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Variables affecting the fatigue resistance of PVD-coated components
International Journal of Fatigue 27 (2005) 1541–1550 www.elsevier.com/locate/ijfatigue
Variables affecting the fatigue resistance of PVD-coated components S. Baragettia,*, G.M. La Vecchiab, A. Terranovac a
Dipartimento di Progettazione e Tecnologie, Universita` degli Studi di Bergamo, Viale Marconi 5, 24044 Dalmine, Italy b Dipartimento di Ingegneria Meccanica, Universita` di Brescia, Via Branze 38, 25123 Brescia, Italy c Dipartimento di Meccanica, Politecnico di Milano, Via la Masa, 34, 20158 Milano, Italy Available online 2 August 2005
Abstract The effect of intrinsic properties of CrN coatings on fatigue behaviour was studied in this paper. The coating layer microhardness and the residual stresses characterising the surface film were measured and the obtained results were introduced in a numerical modelling predicting fatigue life procedure of coated components. The effect of a CrN monolayer film deposited on bulk samples, produced in 2205 duplex stainless steel, H11 tool steel or 6082 aluminium alloy was investigated. The fatigue limit of coated and uncoated samples was experimentally determined while the development of FEM models, confirmed by means of experimental tests, represents a powerful tool to predict fatigue life of coated components. The effects on the fatigue strength of coating and bulk material defects like droplets and nonmetallic inclusions were considered along with the residual stress gradient characterising the coating and evaluated by means of X-ray measurements. The influence of the substrate material plastic deformation on the integrity of the coating was evaluated too. q 2005 Elsevier Ltd. All rights reserved. Keywords: CrN-PVD coating; Fatigue; FEM modelling; Residual stress gradient; Microstructural defects
1. Introduction The beneficial effect of thin ceramic coatings deposited by means of PVD technique is well-known both in terms of improvement in strength and wear resistance and, selecting the proper surface film material, in corrosion resistance. On the other hand any coating system giving improved wear properties, may drastically reduce the fatigue life of a component due to cracks starting in the coating and propagating into the substrate material. Considering, for instance, chrome-plate parts, they show improved wear behaviour but, also, a shorter fatigue life with respect to those of uncoated parts [1]. Furthermore, it is well known that fatigue is a phenomenon strictly related to the residual stress level at the surface of a component and, therefore, coating techniques inducing compressive residual stresses can potentially enhance fatigue behaviour in comparison to surface treatments inducing tensile residual stresses. Recently some studies were addressed to the investigation of the effect of PVD coatings on the fatigue * Corresponding author. Tel.: C39 035 2052 382; fax: C39 035 562 779. E-mail address: [email protected] (S. Baragetti).
0142-1123/$ - see front matter q 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijfatigue.2005.06.011
behaviour in order to inquire into the possibility of improving the use of coated components in applications where the fracture mechanism is the discriminating project choice (typical examples are connecting rods, crankshafts and turbine rotors) [2–7]. Generally these research works have pointed out increments in the fatigue resistance at long life region of coated components with respect to uncoated ones, and such result was interpreted both by means of the constraining effect of slip deformation on the sample surface and by the high level of compressive residual stress affecting the coating [8–11]. However, the choice of the type of coating in terms of deposition technique, chemical composition and thickness is, at the moment, frequently referred to empiric consideration. Therefore, the settlement of previsional numerical models for the fatigue behaviour of coated parts in which the properties characterising substrate and coating materials are explicit would be considered of strategic importance. In fact, according to the applied loads and the type of substrate material, it would allow a simplified and optimised selection of the coating material. The choice of the substrate material is another critical step when the coated component has to guarantee not only high levels of mechanical properties but also a lightness of the final part. Considering, for instance, the automotive industry, a more widespread use of light metal alloys will be guaranteed in tribological applications like bearing plates,
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Nomenclature C a d dc d2 d* E H1 H2 H3 HV KIc Kres DK
material constant crack depth (mm) distance between the microstructural barriers (mm) thickness of coating (mm) depth of maximum hardness data (mm) depth from the surface (mm) Young’s modulus of the material (N/mm2) surface Vickers hardness maximum Vickers hardness core Vickers hardness Vickers hardness pffiffiffiffi fracture toughness (MPa m) stress intensity factor due to residual stresses stress intensity factor range
guide bars or seat supports in the presence of functional surface coatings able to provide good wear protection and stiffness. Nowadays several surface treatment processes have been optimised for aluminium alloys (electroless nickel plating and anodic oxidation are two typical examples of coating industrially used for mechanical applications produced in aluminium alloys) in order to obtain a better behaviour when a high corrosion and wear resistance behaviour are requested. Up to now an unsolved feature characterizes such coatings: no one of them can match the requirement of bearing sliding movements under high applied loads, condition particularly difficult when the deposed coating shows a thin thickness. Therefore, there is a steady demand for tailored coating suitable to overcome such a problem and, in the meantime, to assure high levels of fatigue resistance. The low melting point, peculiar for aluminium alloys, is another factor that must be considered to select the proper coating depositing parameters; a low deposition temperature is, in fact, a solution avoiding severe modification in the microstructure of the substrate material and, consequently, a significant reduction in the mechanical behaviour of the final coated part. A further aspect that is not to be neglected is the strong tendency to oxidation of aluminium alloys requiring a specific and accurate cleaning of the part before the deposition step. A ion cleaning carried out by means of an appropriate gas during the first period of deposition with the aim of contrasting the Al2O3 formation, in order to assure a good adhesion coating/substrate, is also requested. Starting from such assumptions, in the present work, the fatigue behaviour of two steel grades with differences in terms of hardness data and used in different applications (a 2205 duplex stainless steel and a H11 tool steel) was compared to the one of the 6082 alloy. Tests were carried out on both uncoated and CrN-PVD coated samples having a thickness ranging from 2.5 to 5 mm. The experimental
DKapp
DKeff DKth N r Ra u
q n
applied stress intensity factor range (takes into account the reduction of DKeff due to the preload applied during the experimental tests) effective stress intensity factor range (due to the presence of the residual stress field) pffiffiffiffi threshold stress intensity factor range (MPa m) number of cycles distance from the crack tip surface roughness (mm) FEM half relative displacement between the flanks of the crack, for the couple of nodes on the opposite flanks of the crack at the crack tip (mm) angular position taken with reference to the crack tip (rad) Poisson’s ratio
results were interpreted by means of microstructural characterisation and residual stress gradient evaluation. A numerical finite element model was developed and enabled us to predict fatigue life of the coated specimens only in case of absence of delamination between coating and bulk material.
2. Fatigue life evaluation of thin coated components: numerical models It is well known that for components of engines or structures used in car or motorcycle races or in aerospace industry, the capability of predicting the number of cycles until failure of structural components is very important. Bearing in mind this last consideration, authors developed a model able to predict the number of cycles until failure of thin hard coated components. The numerical model, once the initial crack depth from the surface of the component is known, enables one to calculate the stress intensity factor (including residual stress effects) and to simulate crack propagation (in a discrete way) and to evaluate the number of cycles needed to reach specified crack depths until rupture of the component. The quarter-point technique has to be used in order to enhance the accuracy of results obtained from the numerical analyses [12]. As regards the nucleation phase of defects, in this paper only simulation of surface defects was carried out. Surface cracks may be generated by brittle fracture of coating or by the presence of surface droplets or defects at least 5 mm deep. Due to the presence of residual stresses, crack nucleation is always located at surface or subsurface defects, like droplets or inclusions [13]. Numerical evaluation of the stress intensity factor, in the presence of residual stresses, can be carried out by using the value of the half relative displacement, u, of crack
S. Baragetti et al. / International Journal of Fatigue 27 (2005) 1541–1550
surfaces: rffiffiffiffiffiffiffi E 2p u KI Z 1 Cn r f ðqÞ
(1)
q 2 q f ðqÞ Z sin k C 1 K 2cos 2 2
(2)
kZ
3 Kn ; 1 Cn
k Z 3 K 4n;
for plane strain state (3) for plane stress state;
Having established the stress intensity factor from numerical simulation of crack propagation (for each crack depth taken into consideration), it is possible to calculate the stress intensity factor range and to compare it with the threshold stress intensity factor range of the material obtainable from previously published models or formulas [14–17]. If the numerical stress intensity factor range is higher than the threshold stress intensity factor range of the material, fatigue crack propagation is possible. If the crack propagates, then the fatigue crack growth rate, and the number of cycles necessary to reach specified crack depths, can be evaluated by means of existing models. The model is suitable and utilisable in all cases where no delamination between coating and substrate material can be verified, for thin hard coating (with thickness ranging from 1 to 10 mm) to thick coatings (thickness usually over 10 mm). The coating can be of any kind (PVD, CVD, Thermal spray, etc.). The model can be developed by means of a commercial code [18] and mesh refinement has to be applied only at the cracked area. The dimensions of the finite elements at the cracked area can reach the minimum value 0.5–1 mm; care should be put in the phase of mesh refinement in order to reduce the computational time and diminish the total number of degrees of freedom of the model. Substrate material and coating mechanical properties have to be assigned; only Young’s modulus and Poisson’s ratio have to be attributed just because an elastic material behaviour (both for the coating and for the substrate material) was used in all of the analyses, having assumed a limited plasticized area at the crack tip (during the fatigue propagation inside the substrate material such approximation is well verified only in the case of the H11 tool steel grade while during the propagation inside the coating such hypothesis is always substantiated). The residual stress field, induced by coating deposition process has to be applied as an initial condition to the model (residual stresses act before external applied loads). Application of residual stresses can be carried out by means of a temperature gradient procedure: a specified temperature profile, with nodal temperature assignation from surface to in-depth layers of finite elements, enabled simulation of the presence of the residual stress profile. Due to the residual stress field induced by the presence of
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the coating, with rather high compressive residual stresses at the surface, contact elements have to be used in order to avoid overlapping of the mating surfaces of the crack (one flank of the crack is prevented from penetrating the other). Introduction of contact elements enables evaluation of stress intensity factors due to residual stresses present before load application. The stress intensity factor range has to be calculated by taking into consideration the presence of residual stresses and introducing the parameter DKeffZ KIKKres; reduction of DKeff due to the preload applied during the experimental tests has to be considered too, by introducing a numerical applied stress intensity factor range DKapp. The numerical procedure enables determination of the number of cycles until failure but does not allow calculation of the fatigue limit in the case of coated components. Determination of the number of cycles until failure, for crack depths ranging from 0.5 to 1 mm to specified depths, can be carried out by using literature models developed for carburised spur gears [19–21]. Such models require knowledge of the material hardness versus depth from the surface of the component, and the value of the effective stress intensity factor range (evaluated by means of the finite element analyses) da C Z ðDK n K DKthn Þ; dN ð1 K rn Þ
for DKth % DK % KC
da C DK n KIc Z ; for KC !DK!KIc ; dN ð1Krn Þ DKIcn KDK n
(4)
(5)
DK rZ th ; KC ZðDKth KIc Þ1=2 KIc Symbols in Eqs. (4) and (5) have the following meaning: 1 H1 KH3 AZK 2 ln ; d %d2 H2 KH3 d2 AZK
1 550KH3 ln ; d Od2 H2 KH3 ðdc Kd2 Þ2
H ZðH2 KH3 Þexp½KAðd Kd2 Þ2 CH3 DKth Z2:45C3:41!10K3 H KIc Z141K1:64!10K1 H nZ4:31K8:66!10K3 HC1:17!10K5 H2 logðCÞZK10:0C1:09!10K2 HK1:40!10K5 H2 2.1. Experimental data necessary for numerical fatigue life evaluation Mechanical elastic properties (Young’s modulus and Poisson’s ratio) both of bulk material and coating have to be
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introduced as characteristics of the finite elements belonging to each material. Along with material elastic properties, fatigue life evaluation by means of numerical models needs assessment of three important parameters: residual stress gradient, microhardness and surface roughness. The residual stress field is maybe the most important parameter just because of the high compressive surface residual stresses that enable to prevent crack nucleation and propagation [22]. The residual stress field should be known from the surface of the component to inner layers of material. Authors verified, by means of X-ray diffraction measurements, that, in the case of CrN-PVD coatings having thickness 5 mm, the surface film deposition induces a high compressive surface residual stresses (fairly constant in the coating and assuming values ranging from 0.8 to 3 GPa in function of the substrate material [11,22–26]). Structural characterization was executed by means of XRD technique with Bragg–Brentano geometry. 2D images (XRD2) were also obtained by means of a D/max RAPID Rigaku microdiffractometer in order to measure the texture, the lattice constants and the residual stress values characterizing the coating. In particular for the residual stress analysis, the DRAST method was used [22]. Such method permits to measure the residual stresses active in the specimen from the evaluation of the displacement of the Debye ring; details regarding the procedure used during the tests carried out on CrN-PVD coatings can be detected in Ref. [23]. In particular to calculate the residual stress, Eq. [23] dfj K d0 1 Cy y sf sin2 j K ½s11 C s22 Z E E d0
the residual stress field is not much important as the surface value of residual stresses. At different crack depths, the numerical value of the effective stress intensity factor range depends only on the surface residual stress and the application of different experimental trends of subsurface residual stresses did not significantly affect the stress intensity factor range. Due to the steep residual stress gradient present at the interface between coating and bulk material, attention must be paid to the presence of irregularities and defects that could be source of fatigue crack nucleation and propagation. In this case, the proposed model could be suitable if the aim is to simulate the presence of subsurface cracks, especially at the interface between coating and bulk material, emanating from defects, inclusions and irregularities; anyway the crack propagation path has to be perpendicular to the surface of the component. One of the residual stress trend measured by means of X-ray diffraction by the authors is shown in Fig. 1; this is the case of a CrN-PVD coated H11 tool steel sample (four point bending test specimen dimensions 110!21! 4.4 mm) with coating thickness 5 mm. Microhardness has to be measured in order to apply literature models for fatigue crack growth evaluation of hardened surfaces [19–21]. One of the microhardness trend measured by means of a nanoindenter equipped with a Vickers indenter is shown in Fig. 2; this is the case of a microhardness trend for a CrN-PVD coated H11 tool steel sample with coating thickness 5 mm. Considering specimens in which the crack nucleation starts from the coating, surface roughness should be measured too in order to avoid introduction of opportune correction coefficients that would enable to take into consideration the effects of surface roughness. In particular, it is well known that if surface roughness exceeds RaZ 0.3 mm (ISO 1143 Standard), machine design formulas used for fatigue strength analysis of a component has to be
(6)
for biaxial stress analysis, was used introducing the elastic modulus calculated from nanoindentation measurements (CrN coating: EZ303 GPa; Poisson ratio nZ0.2). Considering the high residual stress level characterizing a sizeable part of the coating, the knowledge of the trend of
600
Coating (CrN PVD, thickness 5 µm)
300
10.500
10.490
10.480
−300 −600 −900
Bulk material (H11 tool steel)
−1200 −1500 −1800 −2100 −2400 −2700
Fig. 1. Residual stress trend for a PVD-coated H11 tool steel sample with coating thickness 5 mm.
Residual stresses [MPa]
Distance from the axis of the specimen [mm]
10.470
10.460
10.450
10.440
10.430
10.420
10.410
10.400
10.390
10.380
0
S. Baragetti et al. / International Journal of Fatigue 27 (2005) 1541–1550
1545
3000
X 2700
Coating (CrN PVD, thickness 5 µm)
2100
X Experimental data points
X
1800 1500
Bulk material (H11 tool steel)
X
1200
Hardness [HV]
2400
900
X 10.440
10.450
10.460
10.470
10.480
X
10.490
600
X
300 0 10.500
Distance from the axis of the specimen [mm] Fig. 2. Microhardness trend for a PVD-coated H11 tool steel sample with coating thickness 5 mm.
corrected by using a surface coefficient [27]. Anyway such problem does not affect the results of the proposed model just because, in the case of surface PVD coatings, fundamental is an accurate polishing before the deposition step (Ra!0.2 mm) in order to assure a high adhesion substrate/coating. For the specimens used in the present work, a Ra%0.15 was used. Surface roughness is always less than 0.15 mm (Ra%0.15 mm) for the PVD coated components studied in this paper and, in general, surface roughness has to be as low as possible in order to have good adhesion of the coating with the substrate. If the aim is to evaluate fatigue resistance in presence of surface defects (droplets of cracks) by means of numerical FEM models, that enable one to simulate the presence of such defects having minimum depth 5 mm (much higher than roughness ‘cracks’) than the propagating crack will be the one having depth 5 mm. After brittle rupture of coating (or in presence of defects at the surface of the components), defects having minimum depth 5 mm appear at the surface and surface roughness is no more the driving parameter for fatigue life evaluation. 2.2. Numerical models results: application to four point bending tests Four point bending tests were simulated and fatigue crack propagation from a surface crack having depth equal to the coating thickness was numerically calculated. With reference to H11 tool steed samples, some images related to the finite element model developed are shown in Fig. 3. The finite element models are composed of plane stress four nodes linear finite elements and, due to the symmetry in load and geometry, only one half of the specimens was modelled and symmetry boundary conditions were applied at the symmetry surface.
The values of DKth and DKapp, at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa), are shown in Table 1 [24]. Analysis of Table 1 puts in evidence how, in case of 2205 stainless steel, the stress intensity factor range is higher than the threshold stress intensity factor range only for crack depth over 5 mm. In case of H11 tool steel, the stress intensity factor range is higher than the threshold stress intensity factor range only for crack depth over 55 mm. It is worth underlining that if surface irregularities or inclusions are present in the bulk material, these defect may dramatically affect fatigue life of the component. The number of cycles until failure, calculated at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa), are shown in Table 2. The numerical model enables one to calculate the number of cycles until failure, summing the number of cycles necessary to reach specified crack depths (Table 2). Analysis of Table 2, in case of 2205 stainless steel, shows as the specimen spends almost all of its life in the phase of crack propagation from 5 to 25 mm. This is principally due to the high residual stress field, induced by the coating deposition technique, that reaches high values in the coating (the trend or residual stresses for 2205 stainless steel is similar to the one shown in Fig. 1 for H11 tool steel and in the coating the maximum residual stress measured was K1500 MPa [24]). If fatigue life is set at 2,000,000 cycles, the model enables to say that for 2205 coated specimens, the fatigue limit might be set over 600 MPa (the model predicts a fatigue life equal to 2,930,897 cycles). In case of H11 tool steel, with maximum applied stress 845 MPa, propagation
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Fig. 3. Plane stress finite element models for DK evaluation of coated specimens at different crack depths (developed for H11 tool steel specimens).
of a surface crack would be admissible only if a surface defect would be present and would have a minimum depth from the surface equal to 55 mm. In this case, the predicted number of cycles until failure would be equal to 37,670. Otherwise there is no fatigue crack propagation and the maximum applied stress 845 MPa is the minimum value of the fatigue limit of the coated component.
2.3. Coated components fatigue tests and numerical models validation Comparison between the results obtained from fatigue tests, carried out on steel specimens (H11 tool steel and 2205 duplex stainless steel) was arranged both with the aim of evaluating fatigue resistance of 2205 steel, H11 steel and
Table 1 DKth and DKapp at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa) 2205 Stainless steel 600 MPa
H11 Tool steel 845 MPa
Crack depth (mm)
pffiffiffiffi DKth (MPa m)
pffiffiffiffi DKapp (MPa m)
Crack depth (mm)
pffiffiffiffi DKth (MPa m)
pffiffiffiffi DKapp (MPa m)
5 25 50 100 500 1000
n.d. 3.00 3.06 3.10 3.30 3.30
0.00 3.05 8.54 11.75 27.96 35.71
5 30 55 105 505 1000
n.d. 4.23 5.18 3.39 3.73 3.73
0.00 2.22 2.90 8.19 22.40 43.80
S. Baragetti et al. / International Journal of Fatigue 27 (2005) 1541–1550
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Table 2 Number of cycles until failure, calculated at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa) 2205 Stainless steel, 600 MPa
H11 Tool steel, 845 MPa
Crack depth (mm)
Number of cycles
Crack depth (mm)
Number of cycles
5–25 25–50 50–100 100–500 500–1000 1000-Rupture Fatigue life
2,911,346 8164 6030 2812 2436 109 2,930,897
5–30 30–55 55–105 105–505 505–1000 1000-Rupture Fatigue life
No propagation No propagation 31,360 5520 690 100 37,670
Table 3 Chemical composition (wt%) of the materials used as substrate Material
C
Mn
Si
Cr
Ni
Mo
Mg
Al
Fe
H11 2205 6082
0.43 0.019 –
1.50 1.51 0.73
0.30 0.39 0.99
1.99 22.45 0.05
– 5.5 –
0.21 3.12 –
– – 0.78
0.031 – Balanced
Balanced Balanced 0.46
6082 aluminium alloy coated specimens and with the aim of validating the proposed analytical–numerical procedure for number of cycles to failure calculation. The chemical composition of the three alloys used as substrate materials is shown in Table 3, while the heat treatments carried out before the coating deposition and the measured hardness and tensile values are summarized in Table 4. Independently from the substrate material a monolayer of CrN was deposited using an industrial coating plant equipped with a cathodic arc apparatus. Considering the difference in the bulk material, the standard deposition temperature of 573 K, used for the production of PVD films on steel substrate, was reduced to 483 K during the deposition on 6082 alloy. To oppose the overheating of the aluminium alloy inducing an overaging of the substrate, a reduction in the film thickness, and therefore of the time requested for the deposition step, was imposed: 2.5 mm instead of the 5 mm used for the deposition on the two types of steel investigated in the present work. In the case of the tested steel grades, microhardness data resulted practically independent from the substrate material (2686 HV for H11 steel and 2567 HV for 2205 duplex stainless steel) while a certain decrease in the coating hardness was obtained testing coated 6082 samples (HVZ 2238); such variation in the coating hardness is probably due to trace back to the load applied to the indenter that, at a penetration depth from the sample coated surface equal to 0.5 mm, resulted partially influenced by the hardness of the substrate material [28]. Four point bending fatigue tests were carried out both for 2205 stainless steel and H11 tool steel while for aluminium rotating bending tests were arranged. Only four point bending tests needed the development of a device that,
mounted on the universal test machine, could enable symmetric application of loads on the specimen (Fig. 4). The device consists of a pin connection (with a roller bearing) mounted between the specimen and the test machine (Fig. 4b). Only for H11 tool steel specimens, special supports (in PTFE) had to be developed in order to guarantee an appropriate guide to the specimen during the test (Fig. 4a). Prismatic fatigue samples were produced for testing both 2205 duplex stainless steel (110!11!5.5 mm) and H11 tool steel (110!21!4.4 mm). For 6082 samples, a hourglass geometry was used (minimum diameter 10 mm and length 110 mm according to ISO 1143 standard). Stair case was applied to the results of the fatigue tests carried out on steel and aluminium samples [29]. Experimental fatigue limit, both for coated and uncoated 2205, H11 and aluminium alloy, is reported in Table 5. Numerical results, listed in Tables 1 and 2, show how there is no fatigue crack growth if the crack has a depth less than or equal to 105 mm for H11 tool steel and 25 mm for duplex stainless steel. The number of cycles until failure evaluated by means of the numerical model is 37,670 for H11 tool steel with maximum applied bending stress 845 MPa, and can be calculated only if a surface crack, or a subsurface non metallic inclusion or matrix irregularity, having depth at least 55–105 mm is present at the surface of the specimen. On the other hand, Table 4 Mechanical properties characterising the substrate materials before coating Material
Condition
Hardness
YS (MPa)
UTS (MPa)
H11
QuenchingC tempering Solution annealing T6
32 HRC
893
1006
225 HB 113 HB
513 349
790 365
2205 6082
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Fig. 4. Device developed for four points bending fatigue tests.
simulation of fatigue crack propagation in 2205 stainless steel gives a fatigue life of 2,930,897 at 600 MPa maximum bending applied stress. Experimental four points bending tests confirmed the numerical results. For H11 tool steel specimens only for specimens presenting sub-surface cracks, or inclusions like the ones shown in Fig. 5, rupture occurred before the established fatigue life limit set at 2,000,000 cycles. If the surface of the specimen presents irregularities, surface cracks or subsurface defects with depth from the surface less than 55 mm, no fatigue crack nucleation or propagation for applied stresses less than 845 MPa (the experimental fatigue limit of H11 obtained from application of stair case method) is expected. The proposed model is confirmed by the experimental stair case fatigue tests results: the specimens with initial surface crack depth less than 55–105 mm either presented a number of cycles until failure over 2,000,000 cycles, or did not present any failure (stopped tests): such specimens were considered as unbroken ones during evaluation of the fatigue limit (stair case method). As concerns the 2205 duplex stainless steel the experimental number of cycles until failure is always over 2,000,000 cycles. The experimental results confirm the number of cycles evaluated by means of the numerical approach. The number of cycles, both numerical and experimental, is higher than 2,000,000, thus confirming that the model makes it possible to predict the resistance of specimens beyond the limit number of cycles set at 2,000,000. Moreover, the experimental number of cycles
until failure for the same specimen, in the same conditions as those of the numerical model, is 3,380,500. The difference between the numerical (2,930,897 cycles) and the experimental (3,380,500 cycles) number of cycles until failure is less than 14% and allows a certain confidence in the applicability of the numerical procedure proposed [24]. Coated specimens exhibit a fatigue life increment of about 15%, both in case of stainless steel and H11 tool steel. A slight decrement in the fatigue life was observed in the case of aluminium rotating bending fatigue tests notwithstanding the high level of compressive residual stress measured at the coating (data always higher than K0.8 GPa). To interpret such a difference, the mechanism of rupture induced by the cycling load has to be examined. The nucleation of the fatigue cracks for H11 steel grade, showing quite large non-metallic inclusions, resulted promoted only in a zone sufficiently away from the interface. In such area, the effect of the compressive residual stress induced from the coating procedure can be neglected. On the other hand, 2205 duplex stainless steel grade, being characterised by a higher level of cleanness, shows a nucleation located at the coating, and, in order to
Table 5 Endurance limit values for coated and uncoated samples—experimental results obtained testing the different substrate materials Bulk material
Fatigue limit uncoated samples (MPa)
Fatigue limit coated samples (MPa)
H11 2205 6082
730G26 510G18 165G11
845G43 585G26 142G10
Fig. 5. H11 coated sample—example of nucleation from a large sub-surface non metallic inclusion.
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Fig. 6. (a) Fractured surface of 2205 four-points bending fatigue test coated sample loaded at 600 MPa, and (b) 6082 rotary bending fatigue test coated specimen loaded at 140 MPa. Arrows indicate fatigue nucleation sites.
oppose the effect of the compressive residual stresses, the application of cycling loads more than the yield strength of the substrate is needed. The weakening of the coating increased for the 6082 coated samples, where the surface hard film cracks in multiple points due to the impossibility to sustain the imposed strain as requested from the overaged aluminium alloy substrate. In Fig. 6, a comparison between two fatigue fracture surfaces typical of a coated 2205 sample and a coated 6082 specimen can be observed. Independently from the substrate material, a good adhesion steel/coating was always detected. The droplets, typical defects of PVD coatings, cannot be taken as the area of fatigue nucleation independently from the substrate material.
3. Conclusions Different CrN-PVD coated specimens, made of steel and aluminium alloys, were tested in this paper and fatigue resistance was evaluated. Experimental investigations were also used in the development of a numerical procedure able to predict the number of cycles until failure of coated components. An enhancement in the fatigue limit of steel coated specimens was obtained with respect to uncoated ones. The measured increment is about 15%. For aluminium-coated specimens, no enhancement in the fatigue limit was detected with respect to uncoated ones. However, an enhancement in the fatigue limit might be obtained by opportunely modifying the deposition parameters in order to limit overaging of the substrate material. Such thesis should be verified carrying out further tests. The numerical procedure, based on finite element modeling, enables one to calculate the number of cycles until failure of coated components by simulating, in a discrete way, crack propagation. The trustworthiness of such procedure was verified through the execution of experimental fatigue tests.
Acknowledgements The authors thank S. Troglio of Vacuum Surtec for the preparation of the coated samples and the Chemical
Structural Laboratory of the Brescia University for the X-ray measurements.
References [1] Nascimento MP, Souza RC, Pigatin WL, Voorwald HJC. Effects of surface treatments on the fatigue strength of AISI 4340 aeronautical steel. Int J Fatigue 2001;23:607–18. [2] Ferreira JAM, Costa JDM, Lapa V. Fatigue behaviour of 42CrMo4 steel with PVD coatings. Int J Fatigue 1997;19(4):293–9. [3] Su YL, Yao SH, Wei CS, Wu CT. Evaluation of the tension and fatigue behaviour of various PVD coated materials. Thin Solid Films 1998;322:218–24. [4] Schlund P, Kindermann P, Schulte R, Sockel HG, Schleinkofer U, Gorting K, Heinrich W. Mechanical behaviour of PVD/CVD coated hard metals under cycling loads. Int J Refractory Metal Hard Mater 1999;17:179–85. [5] Berrios JA, Teer DG, Puchi-Cabrera ES. Fatigue properties of a 316L stainless steel coated with different TiNx deposits. Surf Coat Technol 2001;148:179–90. [6] Kim HR, Suh CM, Murakami RI, Chung CW. Effect of intrinsic properties of ceramic coatings on fatigue behavior of Cr–Mo–V steels. Surf Coat Technol 2003;171:15–23. [7] Berrios-Ortiz JA, La Barbera-Sosa JG, Teer DG, Puchi-Cabrera ES. Fatigue properties of a 316L stainless steel coated with different ZrN deposits. Surf Coat Technol 2004;179:145–57. [8] Ejiri S, Sasaki T, Hirose Y. X-ray stress measurement for TiN films evaporated by PVD. Thin Solid Films 1997;307:178–82. [9] Murotani T, Hirose H, Sasaki T, Okazaki K. Study on stress measurement of PVD-coating layer. Thin Solid Films 2000;377–378: 617–20. [10] Suh C-M, Hwang B-W, Murakami R-I. Behaviors of residual stress and high-temperature fatigue life in ceramic coatings produced by PVD. Mater Sci Eng 2003;A343:1–7. [11] Djouadi MA, Nouveau C, Banakh O, Sanjine´s R, Le´vy F, Nouet G. Stress profiles and thermal stability of CrxNy films deposited by magnetron sputtering. Surf Coat Technol 2002;151–152:510–4. [12] Benzley SE. Representation of singularities with isoparametric finite elements. Int J Numer Meth Eng 1974;8:537–45. [13] Gelfi M, La Vecchia GM, Lecis N, Troglio S. Relationship between through thickness residual stress of CrN-PVD coatings and fatigue nucleation sites. Surf Coat Technol 2005;192:263–8. [14] Paris PC, Erdogan F. A critical analysis of crack propagation laws. J Basic 1963;85:528–34. [15] Murakami Y, Endo M. Effects of defects, inclusions and inhomogeneities on fatigue strength. Int J Fatigue 1994;16:519–33. [16] Miller KJ. The short crack problem. Fatigue Eng Mater Struct 1982; 5(3):223–32.
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[17] El-Haddad MH, Smith KN, Topper TH. Fatigue crack propagation of short cracks. ASME J Eng Mater Technol 1979;101:42–6. [18] Structural Research and Analysis Corporation, Cosmos/M 2.7, user’s guide; 2001. [19] Inoue K, Lyu S, Deng G, Kato M. Fracture mechanics based evaluation of the effect of the surface treatments on the strength of carburized gears. Proc VDI Berichte 1996;1320:357–69. [20] Inoue K, Kato M, Yamanaka M. Fatigue strength and crack growth of carburized and shot peened spur gears. Proc Power Transm Eng Conf (ASME) 1989;663–8. [21] Kato M, Deng G, Inoue K, Takatsu N. Evaluation of the strength of carburized spur gear teeth based on fracture mechanics. JSME Int J— Series C 1993;36(2):233–40. [22] Gelfi M, Bontempi E, Roberti R, Depero LE. Acta Mater 2004;52: 583–9. [23] Gelfi M, La Vecchia GM, Lecis N, Troglio S. Surf Coat Technol 2005; 192:263–8.
[24] Baragetti S, La Vecchia GM, Terranova A. Fatigue behavior and FEM modelling of thin-coated components. Int J Fatigue 2003;25: 1229–38. [25] Elstner F, et al. A comparative study of structure and residual stress in chromium nitride films deposited by vacuum arc evaporation, ion plating and DC magnetron sputtering. Phys Status Solidi A 1996;158: 505–21. [26] Cunha L, et al. Residual strees, surface defects and corrosion resistance of CrN hard coatings. Surf Coat Technol 1999;111: 158–62. [27] Juvinall RC, Marshek KM. Fundamentals of machine component design. New York: Wiley; 1991. [28] Ichimura H, Ando I. Mechanical properties of arc-evaporated CrN coatings: nanoindentation hardness and elastic modulus. Surf Coat Technol 2001;145:88–93. [29] Dixon WJ, Massey FJ. Introduction to statistical analysis. New York: McGraw Hill; 1969.
Variables affecting the fatigue resistance of PVD-coated components S. Baragettia,*, G.M. La Vecchiab, A. Terranovac a
Dipartimento di Progettazione e Tecnologie, Universita` degli Studi di Bergamo, Viale Marconi 5, 24044 Dalmine, Italy b Dipartimento di Ingegneria Meccanica, Universita` di Brescia, Via Branze 38, 25123 Brescia, Italy c Dipartimento di Meccanica, Politecnico di Milano, Via la Masa, 34, 20158 Milano, Italy Available online 2 August 2005
Abstract The effect of intrinsic properties of CrN coatings on fatigue behaviour was studied in this paper. The coating layer microhardness and the residual stresses characterising the surface film were measured and the obtained results were introduced in a numerical modelling predicting fatigue life procedure of coated components. The effect of a CrN monolayer film deposited on bulk samples, produced in 2205 duplex stainless steel, H11 tool steel or 6082 aluminium alloy was investigated. The fatigue limit of coated and uncoated samples was experimentally determined while the development of FEM models, confirmed by means of experimental tests, represents a powerful tool to predict fatigue life of coated components. The effects on the fatigue strength of coating and bulk material defects like droplets and nonmetallic inclusions were considered along with the residual stress gradient characterising the coating and evaluated by means of X-ray measurements. The influence of the substrate material plastic deformation on the integrity of the coating was evaluated too. q 2005 Elsevier Ltd. All rights reserved. Keywords: CrN-PVD coating; Fatigue; FEM modelling; Residual stress gradient; Microstructural defects
1. Introduction The beneficial effect of thin ceramic coatings deposited by means of PVD technique is well-known both in terms of improvement in strength and wear resistance and, selecting the proper surface film material, in corrosion resistance. On the other hand any coating system giving improved wear properties, may drastically reduce the fatigue life of a component due to cracks starting in the coating and propagating into the substrate material. Considering, for instance, chrome-plate parts, they show improved wear behaviour but, also, a shorter fatigue life with respect to those of uncoated parts [1]. Furthermore, it is well known that fatigue is a phenomenon strictly related to the residual stress level at the surface of a component and, therefore, coating techniques inducing compressive residual stresses can potentially enhance fatigue behaviour in comparison to surface treatments inducing tensile residual stresses. Recently some studies were addressed to the investigation of the effect of PVD coatings on the fatigue * Corresponding author. Tel.: C39 035 2052 382; fax: C39 035 562 779. E-mail address: [email protected] (S. Baragetti).
0142-1123/$ - see front matter q 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijfatigue.2005.06.011
behaviour in order to inquire into the possibility of improving the use of coated components in applications where the fracture mechanism is the discriminating project choice (typical examples are connecting rods, crankshafts and turbine rotors) [2–7]. Generally these research works have pointed out increments in the fatigue resistance at long life region of coated components with respect to uncoated ones, and such result was interpreted both by means of the constraining effect of slip deformation on the sample surface and by the high level of compressive residual stress affecting the coating [8–11]. However, the choice of the type of coating in terms of deposition technique, chemical composition and thickness is, at the moment, frequently referred to empiric consideration. Therefore, the settlement of previsional numerical models for the fatigue behaviour of coated parts in which the properties characterising substrate and coating materials are explicit would be considered of strategic importance. In fact, according to the applied loads and the type of substrate material, it would allow a simplified and optimised selection of the coating material. The choice of the substrate material is another critical step when the coated component has to guarantee not only high levels of mechanical properties but also a lightness of the final part. Considering, for instance, the automotive industry, a more widespread use of light metal alloys will be guaranteed in tribological applications like bearing plates,
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Nomenclature C a d dc d2 d* E H1 H2 H3 HV KIc Kres DK
material constant crack depth (mm) distance between the microstructural barriers (mm) thickness of coating (mm) depth of maximum hardness data (mm) depth from the surface (mm) Young’s modulus of the material (N/mm2) surface Vickers hardness maximum Vickers hardness core Vickers hardness Vickers hardness pffiffiffiffi fracture toughness (MPa m) stress intensity factor due to residual stresses stress intensity factor range
guide bars or seat supports in the presence of functional surface coatings able to provide good wear protection and stiffness. Nowadays several surface treatment processes have been optimised for aluminium alloys (electroless nickel plating and anodic oxidation are two typical examples of coating industrially used for mechanical applications produced in aluminium alloys) in order to obtain a better behaviour when a high corrosion and wear resistance behaviour are requested. Up to now an unsolved feature characterizes such coatings: no one of them can match the requirement of bearing sliding movements under high applied loads, condition particularly difficult when the deposed coating shows a thin thickness. Therefore, there is a steady demand for tailored coating suitable to overcome such a problem and, in the meantime, to assure high levels of fatigue resistance. The low melting point, peculiar for aluminium alloys, is another factor that must be considered to select the proper coating depositing parameters; a low deposition temperature is, in fact, a solution avoiding severe modification in the microstructure of the substrate material and, consequently, a significant reduction in the mechanical behaviour of the final coated part. A further aspect that is not to be neglected is the strong tendency to oxidation of aluminium alloys requiring a specific and accurate cleaning of the part before the deposition step. A ion cleaning carried out by means of an appropriate gas during the first period of deposition with the aim of contrasting the Al2O3 formation, in order to assure a good adhesion coating/substrate, is also requested. Starting from such assumptions, in the present work, the fatigue behaviour of two steel grades with differences in terms of hardness data and used in different applications (a 2205 duplex stainless steel and a H11 tool steel) was compared to the one of the 6082 alloy. Tests were carried out on both uncoated and CrN-PVD coated samples having a thickness ranging from 2.5 to 5 mm. The experimental
DKapp
DKeff DKth N r Ra u
q n
applied stress intensity factor range (takes into account the reduction of DKeff due to the preload applied during the experimental tests) effective stress intensity factor range (due to the presence of the residual stress field) pffiffiffiffi threshold stress intensity factor range (MPa m) number of cycles distance from the crack tip surface roughness (mm) FEM half relative displacement between the flanks of the crack, for the couple of nodes on the opposite flanks of the crack at the crack tip (mm) angular position taken with reference to the crack tip (rad) Poisson’s ratio
results were interpreted by means of microstructural characterisation and residual stress gradient evaluation. A numerical finite element model was developed and enabled us to predict fatigue life of the coated specimens only in case of absence of delamination between coating and bulk material.
2. Fatigue life evaluation of thin coated components: numerical models It is well known that for components of engines or structures used in car or motorcycle races or in aerospace industry, the capability of predicting the number of cycles until failure of structural components is very important. Bearing in mind this last consideration, authors developed a model able to predict the number of cycles until failure of thin hard coated components. The numerical model, once the initial crack depth from the surface of the component is known, enables one to calculate the stress intensity factor (including residual stress effects) and to simulate crack propagation (in a discrete way) and to evaluate the number of cycles needed to reach specified crack depths until rupture of the component. The quarter-point technique has to be used in order to enhance the accuracy of results obtained from the numerical analyses [12]. As regards the nucleation phase of defects, in this paper only simulation of surface defects was carried out. Surface cracks may be generated by brittle fracture of coating or by the presence of surface droplets or defects at least 5 mm deep. Due to the presence of residual stresses, crack nucleation is always located at surface or subsurface defects, like droplets or inclusions [13]. Numerical evaluation of the stress intensity factor, in the presence of residual stresses, can be carried out by using the value of the half relative displacement, u, of crack
S. Baragetti et al. / International Journal of Fatigue 27 (2005) 1541–1550
surfaces: rffiffiffiffiffiffiffi E 2p u KI Z 1 Cn r f ðqÞ
(1)
q 2 q f ðqÞ Z sin k C 1 K 2cos 2 2
(2)
kZ
3 Kn ; 1 Cn
k Z 3 K 4n;
for plane strain state (3) for plane stress state;
Having established the stress intensity factor from numerical simulation of crack propagation (for each crack depth taken into consideration), it is possible to calculate the stress intensity factor range and to compare it with the threshold stress intensity factor range of the material obtainable from previously published models or formulas [14–17]. If the numerical stress intensity factor range is higher than the threshold stress intensity factor range of the material, fatigue crack propagation is possible. If the crack propagates, then the fatigue crack growth rate, and the number of cycles necessary to reach specified crack depths, can be evaluated by means of existing models. The model is suitable and utilisable in all cases where no delamination between coating and substrate material can be verified, for thin hard coating (with thickness ranging from 1 to 10 mm) to thick coatings (thickness usually over 10 mm). The coating can be of any kind (PVD, CVD, Thermal spray, etc.). The model can be developed by means of a commercial code [18] and mesh refinement has to be applied only at the cracked area. The dimensions of the finite elements at the cracked area can reach the minimum value 0.5–1 mm; care should be put in the phase of mesh refinement in order to reduce the computational time and diminish the total number of degrees of freedom of the model. Substrate material and coating mechanical properties have to be assigned; only Young’s modulus and Poisson’s ratio have to be attributed just because an elastic material behaviour (both for the coating and for the substrate material) was used in all of the analyses, having assumed a limited plasticized area at the crack tip (during the fatigue propagation inside the substrate material such approximation is well verified only in the case of the H11 tool steel grade while during the propagation inside the coating such hypothesis is always substantiated). The residual stress field, induced by coating deposition process has to be applied as an initial condition to the model (residual stresses act before external applied loads). Application of residual stresses can be carried out by means of a temperature gradient procedure: a specified temperature profile, with nodal temperature assignation from surface to in-depth layers of finite elements, enabled simulation of the presence of the residual stress profile. Due to the residual stress field induced by the presence of
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the coating, with rather high compressive residual stresses at the surface, contact elements have to be used in order to avoid overlapping of the mating surfaces of the crack (one flank of the crack is prevented from penetrating the other). Introduction of contact elements enables evaluation of stress intensity factors due to residual stresses present before load application. The stress intensity factor range has to be calculated by taking into consideration the presence of residual stresses and introducing the parameter DKeffZ KIKKres; reduction of DKeff due to the preload applied during the experimental tests has to be considered too, by introducing a numerical applied stress intensity factor range DKapp. The numerical procedure enables determination of the number of cycles until failure but does not allow calculation of the fatigue limit in the case of coated components. Determination of the number of cycles until failure, for crack depths ranging from 0.5 to 1 mm to specified depths, can be carried out by using literature models developed for carburised spur gears [19–21]. Such models require knowledge of the material hardness versus depth from the surface of the component, and the value of the effective stress intensity factor range (evaluated by means of the finite element analyses) da C Z ðDK n K DKthn Þ; dN ð1 K rn Þ
for DKth % DK % KC
da C DK n KIc Z ; for KC !DK!KIc ; dN ð1Krn Þ DKIcn KDK n
(4)
(5)
DK rZ th ; KC ZðDKth KIc Þ1=2 KIc Symbols in Eqs. (4) and (5) have the following meaning: 1 H1 KH3 AZK 2 ln ; d %d2 H2 KH3 d2 AZK
1 550KH3 ln ; d Od2 H2 KH3 ðdc Kd2 Þ2
H ZðH2 KH3 Þexp½KAðd Kd2 Þ2 CH3 DKth Z2:45C3:41!10K3 H KIc Z141K1:64!10K1 H nZ4:31K8:66!10K3 HC1:17!10K5 H2 logðCÞZK10:0C1:09!10K2 HK1:40!10K5 H2 2.1. Experimental data necessary for numerical fatigue life evaluation Mechanical elastic properties (Young’s modulus and Poisson’s ratio) both of bulk material and coating have to be
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introduced as characteristics of the finite elements belonging to each material. Along with material elastic properties, fatigue life evaluation by means of numerical models needs assessment of three important parameters: residual stress gradient, microhardness and surface roughness. The residual stress field is maybe the most important parameter just because of the high compressive surface residual stresses that enable to prevent crack nucleation and propagation [22]. The residual stress field should be known from the surface of the component to inner layers of material. Authors verified, by means of X-ray diffraction measurements, that, in the case of CrN-PVD coatings having thickness 5 mm, the surface film deposition induces a high compressive surface residual stresses (fairly constant in the coating and assuming values ranging from 0.8 to 3 GPa in function of the substrate material [11,22–26]). Structural characterization was executed by means of XRD technique with Bragg–Brentano geometry. 2D images (XRD2) were also obtained by means of a D/max RAPID Rigaku microdiffractometer in order to measure the texture, the lattice constants and the residual stress values characterizing the coating. In particular for the residual stress analysis, the DRAST method was used [22]. Such method permits to measure the residual stresses active in the specimen from the evaluation of the displacement of the Debye ring; details regarding the procedure used during the tests carried out on CrN-PVD coatings can be detected in Ref. [23]. In particular to calculate the residual stress, Eq. [23] dfj K d0 1 Cy y sf sin2 j K ½s11 C s22 Z E E d0
the residual stress field is not much important as the surface value of residual stresses. At different crack depths, the numerical value of the effective stress intensity factor range depends only on the surface residual stress and the application of different experimental trends of subsurface residual stresses did not significantly affect the stress intensity factor range. Due to the steep residual stress gradient present at the interface between coating and bulk material, attention must be paid to the presence of irregularities and defects that could be source of fatigue crack nucleation and propagation. In this case, the proposed model could be suitable if the aim is to simulate the presence of subsurface cracks, especially at the interface between coating and bulk material, emanating from defects, inclusions and irregularities; anyway the crack propagation path has to be perpendicular to the surface of the component. One of the residual stress trend measured by means of X-ray diffraction by the authors is shown in Fig. 1; this is the case of a CrN-PVD coated H11 tool steel sample (four point bending test specimen dimensions 110!21! 4.4 mm) with coating thickness 5 mm. Microhardness has to be measured in order to apply literature models for fatigue crack growth evaluation of hardened surfaces [19–21]. One of the microhardness trend measured by means of a nanoindenter equipped with a Vickers indenter is shown in Fig. 2; this is the case of a microhardness trend for a CrN-PVD coated H11 tool steel sample with coating thickness 5 mm. Considering specimens in which the crack nucleation starts from the coating, surface roughness should be measured too in order to avoid introduction of opportune correction coefficients that would enable to take into consideration the effects of surface roughness. In particular, it is well known that if surface roughness exceeds RaZ 0.3 mm (ISO 1143 Standard), machine design formulas used for fatigue strength analysis of a component has to be
(6)
for biaxial stress analysis, was used introducing the elastic modulus calculated from nanoindentation measurements (CrN coating: EZ303 GPa; Poisson ratio nZ0.2). Considering the high residual stress level characterizing a sizeable part of the coating, the knowledge of the trend of
600
Coating (CrN PVD, thickness 5 µm)
300
10.500
10.490
10.480
−300 −600 −900
Bulk material (H11 tool steel)
−1200 −1500 −1800 −2100 −2400 −2700
Fig. 1. Residual stress trend for a PVD-coated H11 tool steel sample with coating thickness 5 mm.
Residual stresses [MPa]
Distance from the axis of the specimen [mm]
10.470
10.460
10.450
10.440
10.430
10.420
10.410
10.400
10.390
10.380
0
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1545
3000
X 2700
Coating (CrN PVD, thickness 5 µm)
2100
X Experimental data points
X
1800 1500
Bulk material (H11 tool steel)
X
1200
Hardness [HV]
2400
900
X 10.440
10.450
10.460
10.470
10.480
X
10.490
600
X
300 0 10.500
Distance from the axis of the specimen [mm] Fig. 2. Microhardness trend for a PVD-coated H11 tool steel sample with coating thickness 5 mm.
corrected by using a surface coefficient [27]. Anyway such problem does not affect the results of the proposed model just because, in the case of surface PVD coatings, fundamental is an accurate polishing before the deposition step (Ra!0.2 mm) in order to assure a high adhesion substrate/coating. For the specimens used in the present work, a Ra%0.15 was used. Surface roughness is always less than 0.15 mm (Ra%0.15 mm) for the PVD coated components studied in this paper and, in general, surface roughness has to be as low as possible in order to have good adhesion of the coating with the substrate. If the aim is to evaluate fatigue resistance in presence of surface defects (droplets of cracks) by means of numerical FEM models, that enable one to simulate the presence of such defects having minimum depth 5 mm (much higher than roughness ‘cracks’) than the propagating crack will be the one having depth 5 mm. After brittle rupture of coating (or in presence of defects at the surface of the components), defects having minimum depth 5 mm appear at the surface and surface roughness is no more the driving parameter for fatigue life evaluation. 2.2. Numerical models results: application to four point bending tests Four point bending tests were simulated and fatigue crack propagation from a surface crack having depth equal to the coating thickness was numerically calculated. With reference to H11 tool steed samples, some images related to the finite element model developed are shown in Fig. 3. The finite element models are composed of plane stress four nodes linear finite elements and, due to the symmetry in load and geometry, only one half of the specimens was modelled and symmetry boundary conditions were applied at the symmetry surface.
The values of DKth and DKapp, at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa), are shown in Table 1 [24]. Analysis of Table 1 puts in evidence how, in case of 2205 stainless steel, the stress intensity factor range is higher than the threshold stress intensity factor range only for crack depth over 5 mm. In case of H11 tool steel, the stress intensity factor range is higher than the threshold stress intensity factor range only for crack depth over 55 mm. It is worth underlining that if surface irregularities or inclusions are present in the bulk material, these defect may dramatically affect fatigue life of the component. The number of cycles until failure, calculated at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa), are shown in Table 2. The numerical model enables one to calculate the number of cycles until failure, summing the number of cycles necessary to reach specified crack depths (Table 2). Analysis of Table 2, in case of 2205 stainless steel, shows as the specimen spends almost all of its life in the phase of crack propagation from 5 to 25 mm. This is principally due to the high residual stress field, induced by the coating deposition technique, that reaches high values in the coating (the trend or residual stresses for 2205 stainless steel is similar to the one shown in Fig. 1 for H11 tool steel and in the coating the maximum residual stress measured was K1500 MPa [24]). If fatigue life is set at 2,000,000 cycles, the model enables to say that for 2205 coated specimens, the fatigue limit might be set over 600 MPa (the model predicts a fatigue life equal to 2,930,897 cycles). In case of H11 tool steel, with maximum applied stress 845 MPa, propagation
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Fig. 3. Plane stress finite element models for DK evaluation of coated specimens at different crack depths (developed for H11 tool steel specimens).
of a surface crack would be admissible only if a surface defect would be present and would have a minimum depth from the surface equal to 55 mm. In this case, the predicted number of cycles until failure would be equal to 37,670. Otherwise there is no fatigue crack propagation and the maximum applied stress 845 MPa is the minimum value of the fatigue limit of the coated component.
2.3. Coated components fatigue tests and numerical models validation Comparison between the results obtained from fatigue tests, carried out on steel specimens (H11 tool steel and 2205 duplex stainless steel) was arranged both with the aim of evaluating fatigue resistance of 2205 steel, H11 steel and
Table 1 DKth and DKapp at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa) 2205 Stainless steel 600 MPa
H11 Tool steel 845 MPa
Crack depth (mm)
pffiffiffiffi DKth (MPa m)
pffiffiffiffi DKapp (MPa m)
Crack depth (mm)
pffiffiffiffi DKth (MPa m)
pffiffiffiffi DKapp (MPa m)
5 25 50 100 500 1000
n.d. 3.00 3.06 3.10 3.30 3.30
0.00 3.05 8.54 11.75 27.96 35.71
5 30 55 105 505 1000
n.d. 4.23 5.18 3.39 3.73 3.73
0.00 2.22 2.90 8.19 22.40 43.80
S. Baragetti et al. / International Journal of Fatigue 27 (2005) 1541–1550
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Table 2 Number of cycles until failure, calculated at different crack depths, respectively, for 2205 stainless steel (applied maximum bending stress 600 MPa) and H11 tool steel (applied maximum bending stress 845 MPa) 2205 Stainless steel, 600 MPa
H11 Tool steel, 845 MPa
Crack depth (mm)
Number of cycles
Crack depth (mm)
Number of cycles
5–25 25–50 50–100 100–500 500–1000 1000-Rupture Fatigue life
2,911,346 8164 6030 2812 2436 109 2,930,897
5–30 30–55 55–105 105–505 505–1000 1000-Rupture Fatigue life
No propagation No propagation 31,360 5520 690 100 37,670
Table 3 Chemical composition (wt%) of the materials used as substrate Material
C
Mn
Si
Cr
Ni
Mo
Mg
Al
Fe
H11 2205 6082
0.43 0.019 –
1.50 1.51 0.73
0.30 0.39 0.99
1.99 22.45 0.05
– 5.5 –
0.21 3.12 –
– – 0.78
0.031 – Balanced
Balanced Balanced 0.46
6082 aluminium alloy coated specimens and with the aim of validating the proposed analytical–numerical procedure for number of cycles to failure calculation. The chemical composition of the three alloys used as substrate materials is shown in Table 3, while the heat treatments carried out before the coating deposition and the measured hardness and tensile values are summarized in Table 4. Independently from the substrate material a monolayer of CrN was deposited using an industrial coating plant equipped with a cathodic arc apparatus. Considering the difference in the bulk material, the standard deposition temperature of 573 K, used for the production of PVD films on steel substrate, was reduced to 483 K during the deposition on 6082 alloy. To oppose the overheating of the aluminium alloy inducing an overaging of the substrate, a reduction in the film thickness, and therefore of the time requested for the deposition step, was imposed: 2.5 mm instead of the 5 mm used for the deposition on the two types of steel investigated in the present work. In the case of the tested steel grades, microhardness data resulted practically independent from the substrate material (2686 HV for H11 steel and 2567 HV for 2205 duplex stainless steel) while a certain decrease in the coating hardness was obtained testing coated 6082 samples (HVZ 2238); such variation in the coating hardness is probably due to trace back to the load applied to the indenter that, at a penetration depth from the sample coated surface equal to 0.5 mm, resulted partially influenced by the hardness of the substrate material [28]. Four point bending fatigue tests were carried out both for 2205 stainless steel and H11 tool steel while for aluminium rotating bending tests were arranged. Only four point bending tests needed the development of a device that,
mounted on the universal test machine, could enable symmetric application of loads on the specimen (Fig. 4). The device consists of a pin connection (with a roller bearing) mounted between the specimen and the test machine (Fig. 4b). Only for H11 tool steel specimens, special supports (in PTFE) had to be developed in order to guarantee an appropriate guide to the specimen during the test (Fig. 4a). Prismatic fatigue samples were produced for testing both 2205 duplex stainless steel (110!11!5.5 mm) and H11 tool steel (110!21!4.4 mm). For 6082 samples, a hourglass geometry was used (minimum diameter 10 mm and length 110 mm according to ISO 1143 standard). Stair case was applied to the results of the fatigue tests carried out on steel and aluminium samples [29]. Experimental fatigue limit, both for coated and uncoated 2205, H11 and aluminium alloy, is reported in Table 5. Numerical results, listed in Tables 1 and 2, show how there is no fatigue crack growth if the crack has a depth less than or equal to 105 mm for H11 tool steel and 25 mm for duplex stainless steel. The number of cycles until failure evaluated by means of the numerical model is 37,670 for H11 tool steel with maximum applied bending stress 845 MPa, and can be calculated only if a surface crack, or a subsurface non metallic inclusion or matrix irregularity, having depth at least 55–105 mm is present at the surface of the specimen. On the other hand, Table 4 Mechanical properties characterising the substrate materials before coating Material
Condition
Hardness
YS (MPa)
UTS (MPa)
H11
QuenchingC tempering Solution annealing T6
32 HRC
893
1006
225 HB 113 HB
513 349
790 365
2205 6082
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Fig. 4. Device developed for four points bending fatigue tests.
simulation of fatigue crack propagation in 2205 stainless steel gives a fatigue life of 2,930,897 at 600 MPa maximum bending applied stress. Experimental four points bending tests confirmed the numerical results. For H11 tool steel specimens only for specimens presenting sub-surface cracks, or inclusions like the ones shown in Fig. 5, rupture occurred before the established fatigue life limit set at 2,000,000 cycles. If the surface of the specimen presents irregularities, surface cracks or subsurface defects with depth from the surface less than 55 mm, no fatigue crack nucleation or propagation for applied stresses less than 845 MPa (the experimental fatigue limit of H11 obtained from application of stair case method) is expected. The proposed model is confirmed by the experimental stair case fatigue tests results: the specimens with initial surface crack depth less than 55–105 mm either presented a number of cycles until failure over 2,000,000 cycles, or did not present any failure (stopped tests): such specimens were considered as unbroken ones during evaluation of the fatigue limit (stair case method). As concerns the 2205 duplex stainless steel the experimental number of cycles until failure is always over 2,000,000 cycles. The experimental results confirm the number of cycles evaluated by means of the numerical approach. The number of cycles, both numerical and experimental, is higher than 2,000,000, thus confirming that the model makes it possible to predict the resistance of specimens beyond the limit number of cycles set at 2,000,000. Moreover, the experimental number of cycles
until failure for the same specimen, in the same conditions as those of the numerical model, is 3,380,500. The difference between the numerical (2,930,897 cycles) and the experimental (3,380,500 cycles) number of cycles until failure is less than 14% and allows a certain confidence in the applicability of the numerical procedure proposed [24]. Coated specimens exhibit a fatigue life increment of about 15%, both in case of stainless steel and H11 tool steel. A slight decrement in the fatigue life was observed in the case of aluminium rotating bending fatigue tests notwithstanding the high level of compressive residual stress measured at the coating (data always higher than K0.8 GPa). To interpret such a difference, the mechanism of rupture induced by the cycling load has to be examined. The nucleation of the fatigue cracks for H11 steel grade, showing quite large non-metallic inclusions, resulted promoted only in a zone sufficiently away from the interface. In such area, the effect of the compressive residual stress induced from the coating procedure can be neglected. On the other hand, 2205 duplex stainless steel grade, being characterised by a higher level of cleanness, shows a nucleation located at the coating, and, in order to
Table 5 Endurance limit values for coated and uncoated samples—experimental results obtained testing the different substrate materials Bulk material
Fatigue limit uncoated samples (MPa)
Fatigue limit coated samples (MPa)
H11 2205 6082
730G26 510G18 165G11
845G43 585G26 142G10
Fig. 5. H11 coated sample—example of nucleation from a large sub-surface non metallic inclusion.
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Fig. 6. (a) Fractured surface of 2205 four-points bending fatigue test coated sample loaded at 600 MPa, and (b) 6082 rotary bending fatigue test coated specimen loaded at 140 MPa. Arrows indicate fatigue nucleation sites.
oppose the effect of the compressive residual stresses, the application of cycling loads more than the yield strength of the substrate is needed. The weakening of the coating increased for the 6082 coated samples, where the surface hard film cracks in multiple points due to the impossibility to sustain the imposed strain as requested from the overaged aluminium alloy substrate. In Fig. 6, a comparison between two fatigue fracture surfaces typical of a coated 2205 sample and a coated 6082 specimen can be observed. Independently from the substrate material, a good adhesion steel/coating was always detected. The droplets, typical defects of PVD coatings, cannot be taken as the area of fatigue nucleation independently from the substrate material.
3. Conclusions Different CrN-PVD coated specimens, made of steel and aluminium alloys, were tested in this paper and fatigue resistance was evaluated. Experimental investigations were also used in the development of a numerical procedure able to predict the number of cycles until failure of coated components. An enhancement in the fatigue limit of steel coated specimens was obtained with respect to uncoated ones. The measured increment is about 15%. For aluminium-coated specimens, no enhancement in the fatigue limit was detected with respect to uncoated ones. However, an enhancement in the fatigue limit might be obtained by opportunely modifying the deposition parameters in order to limit overaging of the substrate material. Such thesis should be verified carrying out further tests. The numerical procedure, based on finite element modeling, enables one to calculate the number of cycles until failure of coated components by simulating, in a discrete way, crack propagation. The trustworthiness of such procedure was verified through the execution of experimental fatigue tests.
Acknowledgements The authors thank S. Troglio of Vacuum Surtec for the preparation of the coated samples and the Chemical
Structural Laboratory of the Brescia University for the X-ray measurements.
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