at the early propagation stage and then remains approximately constant. The maximum differences are about 7%, as observed before in figure 12b. With crack closure, the standard deviation is always greater. At the

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R. Branco, F.V. Antunes, J.A. Martins Ferreira et al.

beginning of the propagation st increases suddenly and then drops steadily to values about 9%. It is interesting to observe that both initial periods (characterized by high st gradients) have, apparently, the same range. The effect of the initial crack shape is shown again in this figure.

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a

b

Figure 15. Evolution of: a) di with  during the crack growth with U=0.48; b) st with the r‟ with U=0.48 during the crack growth for double-U specimen.

CONCLUSIONS

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Effect of Residual Stresses on Crack Shape of Corner Cracks…

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In this paper, the effects of residual stresses on the crack shape evolution in double U and central hole specimens were investigated, employing both numerical and experimental approaches. The main conclusions are: •

•

•

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•

•

•

an automatic crack growth technique consisting of a three iterative steps (3D FEM analysis, K calculation and crack propagation) was successfully implemented and used to study the crack shape development; experimental work was carried out using the RR1000 nickel base superalloy. The specimens were shot peened on all the surfaces with peening parameter 110H6-8A; fatigue tests were carried out with constant load ratio (R=0.1). During these tests the loading conditions were changed to record visible marks on the fracture surface which were used to validation purposes. Through the comparison of the numerical predictions and these visible crack shapes it was found that both data are quite close; the crack shape development is strongly dependent on the initial crack shape but after a short initial propagation stage tends towards a preferred propagation path independently on the initial crack shape. The crack grows under a not balanced fashion which can be explained by the non-uniform stress concentration factors at the hole edge and the front surface. At the hole edge the normal stress is approximately three times the remote tension whereas at the front surface the normal tension tends towards the remote tension; The geometrical proximity of the crack profiles to the quarterelliptical crack shape was investigated. It was found that this assumption is adequate and can be efficiently used to estimate the crack shape propagation; The effect of residual stresses was studied by the introduction of crack closure on the numerical model. The residual stresses affect the stress ratio and consequently it directly affects the crack closure level as well. For U=0.48 at the free surfaces, the results show that the crack shape is more distant from the quarter-elliptical crack shape and an approach based on this presupposition is inadequate. The tunneling effect rises significantly as well as the delay at the free surfaces, especially at the front surface.

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R. Branco, F.V. Antunes, J.A. Martins Ferreira et al.

The authors gratefully acknowledge the financial support provided by the FCT (Fundação para a Ciência e a Tecnologia) through the research project PTDC/EME-PME/114892/2009.

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REFERENCES Antunes FV, J.A.M. Ferreira, C.M. Branco e J. Byrne, “Influence of stress state on high temperature fatigue crack growth in Inconel 718”, Fatigue and Fracture of Engineering Materials and Structures 24, pp.127-135, 2001. Branco R, Antunes FV (2008). Finite element modelling and analysis of crack shape evolution in mode-I fatigue middle-cracked tension specimens. Engineering Fracture Mechanics, 75, pp. 3020–337. Branco R, Antunes FV, Martins RF (2008). Modelling fatigue crack propagation in CT specimen. Fatigue and Fracture of Engineering Materials and Structures, 31, pp. 452–465. Branco R, Antunes FJV, Martins Ferreira JM, Silva JM (2007). Automatic fatigue crack growth in nickel base supperalloy at elevated temperature. The 6th Engineering Integrity Society International Conference on Durability and Fatigue, 26-28 March, Queens‟ College, Cambridge, UK. Brown CW, Hicks HA (1982). Fatigue Growth of Surface Cracks in NickelBase Superalloys. Int. J. Fatigue, 4, 73-81. Byrne J, Burgess A (2002). Surface improvement for structural integrity and life extension. 8as Jornadas de Fractura, UTAD, Vila Real de Trás os Montes, Portugal, Ed. by SPM (Portuguese Society of Materials) Cláudio RA (2005). Fatigue Behaviour and Structural Integrity of Scratch Damaged Shot Peened Surfaces at Elevated Temperature. PhD Thesis, University of Portsmouth, UK. Couroneau N, Royer J (1998). Simplified model for the fatigue crack growth analysis of surface cracks in round bars under mode I. International Journal of Fatigue, 20, 711-718. Evans WJ, Jones JP, Williams S (2005). The interactions between fatigue, creep and environment damage in Ti 6246 and Udimet 720Li. International Journal of Fatigue, 27, 1473-1484. Gilchrist MD, Smith RA (1991). Finite element modelling of fatigue crack shapes. Fatigue and Fracture of Engineering Materials and Structures, 6, 617-626.

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Koning AU (1981). A simple crack closure model for prediction of fatigue crack growth rates under variable amplitude loading. ASTM STP 743, 6385. Lazarus V (1999). Fatigue propagation path of 3D plane cracks under mode I loading. Comptes Rendus de l‟Académie des Sciences Paris, t.327, Série IIb, 1319-1324. Lee WY, Lee JJ (2004). Successive 3D analysis technique for characterization of fatigue crack growth behaviour in composite-repaired aluminum plate. Composite Structures, 66, 513-520. Lin XB, Smith RA (1997). An improved numerical technique for simulating the growth of planar fatigue cracks. Fatigue and Fracture of Engineering Materials and Structures, 20, 1363-1373. Lin XB, Smith RA (1998). Fatigue shape analysis for corner cracks at fastener holes. Engineering Fracture Mechanics, 59, 73-87. Lin XB, Smith RA (1998a). Fatigue growth simulation for cracks in notched and unnotched round bars. International Journal of Mechanical Sciences, 40, 405-419. Lin XB, Smith RA (1999). Finite element modelling of fatigue crack growth of surface cracked plates. Part I: The numerical technique. Engineering Fracture Mechanics, 63, 503-522. Newman Jr JC (1976). A finite element analysis of the fatigue crack closure. Mechanisms of crack growth, ASTM STP 590, 281-301. Pickard AC (1986). The Application of 3-Dimensional Finite Element Methods to Fracture Mechanics and Fatigue Life Prediction, EMAS. Raju IS, Newman Jr JC (1979). Stress intensity factors for two symmetric corner cracks. Fracture Mechanics, ASTM STP-677, Ed. by Smith CW, 411-430. Richard HA et al (2007). Development of fatigue crack growth in real structures. Engineering Fracture Mechanics. doi:10.1016/j.engfracmech. 2007.01.017 Sander M, Richard HA (2005). Finite element analysis of fatigue crack growth with interspersed mode I and mixed mode overloads. International Journal of Fatigue, 27, 905–913. Sander M, Richard HA (2006). Experimental and numerical investigations on the influence of the loading direction on the fatigue crack growth. International Journal of Fatigue, 28, 583–591. Schijve J (1981). Some formulas for the crack opening stress level. Eng. Fracture Mechanics, 14, 461-465.

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Silva JM, Cláudio RA, Moura Branco C, Martins Ferreira JA (2010). Creepfatigue behaviour of a new generation Ni-base superalloy for aeroengine usage. Procedia Engineering, 2,1865-1875; Tong J, Byrne J, Hall R, Aliabadi MH (1999). A Comparison of Corner Notched and Compact Tension Specimens for High Temperature Fatigue Testing, Proc. Conference Eng. Against Fatigue, 17-21 March, University of Sheffield, UK Zhu WX, Smith DJ (1995). On the use of displacements extrapolation to obtain crack tip singular stresses and stress intensity factors. Engineering Fracture Mechanics, 51, 391-400.

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In: Superalloys: Production, Properties… ISBN: 978-1-61209-536-3 Editor: Jeremy E. Watson, pp. 25-43 © 2011 Nova Science Publishers, Inc.

Chapter 2

WROUGHT SUPERALLOYS – SCIENCE, TECHNOLOGY AND APPLICATIONS M. Nageswara Rao*

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Senior Professor Emeritus, School of Mechanical & Building Sciences VIT University, Vellore, Tamil Nadu, India – 632014.

ABSTRACT Superalloys are heat-resistant alloys based on iron, nickel-iron, nickel or cobalt, exhibiting high creep and rupture strength at elevated temperatures. A number of both wrought (produced through ingot metallurgy route) and cast grades of superalloys have been developed over the years. There is also a group of superalloys produced through powder metallurgy route. This article focuses on the metallurgy, high temperature mechanical properties and applications of conventional wrought superalloys based on iron, nickel and cobalt. A number of wrought superalloy grades have become commercially available over the last seven decades; the emphasis was on the development of alloys with higher and higher creep resistance. A noteworthy feature of wrought nickel base superalloys available today is their heat-resistance upto 80% of their incipient melting temperature, a fraction that is higher than for any other class of engineering alloys. There have been a number of processing discoveries which helped in development of wrought superalloys with higher and higher performance levels. The strengthening mechanisms that come into play in different groups of wrought *

Email: [email protected]

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M. Nageswara Rao

superalloys have been well understood. Certain trace elements play a major role in superalloys and there have been a number of studies dealing with underlying mechanisms. Considerable progress has been achieved with reference to understanding the metallurgical stability of wrought superalloys under high temperature loading conditions and its relation to chemical composition. While non-aerospace applications have been there, prominently in the power generation sector, aerospace accounts for the major part of the consumption of the wrought superalloys. Developments in wrought nickel-iron and nickel base superalloys over the decades have importantly contributed to the evolution of modern high performance jet engines.

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1. INTRODUCTION Superalloys are heat-resisting alloys based on nickel, nickel-iron, iron or cobalt that exhibit an excellent capability to withstand sustained loading at elevated temperatures. Superalloys are primarily used in aircraft gas turbine engines, power generation sector, and chemical process industries, and for other specialized applications requiring heat resistance. The modern highperformance aircraft engines would not have been there but for the major advances made in superalloy development over the past 50 years. Excellent monographs / handbooks / technical guides are available [1-6] on the subject of superalloys, covering different aspects – metallurgy, processing, properties and applications. The reader is referred to them for more detailed information. Based on the matrix element(s), superalloys can be broadly be classified into four families: I. II. III. IV.

Iron-base (Fe-base) Nickel-iron base (Ni-Fe-base) Nickel-base (Ni-base) Cobalt-base (Co-base)

Ni-base superalloys are the most complex, the most widely used for the gas turbine engine parts experiencing the highest temperatures. They have the highest creep strength among the above families. A remarkable feature of Nibase superalloys is their use in load-bearing applications at temperatures as high as 80% of their incipient melting temperatures, a percentage that is higher than for any other class of engineering alloys. However, they are costlier than

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Wrought Superalloys – Science, Technology and Applications

27

Ni-Fe-base and Fe-base superalloys. Ni-base and Ni-Fe-base superalloys put together constitute upto 50% of the weight of advanced aircraft engines. Based on the method of manufacture, the superalloys can be broadly classified into three categories (i) wrought superalloys, processed through ingot metallurgy route and adopting hot working process(es) to convert the material into the required size and shape (ii) cast superalloys, where the component is cast directly into shape, without any hot working coming into picture and (iii) powder metallurgy based superalloys, including oxide dispersion strengthened grades. In this chapter, the treatment is confined to wrought superalloys. Many coatings have been developed to extend the usage of superalloys to higher temperature applications and corrosive environments; this subject is however outside the scope of this chapter.

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2. IRON-BASE SUPERALLOYS Fe-base superalloys evolved from austenitic stainless steels and are based on the principle of combining a closed-packed face centered cubic (FCC) matrix with (in most cases) both solid-solution hardening and precipitateforming elements [6]. The austenitic matrix is based on nickel and iron, with at least 25% Ni needed to stabilize the FCC phase. Other alloying elements, such as chromium, partition primarily to the austenite for solid-solution hardening. The strengthening precipitates are primarily ordered intermetallics, such as γ′ Ni3Al, η Ni3Ti, and γ″ Ni3Nb [6].

3. NICKEL-BASE AND NICKEL-IRON-BASE SUPERALLOYS Ni-base superalloys in commercial service range from solid solution strengthened single-phase alloys to precipitation-hardened alloys. Inconel 600 and Hastelloy C are examples of solid solution strengthened Ni-base superalloys. Inconel 706 is an example of precipitation hardened Ni-base superalloys. Inconel 718 is a Ni-Fe-base precipitation hardened superalloy and accounts for major part of the total superalloy tonnage produced / consumed. The importance of this alloy and its derivatives is seen by several international conferences devoted to them [7-12].

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M. Nageswara Rao

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As illustrated in Fig 1, which schematically compares the rupture strengths of different families of wrought superalloys, the precipitation hardening Ni-base and Fe-Ni-base alloys have considerably higher rupture strength values, when compared to solid solution strengthened Fe- and Ni-base alloys. The latter group of alloys is preferred when service conditions allow their use, because of their ease of fabrication, especially weldability. For the most demanding of elevated temperature applications, precipitation strengthened alloys are preferred.

Figure 1. Stress-rupture strengths of different families of wrought superallows (schematic).

Table 1 gives the details of chemical composition for representative wrought Ni-base superalloy grades. It includes one grade of solid solution strengthened Russian superalloy (Alloy 868). The Russian superalloy grades have evolved differently; a notable difference is the presence of substantial levels of tungsten as an alloying element. Studies on two Russian based solid solution strengthened superalloys have been published by Nageswara Rao and Mayadeo [13]. The treatment in the present chapter will be with emphasis to Western grades. Some Ni-base superalloy compositions show excellent resistance to corrosive media such as aqueous solutions and acids at low temperatures. The Ni-Cr-Mo compositions such as Alloys 625, C276, 59 and 686 have been developed for such applications. Information on chemical composition of these grades is included in Table 1. A comprehensive coverage of these grades and other nickel alloys for corrosion resistance applications can be found at [14].

Superalloys : Production, Properties and Applications, Nova Science Publishers, Incorporated, 2011. ProQuest Ebook Central,

Copyright © 2011. Nova Science Publishers, Incorporated. All rights reserved.

Table 1. Chemical composition of representative wrought nickel base superalloy grades (wt%). Alloy Solid solution alloys Inconel 600 Hastelloy C276

Ni

Cr

76 58

15.5 16

Inconel 625 Alloy 868 (Russian) Grade 59 Grade 686

61 Bal. 59 57

21.5 25 23 21

Precipitation-hardening alloys Inconel X750 73 Udimet 500 53.6 Udimet 700 53.4 Udimet 720 Bal. Waspaloy 58.3 Astraloy 55.1 Rene 41 55.3 Nimonic 80A 74.7 Nimonic 90 57.4 Nimonic 105 53.3 Nimonic 115 57.3

15 18 15 18 19.5 15.0 19.0 19.5 19.5 14.5 15.0

Co

Mo

Al

Ti

Nb

16 16

18.5 18.5 14.8 13.5 17.0 11.0 1.1 18.0 20.0 15.0

4.0 5.2 3.0 4.3 5.2 10.0 5.0 3.5

B

Zr

0.08