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Fracture initiation mechanisms of thin hard coatings during the impact test
Surface & Coatings Technology 185 (2004) 150 – 159 www.elsevier.com/locate/surfcoat
Fracture initiation mechanisms of thin hard coatings during the impact test K.-D. Bouzakis *, A. Siganos Laboratory for Machine Tools and Manufacturing Engineering, Mechanical Engineering Department, Aristoteles University of Thessaloniki, GR-54124 Thessaloniki, Greece Received 25 February 2003; accepted in revised form 24 December 2003 Available online 18 March 2004
Abstract The impact test can be successfully applied for the characterization of thin and thick coatings fatigue and creep properties, respectively. In the present paper experimental and analytical investigations concerning thin films fracture initiation during this test are introduced. The damage initiation of thin hard coatings deposited on smooth surfaces is mainly induced by fatigue mechanisms and characterized by the removal of film segments in the contact area between the ball indenter and the specimen in highly superficially loaded contact regions during the impact test. Herein the currently used coating stress determination methodologies, when applied, lead in most cases to sufficiently accurate results. Superficial PVD layers deposited on a basic coating do not affect the occurring stress field practically and thus no differences concerning the fracture initiation in the basic coating are encountered. When the surface roughness increases, localized stress concentrations throughout the contact area occur, resulting in film failures in randomly distributed spots, at predictable high impact loads. In cases of relatively soft PVD films or at elevated substrate roughness, the abrasive wear contribution to the coating damage initiation has to be considered. In such cases, as will be explained, the first film fracture appears in a region, close to the middle of the contact radius. D 2004 Elsevier B.V. All rights reserved. Keywords: Impact test; Thin hard coatings; Fracture initiation
1. Introduction Coatings comprise one of the most up-to-date technologies used to enhance the superficial properties of machine elements and components of tools, etc. [1]. In all these applications the knowledge of the coating material properties is of great importance. The impact test is used successfully for coating fatigue and creep property characterization [2 – 6]. Its working principle is shown in Fig. 1. An oscillating indenter, usually a cemented carbide ball, penetrates successively into the coated specimen at a constant maximum force. During the test the impact force is continuously monitored with the aid of a piezoelectric dynamometer. The impact load and the number of impacts may induce a coating failure. Fig. 2 shows the impact force vs. the number of impacts leading to coating failure. The coating failure of thin hard coatings on low roughness substrates, is initiated mainly by fatigue mechanisms and a continuous endurance load can be * Corresponding author: Tel.: +30-2310-996079, +30-2310-996021; fax: +30-2310-996059. E-mail address: [email protected] (K.-D. Bouzakis). 0257-8972/$ - see front matter D 2004 Elsevier B.V. All rights reserved. doi:10.1016/j.surfcoat.2003.12.028
determined, corresponding to the maximum force, at which no coating damage after 106 impacts is encountered. With the aid of FEM simulation of the impact test at the maximum ball-indenter penetration and during the specimen relaxation between successive impacts, Smith and Woehler diagrams of the examined coating materials have been established, as indicated in Fig. 2 [3,4,7,8]. In high substrate roughness cases, stress concentrations occur, resulting in potential local coating overloads, beyond the film fracture strength. Moreover, in soft thin coating cases or at elevated substrate roughness, abrasive wear may contribute also to the coating damage initiation during the impact test. In the following sections, experimental and analytical investigations explaining the film failure mechanisms previously mentioned, in various film cases, will be introduced.
2. Coating fracture initiation mechanisms due to fatigue During the impact test a severe plastic deformation of the substrate may occur. In order to determine the stress field in the coating – substrate compound, FEM supported algo-
K.-D. Bouzakis, A. Siganos / Surface & Coatings Technology 185 (2004) 150–159
Fig. 1. The impact tester in the investigations applied.
rithms are used. The FEM methodologies in [4,7,8] presented are based on the application of a semi-elliptical pressure distribution on the contact circle, the radius of which is experimentally determined. The coating stress field is only at the maximum contact force and during the relaxation stage examined. However, in cases of elastic – plastic contact, the pressure distribution is flattened and not semi-elliptical [9]. In the frame of the investigations de-
Fig. 2. Impact test simulation to elaborate coatings Smith and Woehler diagrams.
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scribed, a new FEM model has been developed, presented in Fig. 3, considering not a pressure distribution, but an elastic – plastic ball indenter penetration into an elastic – plastic film-substrate compound as well. The indenter contact with the coated specimen was described with the aid of contact elements and the occurring pressure distribution determined [10]. The validity of this impact test simulation was verified through the comparisons between the calculated and the measured contact area magnitudes. Using this model, the fracture initiation during the impact test of a TiAlN film deposited on the typical bearing steel 100Cr6, was investigated. Multi-linear hardening laws associated to the applied materials, i.e. the carbide ball indenter, the coatings and the substrate have been applied. These laws were obtained through nanoindentation tests [11– 13]. Fig. 4 illustrates the superficial stress distribution on the coating at successive stages of the ball indentation, as well as during the relaxation stage. In order to determine the most loaded region inside the contact circle, the stress field development at all these ball penetration stages is considered and the deriving maximum superficial stress distribution is shown in Fig. 5a. This stress distribution does not coincide with the corresponding one when the impact force reaches its maximum value, which is also illustrated in the figure. According to these results, the most loaded region is shifted slightly towards the contact area center, just before the maximum impact load. Hence, the maximum von Mises equivalent stresses, according to the differences in principal stresses, are encountered very close to the imprint vicinity.
Fig. 3. The developed impact test FEM model with deformable ball indenter.
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between the results through the elastic plastic deformable ball indenter and the semi-elliptical pressure distribution FEM models achieved is less than 7%. Due to this fact both FEM models can be reliably applied to determine the contact stress and to elaborate fatigue diagrams. If a high substrate plastic deformation during the impact test occurs, as for example in the case of a 2.5 mm diameter ball indenter application at elevated impact loads, the results deviation increases and the use of the deformable ball model in order to obtain more accurate results is recommended [14]. Experimental results concerning the film failure initiation mechanisms on smooth substrates (Raf0.017 Am) are in Fig. 7 illustrated. The continuous endurance of TiAlN coating was found to be at an impact force of 45 daN. The coating fatigue fracture initiation can be accurately revealed by means of SEM micrographs, whereas in the indicated regions in Fig. 7, the substrate material could be detected by means of EDX microanalyses. It is obvious that at the coating damage initiation small coating segments are removed close to the imprint vicinity. In these contact area regions the superficial equivalent stresses possess a maximum, according to the previously described FEM calculations results. In general, the tribological behaviour can be enhanced through the deposition of top DLC layers, due to their low friction coefficient. In order to investigate the influence of
Fig. 4. Superficial stress distribution at intermediate stages during the ball indentation.
The equivalent stress increase at the crater vicinity is caused by the deformation occurring in the substrate, which forces the coating to ‘bend’. However, the principle stress absolute values are maximized in the center of the contact circle (see Fig. 5b). As already mentioned, the FEM methodologies applied to determine the stress field in the coatings during the impact test [4,7,8] assume a semi-elliptical pressure distribution at the maximum impact force. A comparison of the maximum stress values obtained with the aid of the described deformable ball indenter FEM model and calculated by means of the methods in [4,7,8], has been performed in several coating – substrate cases. Some characteristic results are demonstrated in Fig. 6. For the applied impact forces, which are usual for the impact test, the substrate plastic deformation does not reach extreme values and the deviation
Fig. 5. Superficial maximum and minimum stress distribution during the impact test.
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Fig. 6. Maximum superficial Seqv stresses, by means of various FEM models calculated.
such layers on the impact performance of TiAlN coatings, a top amorphous carbon (aC) layer was deposited, through the application of a carbon target during the PVD process. The effect of the top layers on the stress field developed in the basic coating was determined by means of the FEM model presented in Fig. 3, which was properly modified. As in Fig. 8 illustrated, the FEM simulation takes into account two distinct layers, the basic TiAlN and the superficial aC one. In the figure table the main data of the examined coating as well as of the superficial layer are presented. Moreover, multi-linear hardening laws determined through nanoindentations as already mentioned [11 –13], describing the material involved elastic –plastic properties are also shown. The superficial equivalent stress distribution on the basic TiAlN layer in comparison to the corresponding one of the mono-
Fig. 8. FEM model to simulate the stress field during the impact test of a coating with a superficial layer.
layer TiAlN is presented in Fig. 9a. It is evident that the superficial layer does not affect practically the stress distribution on the basic coating. The related maximum and minimum stress values at the most endangered region on the basic TiAlN coating were inserted in the already elaborated Smith diagram of the monolayer TiAlN (see Fig. 9b). It can be observed that they lie very close to the boundaries of the fracture safe diagram region of this coating. These theoretical considerations were confirmed by the experimental results. The deposition of the superficial layer does not affect the impact test performance of the basic coating as far its failure initiation concerns. The performance of the TiAlN-aC film is presented in Fig. 10. The continuous endurance load remains at the 45 daN impact force level, already in the case of the monolayer TiAlN coating found out, while the first signs of the coating damage appear close to the imprint vicinity, as the micrographs in the bottom of the figure indicate.
3. Coating fracture initiation due to local overloadings, caused by the substrate roughness
Fig. 7. Failure initiation of a TiAlN coating on a smooth surface deposited.
Roughness peaks can lead to local stress concentrations on the surface of bodies in contact [15], and thus coating damage due to overloading may occur. In order to determine the stress field during the impact test considering this effect, the surface topomorphy was taken into account in the FEM simulation. Herein various simplifications have been ap-
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radius of the contact area and the maximum pressure, depend only on the equivalent elastic properties of the contact pair. However, the bodies’ deformation differs, since it depends on the individual material stress –strain properties. Taking these considerations into account an appropriate FEM model, simulating the real contact between a ball and a plane has been established (see Fig. 11a). In this model the indenter is assumed to be rigid, since in such a case the contact elements function is more efficient [10]. The substrate has equivalent material properties (see Fig. 11b), and the pressure distribution during successive impact test stages can be determined (see Fig. 11c). However, due to the fact that the coating stress depends on the substrate deformation, which in this model is equivalent, the stress distribution in the coating cannot be accurately determined. In order to overcome this problem the pressure calculated as previously described, is applied in a further FEM model, with coating and substrate material properties, corresponding to the real ones (see Fig. 11d). An indication that the above methodology is sufficient is that in the case of smooth coating surfaces, there is a deviation in the value of the maximum equivalent stress less than 5%, compared to the corresponding one, by means of the accurate deformable indenter FEM model obtained. Various surface roughness cases have been examined according to the methodology previously described (see Fig. 12). The first one corresponds to a perfectly smooth specimen, while the further ones to specimens with different rough surfaces. The indenter has been assumed to be perfectly smooth, since its surface roughness has been found
Fig. 9. Superficial stress field on TiAlN coating and on a TiAlN-aC layered film for the continuous endurance load.
plied [9,15]. The FEM algorithm in Fig. 3 described, with deformable both bodies of the contact pair, faces serious difficulties regarding the solution convergence, if the surfaces in contact are not even. Furthermore, the roughness geometry is not axisymmetric, since its peaks in general do not revolve around the center of the contact area. A FEM model that can simulate precisely the surface roughness topomorphy should be a 3-dimensional one, which in turn is prohibitive concerning computer resources. In order to approach the stress field during the impact test of coatings on rough substrates developed, the following procedure has been carried out. According to the Hertzian theory, the pair of two elastic bodies in contact can be substituted by a rigid body in contact with a further one, having equivalent material properties [9]. In this case, the stress field remains unaffected and the magnitudes describing the pressure distribution, i.e. the
Fig. 10. Impact test performance of a TiAlN-aC coating on a smooth surface deposited.
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As a result the first signs of the coating damage initiation are expected in spots randomly distributed throughout the contact area. However, at lower roughness values, as for example in the specimen cases with Raf0.017 Am and 0.07 Am, respectively, the stress peaks due to roughness are smaller than the corresponding ones at the imprint vicinity and the coating seems to behave as deposited on a perfectly smooth surface. The increase factor of the maximum equivalent stress due to roughness is defined as the ratio of the maximum equivalent stress on the coating surface by the corresponding one of an ideal smooth coating surface divided. This ratio vs. the impact load and the surface roughness Ra is presented in Fig. 14. It is evident that at low impact loads the developed stress field is very sensitive against roughness peaks, mainly due to the limited number of peaks of the coating surface coming in contact with the indenter. As the contact load increases, the material in the aforementioned roughness peaks regions is elastically and furthermore plastically deformed. Whenever a greater number of peaks are coming in contact with the ball indenter, the contact area increases and the local stress concentrations are diminished. The impact test performance of a TiAlN, 2 Am thick coating on a rather rough surface of a 100Cr6 steel
Fig. 11. Impact test FEM model to simulate the contact in case of a coating on a rough surface deposited.
out through measurements to be very low (Ra0.01 Am). As the surface roughness increases, the pressure distribution becomes more uneven, as demonstrated in the figure. The pressure distribution peaks cause local high stress concentrations, close to the surface, as Fig. 13 demonstrates, according to the superficial equivalent stress distribution. In the case that the coating is deposited on rather rough surfaces (Raf0.15 Am), the stress peaks caused by roughness, are high enough to overpass the maximum stress value in the circumference encountered, due to coating bending.
Fig. 12. Roughness and pressure distribution during the impact test on coated surfaces.
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Fig. 15. Impact test performance of a TiAlN coating on a relatively rough specimen deposited.
Fig. 13. Stress distribution during the impact test on various rough surfaces.
(Raf0.15 Am) deposited, was examined and the corresponding results are presented in Fig. 15. The increased surface roughness leads to the reduction of the continuous endurance load from 45 daN (see Fig. 7) down to 35 daN. At the initial stage of the coating failure, damaged coating regions occur throughout the contact area, as the micrographs in the figure indicate. This fact is in good agreement with the previously described FEM supported calculation results. The visible spots correspond to surface asperities, which undertake the main portion of the contact pressure and, therefore fail faster. 4. Contribution of the abrasive wear to the coating damage initiation during the impact test
Fig. 14. Roughness effect on the maximum stress on the coating during the impact test encountered.
The failure initiation during the impact test of PVD and CVD hard coatings on smooth surfaces usually derives from fatigue phenomena. However, in relatively soft PVD coating cases, especially when the surface roughness is increased, the abrasive wear can represent a predominant film failure initiation mechanism. In a ball-on-flat contact case, the radial displacement of a body depends on its material properties [9], and a relative sliding in the radial direction between the indenter and the coating – substrate compound is expected. This relative sliding is directly related to wear phenomena and its reliable estimation cannot be accurately and effectively achieved by means of FEM supported numerical methodologies. However, the corresponding analytical equations for the case of an elastic contact between a ball and a plane lead more effectively to precise results.
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Fig. 16. Impact test imprint geometry approximation.
Due to Hertzian theory limitations, which assume a zero friction coefficient, the friction effect is neglected in the following relative sliding calculations. Herein the contact between the impact test indenter and the coating –substrate compound cannot be considered as a point contact of elastic half-spaces, due to the presence of the coating. However, when the coating thickness is small compared to the contact circle radius, its influence on the pressure distribution is small [16] and, therefore the coating contribution to the displacement of the two bodies can be neglected. Moreover the contact occurring during the impact test is usually accompanied by severe plastic deformation of the substrate, which after the first impact behaves as a hardened elastic body, with modified geometry. In Fig. 16 the imprint on a 100Cr6 steel surface, by a K05/K20 ball (Rball = 2.5 mm) with a 45 daN force induced, is demonstrated. This imprint is well approached, at least in the major part of the contact area, by a concave spherical surface, having in the present case a 6.532 mm large radius. In order to estimate the occurring radial sliding between a K05/K20 ball indenter (Rball = 2.5 mm) and a thin coating on a 100Cr6 steel deposited during the impact test at a force of Fmax = 45 daN, an elastic contact between the indenter and an elastic concave specimen surface with a radius of 6.532 mm is assumed. The radial displacement of a surface point at a distance x from the contact circle center, is given by the following equation [9]:
ur ð xÞ¼
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As the indenter and the substrate have different elastic modulus, a radial displacement difference Dur between the points of the bodies coming in contact is expected. During one impact this difference is encountered twice, once during the loading and once in the unloading stage. In Fig. 17 the total radial displacement difference after 106 impacts is presented as well as the pressure distribution at the maximum impact force calculated. The radial displacement difference, however, does not correspond directly to a sliding, as in Fig. 18 explained. As the sphere penetrates into the substrate, the contact area radius is increasing from a zero value to the maximum one. Herein ball indenter point 1, at which the maximum radial displacement difference is encountered, comes directly in contact with the specimen at the final stage of penetration without previous sliding. As a result the sliding displacement at each indenter or specimen point, after their contact up to the maximum contact region occurring has to be considered for the abrasive wear calculations. Hence the penetration procedure is divided into a number of steps. For each step, the elementary change of the radial displacement difference dDur is calculated and multiplied by the instant pressure at this point, after its contact with the counter surface. In this way the elemen-
ð1 2vÞð1 þ vÞ a2 pmax f1ð1 x2 =a2 Þ3=2 g; 3E x ð1aÞ
for x V a
for points inside the contact circle. For points outside the contact region the corresponding equation is:
ur ð x Þ ¼
ð1 2vÞð1 þ vÞ a2 pmax ; 3E x
for x > a
ð1bÞ
Fig. 17. Overall radial displacement difference and pressure distribution after 106 impacts.
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neglected in the previous considerations, reduces the occurring sliding. As a result the wear is reduced, while a deceleration of the wear development rate should be expected throughout the contact area. In this way, in soft coatings cases or at elevated roughness values, and if the substrate and the indenter have different elasticity moduli the contribution of abrasive wear to the film failure during the impact test is potential. A related experimental result is illustrated in the lower part of Fig. 19, where the failure mechanism of the superficial soft aC:H layer on a CrN basic coating deposited is demonstrated.
5. Conclusions
Fig. 18. Radial displacement at different stages of the indentation during the impact test.
The film failure initiation mechanisms during the impact test encountered, were examined experimentally and analytically. The damage of thin hard PVD coatings on relatively smooth steel surfaces deposited, starts very close to the imprint vicinity. The introduced FEM supported analysis revealed a stress concentration in these regions. If a top layer is deposited on a basic coating, the film failure initiation is not affected, since the stress field in the basic film remains almost stable. As the substrate roughness increases, the pressure distribution becomes uneven and local stress concentrations all over the contact area are encountered. At higher roughness values these concentrations can cause a coating material overload and become the predominant damage reason with a
tary surface abrasive wear can be determined according to the following equation: dW ¼ pd D ur
ð2Þ
Fig. 19 shows the occurring abrasive wear magnitude (WM) on the radial contact region direction, defined as the sum of the products of Eq. (2). X WM ¼ ð p*d D ur Þ ð3Þ This relation is very similar to the Archard equation [17], according to which, the wear w is directly proportional to the applied load P and the sliding distance s: w¼K
P *s H
ð4Þ
whereas H is the material hardness and K a parameter, depending on the bodies properties coming in contact. According to the results presented in Fig. 19, if the abrasive wear is a predominant film failure initiation mechanism, then the coating damage should be expected in the middle between the contact area centre and the imprint circumference. It should be stressed that friction, which is
Fig. 19. Estimated wear magnitude distribution in the contact area after 106 impacts.
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significant decrease of the coating performance during the impact test. In cases of relatively soft coatings or of surfaces with elevated roughness and different indenter and substrate elastic moduli, the contribution of the abrasive wear to the film fracture initiation during the impact test has to be considered. In such cases the coating failure is observed in regions, between the center and the circumference of the contact region.
6. Nomenclature aC DLC E EDX FEM Fmax p pmax PVD Ra Rball SEM Seqv t ur v WM x a Dur rr ru rz
amorphous carbon diamond like carbon elasticity modulus (GPa) energy dispersive X-ray spectroscopy finite elements method maximum impact force (daN) pressure (GPa) maximum pressure (GPa) physical vapour deposition center line average (Am) indenter diameter (mm) scanning electron microscopy von Mises equivalent stress (GPa) coating thickness (Am) superficial radial displacement (mm) Poisson ratio wear magnitude (kN/mm) radial distance from the center of contact area (mm) contact area radius (mm) radial displacement difference (mm) principle stress in the radial direction (GPa) principle stress in the circumferential direction (GPa) principle stress in the direction perpendicular to contact area (GPa)
Acknowledgments The research works in the present paper, have been conducted in the frame of the BRITE-EURAM project BE96-3398- ‘REFINE’ (Reliable Environmentally Friendly spINdle bEarings for high speed applications). The authors would especially like to thank, among others, the participating company to this project CemeCon AG (Germany).
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References [1] R. Bunshah (Ed.), Handbooks of Hard Coatings, Noyes Publications, 2001. [2] O. Knotek, B. Bosserhoff, A. Schrey, T. Leyendecker, O. Lemmer, S. Esser, A new technique for testing the impact load of thin films: the coating impact test, Surf. Coat. Technol. 54 – 55 (1992) 102 – 107. [3] K.-D. Bouzakis, N. Michailidis, A. Lontos, A. Siganos, S. Hadjiyiannis, G. Giannopoulos, et al., Coating dynamic cohesion, adhesion and creep properties characterization through the impact tester, Z. Metallkd. 92 (10) (2001) 1180 – 1185. [4] K.D. Bouzakis, N. Vidakis, K. David, The concept of an advanced impact tester supported by evaluation software for the fatigue strength characterization of hard layered media, Thin Solid Films 355 – 356 (1999) 322 – 329. [5] K.-D. Bouzakis, N. Michailidis, S. Hadjiyiannis, K. Efstathiou, E. Pavlidou, G. Erkens, et al., Improvement of PVD coated inserts cutting performance, through appropriate mechanical treatments of substrate and coating surface, Surf. Coat. Technol. 146 – 147 (2001) 443 – 450. [6] E. Lugscheider, O. Knotek, C. Wolff, S. Ba¨rwulf, Structure and properties of PVD-coatings by means of impact tester, Surf. Coat. Technol. 116 – 119 (1999) 141 – 146. [7] K.-D. Bouzakis, N. Vidakis, T. Leyendecker, O. Lemmer, H.-G. Fuss, G. Erkens, Determination of the fatigue behaviour of thin hard coatings using the impact test and a FEM simulation, Surf. Coat. Technol. 86 – 87 (1996) 549 – 556. [8] K.-D. Bouzakis, N. Vidakis, T. Leyendecker, G. Erkens, R. Wenke, Determination of the fatigue properties of multiplayer PVD coatings on various substrates, based on the impact test and its FEM simulation, Thin Solid Films 308 – 309 (1997) 315 – 322. [9] K.L. Johnson, Contact Mechanics, Cambridge University Press, 1992. [10] ANSYS 5.7.1 Online Help Documentation, Element Reference. [11] K.-D. Bouzakis, N. Michailidis, G. Erkens, Thin hard coatings stress – strain curves determination through a FEM supported evaluation of nanohardness results, Surf. Coat. Technol. 142 – 144 (2001) 102 – 109. [12] K.-D. Bouzakis, N. Michailidis, S. Hadjiyiannis, G. Skordaris, G. Erkens, The effect of the specimen roughness and the indenter tip geometry on the determination accuracy of thin hard coatings stress – strain laws by means of nanoindentations, Mater. Charact. 49 (2) (2003) 149 – 156. [13] K.-D. Bouzakis, N. Michailidis, S. Hadjiyiannis, G. Skordaris, G. Erkens, A continuous FEM simulation of the nanoindentation to determine actual indenter tip geometries, material elastoplastic deformation laws and universal hardness, Z. Metallkd. 93 (2002) 862 – 869. [14] A. Siganos, Experimental and analytical investigation of failure mechanisms encountered during the impact test for the determination of cohesion properties of thin hard PVD coatings, PhD dissertation, Aristotles University of Thessaloniki, 2003. [15] A. Michailidis, V. Bakolas, N. Drivakos, Subsurface stress field of a dry line contact, 12th International Colloquium of Tribology, Ostfildern, Germany (1999). [16] N. Vidakis, Determination of the fatigue strength of thin hard coatings for the prediction of their life time on hybrid bearing steel races, used in high speed spindles of machine tools, PhD dissertation, Aristotles University of Thessaloniki, 1997. [17] R. Liu, D.Y. Li, Modification of Archard’s equation by taking into account of elastic/pseudoelastic properties of materials, Wear 251 (2001) 956 – 964.
Fracture initiation mechanisms of thin hard coatings during the impact test K.-D. Bouzakis *, A. Siganos Laboratory for Machine Tools and Manufacturing Engineering, Mechanical Engineering Department, Aristoteles University of Thessaloniki, GR-54124 Thessaloniki, Greece Received 25 February 2003; accepted in revised form 24 December 2003 Available online 18 March 2004
Abstract The impact test can be successfully applied for the characterization of thin and thick coatings fatigue and creep properties, respectively. In the present paper experimental and analytical investigations concerning thin films fracture initiation during this test are introduced. The damage initiation of thin hard coatings deposited on smooth surfaces is mainly induced by fatigue mechanisms and characterized by the removal of film segments in the contact area between the ball indenter and the specimen in highly superficially loaded contact regions during the impact test. Herein the currently used coating stress determination methodologies, when applied, lead in most cases to sufficiently accurate results. Superficial PVD layers deposited on a basic coating do not affect the occurring stress field practically and thus no differences concerning the fracture initiation in the basic coating are encountered. When the surface roughness increases, localized stress concentrations throughout the contact area occur, resulting in film failures in randomly distributed spots, at predictable high impact loads. In cases of relatively soft PVD films or at elevated substrate roughness, the abrasive wear contribution to the coating damage initiation has to be considered. In such cases, as will be explained, the first film fracture appears in a region, close to the middle of the contact radius. D 2004 Elsevier B.V. All rights reserved. Keywords: Impact test; Thin hard coatings; Fracture initiation
1. Introduction Coatings comprise one of the most up-to-date technologies used to enhance the superficial properties of machine elements and components of tools, etc. [1]. In all these applications the knowledge of the coating material properties is of great importance. The impact test is used successfully for coating fatigue and creep property characterization [2 – 6]. Its working principle is shown in Fig. 1. An oscillating indenter, usually a cemented carbide ball, penetrates successively into the coated specimen at a constant maximum force. During the test the impact force is continuously monitored with the aid of a piezoelectric dynamometer. The impact load and the number of impacts may induce a coating failure. Fig. 2 shows the impact force vs. the number of impacts leading to coating failure. The coating failure of thin hard coatings on low roughness substrates, is initiated mainly by fatigue mechanisms and a continuous endurance load can be * Corresponding author: Tel.: +30-2310-996079, +30-2310-996021; fax: +30-2310-996059. E-mail address: [email protected] (K.-D. Bouzakis). 0257-8972/$ - see front matter D 2004 Elsevier B.V. All rights reserved. doi:10.1016/j.surfcoat.2003.12.028
determined, corresponding to the maximum force, at which no coating damage after 106 impacts is encountered. With the aid of FEM simulation of the impact test at the maximum ball-indenter penetration and during the specimen relaxation between successive impacts, Smith and Woehler diagrams of the examined coating materials have been established, as indicated in Fig. 2 [3,4,7,8]. In high substrate roughness cases, stress concentrations occur, resulting in potential local coating overloads, beyond the film fracture strength. Moreover, in soft thin coating cases or at elevated substrate roughness, abrasive wear may contribute also to the coating damage initiation during the impact test. In the following sections, experimental and analytical investigations explaining the film failure mechanisms previously mentioned, in various film cases, will be introduced.
2. Coating fracture initiation mechanisms due to fatigue During the impact test a severe plastic deformation of the substrate may occur. In order to determine the stress field in the coating – substrate compound, FEM supported algo-
K.-D. Bouzakis, A. Siganos / Surface & Coatings Technology 185 (2004) 150–159
Fig. 1. The impact tester in the investigations applied.
rithms are used. The FEM methodologies in [4,7,8] presented are based on the application of a semi-elliptical pressure distribution on the contact circle, the radius of which is experimentally determined. The coating stress field is only at the maximum contact force and during the relaxation stage examined. However, in cases of elastic – plastic contact, the pressure distribution is flattened and not semi-elliptical [9]. In the frame of the investigations de-
Fig. 2. Impact test simulation to elaborate coatings Smith and Woehler diagrams.
151
scribed, a new FEM model has been developed, presented in Fig. 3, considering not a pressure distribution, but an elastic – plastic ball indenter penetration into an elastic – plastic film-substrate compound as well. The indenter contact with the coated specimen was described with the aid of contact elements and the occurring pressure distribution determined [10]. The validity of this impact test simulation was verified through the comparisons between the calculated and the measured contact area magnitudes. Using this model, the fracture initiation during the impact test of a TiAlN film deposited on the typical bearing steel 100Cr6, was investigated. Multi-linear hardening laws associated to the applied materials, i.e. the carbide ball indenter, the coatings and the substrate have been applied. These laws were obtained through nanoindentation tests [11– 13]. Fig. 4 illustrates the superficial stress distribution on the coating at successive stages of the ball indentation, as well as during the relaxation stage. In order to determine the most loaded region inside the contact circle, the stress field development at all these ball penetration stages is considered and the deriving maximum superficial stress distribution is shown in Fig. 5a. This stress distribution does not coincide with the corresponding one when the impact force reaches its maximum value, which is also illustrated in the figure. According to these results, the most loaded region is shifted slightly towards the contact area center, just before the maximum impact load. Hence, the maximum von Mises equivalent stresses, according to the differences in principal stresses, are encountered very close to the imprint vicinity.
Fig. 3. The developed impact test FEM model with deformable ball indenter.
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between the results through the elastic plastic deformable ball indenter and the semi-elliptical pressure distribution FEM models achieved is less than 7%. Due to this fact both FEM models can be reliably applied to determine the contact stress and to elaborate fatigue diagrams. If a high substrate plastic deformation during the impact test occurs, as for example in the case of a 2.5 mm diameter ball indenter application at elevated impact loads, the results deviation increases and the use of the deformable ball model in order to obtain more accurate results is recommended [14]. Experimental results concerning the film failure initiation mechanisms on smooth substrates (Raf0.017 Am) are in Fig. 7 illustrated. The continuous endurance of TiAlN coating was found to be at an impact force of 45 daN. The coating fatigue fracture initiation can be accurately revealed by means of SEM micrographs, whereas in the indicated regions in Fig. 7, the substrate material could be detected by means of EDX microanalyses. It is obvious that at the coating damage initiation small coating segments are removed close to the imprint vicinity. In these contact area regions the superficial equivalent stresses possess a maximum, according to the previously described FEM calculations results. In general, the tribological behaviour can be enhanced through the deposition of top DLC layers, due to their low friction coefficient. In order to investigate the influence of
Fig. 4. Superficial stress distribution at intermediate stages during the ball indentation.
The equivalent stress increase at the crater vicinity is caused by the deformation occurring in the substrate, which forces the coating to ‘bend’. However, the principle stress absolute values are maximized in the center of the contact circle (see Fig. 5b). As already mentioned, the FEM methodologies applied to determine the stress field in the coatings during the impact test [4,7,8] assume a semi-elliptical pressure distribution at the maximum impact force. A comparison of the maximum stress values obtained with the aid of the described deformable ball indenter FEM model and calculated by means of the methods in [4,7,8], has been performed in several coating – substrate cases. Some characteristic results are demonstrated in Fig. 6. For the applied impact forces, which are usual for the impact test, the substrate plastic deformation does not reach extreme values and the deviation
Fig. 5. Superficial maximum and minimum stress distribution during the impact test.
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Fig. 6. Maximum superficial Seqv stresses, by means of various FEM models calculated.
such layers on the impact performance of TiAlN coatings, a top amorphous carbon (aC) layer was deposited, through the application of a carbon target during the PVD process. The effect of the top layers on the stress field developed in the basic coating was determined by means of the FEM model presented in Fig. 3, which was properly modified. As in Fig. 8 illustrated, the FEM simulation takes into account two distinct layers, the basic TiAlN and the superficial aC one. In the figure table the main data of the examined coating as well as of the superficial layer are presented. Moreover, multi-linear hardening laws determined through nanoindentations as already mentioned [11 –13], describing the material involved elastic –plastic properties are also shown. The superficial equivalent stress distribution on the basic TiAlN layer in comparison to the corresponding one of the mono-
Fig. 8. FEM model to simulate the stress field during the impact test of a coating with a superficial layer.
layer TiAlN is presented in Fig. 9a. It is evident that the superficial layer does not affect practically the stress distribution on the basic coating. The related maximum and minimum stress values at the most endangered region on the basic TiAlN coating were inserted in the already elaborated Smith diagram of the monolayer TiAlN (see Fig. 9b). It can be observed that they lie very close to the boundaries of the fracture safe diagram region of this coating. These theoretical considerations were confirmed by the experimental results. The deposition of the superficial layer does not affect the impact test performance of the basic coating as far its failure initiation concerns. The performance of the TiAlN-aC film is presented in Fig. 10. The continuous endurance load remains at the 45 daN impact force level, already in the case of the monolayer TiAlN coating found out, while the first signs of the coating damage appear close to the imprint vicinity, as the micrographs in the bottom of the figure indicate.
3. Coating fracture initiation due to local overloadings, caused by the substrate roughness
Fig. 7. Failure initiation of a TiAlN coating on a smooth surface deposited.
Roughness peaks can lead to local stress concentrations on the surface of bodies in contact [15], and thus coating damage due to overloading may occur. In order to determine the stress field during the impact test considering this effect, the surface topomorphy was taken into account in the FEM simulation. Herein various simplifications have been ap-
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radius of the contact area and the maximum pressure, depend only on the equivalent elastic properties of the contact pair. However, the bodies’ deformation differs, since it depends on the individual material stress –strain properties. Taking these considerations into account an appropriate FEM model, simulating the real contact between a ball and a plane has been established (see Fig. 11a). In this model the indenter is assumed to be rigid, since in such a case the contact elements function is more efficient [10]. The substrate has equivalent material properties (see Fig. 11b), and the pressure distribution during successive impact test stages can be determined (see Fig. 11c). However, due to the fact that the coating stress depends on the substrate deformation, which in this model is equivalent, the stress distribution in the coating cannot be accurately determined. In order to overcome this problem the pressure calculated as previously described, is applied in a further FEM model, with coating and substrate material properties, corresponding to the real ones (see Fig. 11d). An indication that the above methodology is sufficient is that in the case of smooth coating surfaces, there is a deviation in the value of the maximum equivalent stress less than 5%, compared to the corresponding one, by means of the accurate deformable indenter FEM model obtained. Various surface roughness cases have been examined according to the methodology previously described (see Fig. 12). The first one corresponds to a perfectly smooth specimen, while the further ones to specimens with different rough surfaces. The indenter has been assumed to be perfectly smooth, since its surface roughness has been found
Fig. 9. Superficial stress field on TiAlN coating and on a TiAlN-aC layered film for the continuous endurance load.
plied [9,15]. The FEM algorithm in Fig. 3 described, with deformable both bodies of the contact pair, faces serious difficulties regarding the solution convergence, if the surfaces in contact are not even. Furthermore, the roughness geometry is not axisymmetric, since its peaks in general do not revolve around the center of the contact area. A FEM model that can simulate precisely the surface roughness topomorphy should be a 3-dimensional one, which in turn is prohibitive concerning computer resources. In order to approach the stress field during the impact test of coatings on rough substrates developed, the following procedure has been carried out. According to the Hertzian theory, the pair of two elastic bodies in contact can be substituted by a rigid body in contact with a further one, having equivalent material properties [9]. In this case, the stress field remains unaffected and the magnitudes describing the pressure distribution, i.e. the
Fig. 10. Impact test performance of a TiAlN-aC coating on a smooth surface deposited.
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As a result the first signs of the coating damage initiation are expected in spots randomly distributed throughout the contact area. However, at lower roughness values, as for example in the specimen cases with Raf0.017 Am and 0.07 Am, respectively, the stress peaks due to roughness are smaller than the corresponding ones at the imprint vicinity and the coating seems to behave as deposited on a perfectly smooth surface. The increase factor of the maximum equivalent stress due to roughness is defined as the ratio of the maximum equivalent stress on the coating surface by the corresponding one of an ideal smooth coating surface divided. This ratio vs. the impact load and the surface roughness Ra is presented in Fig. 14. It is evident that at low impact loads the developed stress field is very sensitive against roughness peaks, mainly due to the limited number of peaks of the coating surface coming in contact with the indenter. As the contact load increases, the material in the aforementioned roughness peaks regions is elastically and furthermore plastically deformed. Whenever a greater number of peaks are coming in contact with the ball indenter, the contact area increases and the local stress concentrations are diminished. The impact test performance of a TiAlN, 2 Am thick coating on a rather rough surface of a 100Cr6 steel
Fig. 11. Impact test FEM model to simulate the contact in case of a coating on a rough surface deposited.
out through measurements to be very low (Ra0.01 Am). As the surface roughness increases, the pressure distribution becomes more uneven, as demonstrated in the figure. The pressure distribution peaks cause local high stress concentrations, close to the surface, as Fig. 13 demonstrates, according to the superficial equivalent stress distribution. In the case that the coating is deposited on rather rough surfaces (Raf0.15 Am), the stress peaks caused by roughness, are high enough to overpass the maximum stress value in the circumference encountered, due to coating bending.
Fig. 12. Roughness and pressure distribution during the impact test on coated surfaces.
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Fig. 15. Impact test performance of a TiAlN coating on a relatively rough specimen deposited.
Fig. 13. Stress distribution during the impact test on various rough surfaces.
(Raf0.15 Am) deposited, was examined and the corresponding results are presented in Fig. 15. The increased surface roughness leads to the reduction of the continuous endurance load from 45 daN (see Fig. 7) down to 35 daN. At the initial stage of the coating failure, damaged coating regions occur throughout the contact area, as the micrographs in the figure indicate. This fact is in good agreement with the previously described FEM supported calculation results. The visible spots correspond to surface asperities, which undertake the main portion of the contact pressure and, therefore fail faster. 4. Contribution of the abrasive wear to the coating damage initiation during the impact test
Fig. 14. Roughness effect on the maximum stress on the coating during the impact test encountered.
The failure initiation during the impact test of PVD and CVD hard coatings on smooth surfaces usually derives from fatigue phenomena. However, in relatively soft PVD coating cases, especially when the surface roughness is increased, the abrasive wear can represent a predominant film failure initiation mechanism. In a ball-on-flat contact case, the radial displacement of a body depends on its material properties [9], and a relative sliding in the radial direction between the indenter and the coating – substrate compound is expected. This relative sliding is directly related to wear phenomena and its reliable estimation cannot be accurately and effectively achieved by means of FEM supported numerical methodologies. However, the corresponding analytical equations for the case of an elastic contact between a ball and a plane lead more effectively to precise results.
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Fig. 16. Impact test imprint geometry approximation.
Due to Hertzian theory limitations, which assume a zero friction coefficient, the friction effect is neglected in the following relative sliding calculations. Herein the contact between the impact test indenter and the coating –substrate compound cannot be considered as a point contact of elastic half-spaces, due to the presence of the coating. However, when the coating thickness is small compared to the contact circle radius, its influence on the pressure distribution is small [16] and, therefore the coating contribution to the displacement of the two bodies can be neglected. Moreover the contact occurring during the impact test is usually accompanied by severe plastic deformation of the substrate, which after the first impact behaves as a hardened elastic body, with modified geometry. In Fig. 16 the imprint on a 100Cr6 steel surface, by a K05/K20 ball (Rball = 2.5 mm) with a 45 daN force induced, is demonstrated. This imprint is well approached, at least in the major part of the contact area, by a concave spherical surface, having in the present case a 6.532 mm large radius. In order to estimate the occurring radial sliding between a K05/K20 ball indenter (Rball = 2.5 mm) and a thin coating on a 100Cr6 steel deposited during the impact test at a force of Fmax = 45 daN, an elastic contact between the indenter and an elastic concave specimen surface with a radius of 6.532 mm is assumed. The radial displacement of a surface point at a distance x from the contact circle center, is given by the following equation [9]:
ur ð xÞ¼
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As the indenter and the substrate have different elastic modulus, a radial displacement difference Dur between the points of the bodies coming in contact is expected. During one impact this difference is encountered twice, once during the loading and once in the unloading stage. In Fig. 17 the total radial displacement difference after 106 impacts is presented as well as the pressure distribution at the maximum impact force calculated. The radial displacement difference, however, does not correspond directly to a sliding, as in Fig. 18 explained. As the sphere penetrates into the substrate, the contact area radius is increasing from a zero value to the maximum one. Herein ball indenter point 1, at which the maximum radial displacement difference is encountered, comes directly in contact with the specimen at the final stage of penetration without previous sliding. As a result the sliding displacement at each indenter or specimen point, after their contact up to the maximum contact region occurring has to be considered for the abrasive wear calculations. Hence the penetration procedure is divided into a number of steps. For each step, the elementary change of the radial displacement difference dDur is calculated and multiplied by the instant pressure at this point, after its contact with the counter surface. In this way the elemen-
ð1 2vÞð1 þ vÞ a2 pmax f1ð1 x2 =a2 Þ3=2 g; 3E x ð1aÞ
for x V a
for points inside the contact circle. For points outside the contact region the corresponding equation is:
ur ð x Þ ¼
ð1 2vÞð1 þ vÞ a2 pmax ; 3E x
for x > a
ð1bÞ
Fig. 17. Overall radial displacement difference and pressure distribution after 106 impacts.
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neglected in the previous considerations, reduces the occurring sliding. As a result the wear is reduced, while a deceleration of the wear development rate should be expected throughout the contact area. In this way, in soft coatings cases or at elevated roughness values, and if the substrate and the indenter have different elasticity moduli the contribution of abrasive wear to the film failure during the impact test is potential. A related experimental result is illustrated in the lower part of Fig. 19, where the failure mechanism of the superficial soft aC:H layer on a CrN basic coating deposited is demonstrated.
5. Conclusions
Fig. 18. Radial displacement at different stages of the indentation during the impact test.
The film failure initiation mechanisms during the impact test encountered, were examined experimentally and analytically. The damage of thin hard PVD coatings on relatively smooth steel surfaces deposited, starts very close to the imprint vicinity. The introduced FEM supported analysis revealed a stress concentration in these regions. If a top layer is deposited on a basic coating, the film failure initiation is not affected, since the stress field in the basic film remains almost stable. As the substrate roughness increases, the pressure distribution becomes uneven and local stress concentrations all over the contact area are encountered. At higher roughness values these concentrations can cause a coating material overload and become the predominant damage reason with a
tary surface abrasive wear can be determined according to the following equation: dW ¼ pd D ur
ð2Þ
Fig. 19 shows the occurring abrasive wear magnitude (WM) on the radial contact region direction, defined as the sum of the products of Eq. (2). X WM ¼ ð p*d D ur Þ ð3Þ This relation is very similar to the Archard equation [17], according to which, the wear w is directly proportional to the applied load P and the sliding distance s: w¼K
P *s H
ð4Þ
whereas H is the material hardness and K a parameter, depending on the bodies properties coming in contact. According to the results presented in Fig. 19, if the abrasive wear is a predominant film failure initiation mechanism, then the coating damage should be expected in the middle between the contact area centre and the imprint circumference. It should be stressed that friction, which is
Fig. 19. Estimated wear magnitude distribution in the contact area after 106 impacts.
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significant decrease of the coating performance during the impact test. In cases of relatively soft coatings or of surfaces with elevated roughness and different indenter and substrate elastic moduli, the contribution of the abrasive wear to the film fracture initiation during the impact test has to be considered. In such cases the coating failure is observed in regions, between the center and the circumference of the contact region.
6. Nomenclature aC DLC E EDX FEM Fmax p pmax PVD Ra Rball SEM Seqv t ur v WM x a Dur rr ru rz
amorphous carbon diamond like carbon elasticity modulus (GPa) energy dispersive X-ray spectroscopy finite elements method maximum impact force (daN) pressure (GPa) maximum pressure (GPa) physical vapour deposition center line average (Am) indenter diameter (mm) scanning electron microscopy von Mises equivalent stress (GPa) coating thickness (Am) superficial radial displacement (mm) Poisson ratio wear magnitude (kN/mm) radial distance from the center of contact area (mm) contact area radius (mm) radial displacement difference (mm) principle stress in the radial direction (GPa) principle stress in the circumferential direction (GPa) principle stress in the direction perpendicular to contact area (GPa)
Acknowledgments The research works in the present paper, have been conducted in the frame of the BRITE-EURAM project BE96-3398- ‘REFINE’ (Reliable Environmentally Friendly spINdle bEarings for high speed applications). The authors would especially like to thank, among others, the participating company to this project CemeCon AG (Germany).
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