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Which hardness (nano or macrohardness) should be evaluated in cavitation?
ARTICLE IN PRESS Tribology International 42 (2009) 1021–1028
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Which hardness (nano or macrohardness) should be evaluated in cavitation? C. Godoy a,, R.D. Mancosu a, R.R. Machado b, P.J. Modenesi a, J.C. Avelar-Batista c a Department of Metallurgical and Materials Engineering, School of Engineering, Universidade Federal de Minas Gerais, Rua Espı´rito Santo 35, Belo Horizonte-MG – 30160-030, Brazil b National Institute of Metrology-Brazil, Standardization and Industrial Quality, Inmetro, Brazil c Tecvac Ltd., Buckingway Business Park, Swavesey, Cambridge CB4 5UG, UK
a r t i c l e in fo
abstract
Article history: Received 25 November 2006 Received in revised form 22 August 2008 Accepted 2 September 2008 Available online 10 October 2008
This paper is concerned with the analysis of data obtained from instrumented depth-sensing nanoindentation testing performed on materials whose hardness varies with depth from surface. The objective is to determine which hardness (nano or macrohardness) should be evaluated in functionally graded materials to correlate the material properties to its performance in cavitation. A linear relationship between H2/E and cavitation erosion resistance could be established. Significant improvements in the cavitation erosion resistance can be achieved when deep nitrided cases are present. & 2008 Published by Elsevier Ltd.
Keywords: Instrumented indentation testing Cavitation erosion resistance Stylus-based profilometer
1. Introduction Cavitation erosion is often defined as the rapid formation and collapse of vapor bubbles in liquids as a result of strong pressure fluctuations. As bubbles generated in the liquid implode on a solid surface, stresses of several gigapascals with the duration of a few microseconds are induced, resulting in the formation of cavities and therefore material wear. Water and sewage pumps, propellers and impellers are typical examples of parts that experience cavitation wear. It is well known that cavitation erosion rates are often influenced by yield properties (hardness and rate of strainhardening), elastic properties (elastic modulus, resilience, superelasticity) and surface topography. Materials exhibiting compressive residual stresses and surface texture characterized by a low number of nucleation sites for the implosion of bubbles can also enhance the cavitation erosion resistance. Therefore, surface treatments such as plasma nitriding and coatings produced by plasma-assisted physical vapor deposition (PAPVD), which promote a significant increase in surface hardness, have potential to improve cavitation wear. PAPVD coatings can also be tailored to exhibit high hardness and low elastic modulus, which translates into a higher elastic strain to failure (i.e., higher resilience). Moreover, compressive residual stresses that often result from both treatments are expected to play a positive role in enhancing the cavitation erosion resistance. Bearing in mind all benefits
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described above, it is envisaged that PAPVD and plasma nitriding treatments, either isolated or as a combination, could be successfully used in sewage systems and water pumps to combat cavitation erosion. The hardness and elastic modulus of a ceramic coating are critical factors determining the performance of the coated product. Indeed many coatings are specifically developed to provide wear resistance which is usually conferred by their elevated hardness. Measurement of coating hardness is often used as a quality control check. ISO 14577-4:2002 International Standard describes a method of measuring hardness and modulus of elasticity of coatings by means of instrumented indentation testing (IIT), using instruments capable of measuring force and displacement as a function of time during the indentation process [1]. It is relatively uncomplicated to determine the hardness of bulk materials using instrumented indentation. However, since these measurements are made normal to a coated surface, depending on the force applied and the thickness of the coating, the substrate properties influence the result. In this case, the mechanical properties are depth-dependent and hardness becomes a function of the applied force. Under very low applied force conditions, the indentation depth is small and hardness will be evaluated in a micro- or nano-scale reflecting, predominantly, the coating hardness. In contrast to this, if we use intermediate applied force conditions, we will have more influence of the substrate and the measured hardness relates to the superficial substrate hardness. If superficial modifications such as plasma nitriding were carried out, we obtain hardness values dependent on depth in all nitriding case depths. Nevertheless, it is not clear which hardness (nano, micro or macrohardness) should be
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evaluated in a particular mechanical application, to successfully correlate the material properties to its performance. Conversely the cavitation erosion kinetics is influenced by the mechanical properties and surface conditions of the material. Our previous results in cavitation erosion studies [2] indicated that the cavitation erosion resistance of steels can be further improved by two treatments: plasma nitriding and PAPVD coating. Although the plasma nitriding treatment increased the surface roughness, the cavitation erosion resistance was not adversely affected. Moreover, the results clearly showed that the deposition of thin PAPVD coatings provided only a limited improvement in the cavitation performance of the steel substrate. Plasma nitriding was more effective in further enhancing the cavitation erosion resistance, demonstrating that not only hard but sufficiently thick (i.e., deep) layers are required to reduce wear produced by the cavitation erosion process. There is however little information available to the engineer regarding the depth required for specific purposes [3]. In order to understand the variables involved in cavitation erosion process, in this study we have investigated the relationship between depth of cavitation damage and hardness at various depths. 3D profilometry was utilized to quantify the cavitation craters and evaluate the depth of cavitation damage. A contact profilometer, using a lightly-loaded stylus with appropriate geometry, was employed. Hardness measurements were performed using a Nanoindenter XP (MTS Systems Corporation, Oak Ridge, TN, USA). Forces applied varied from 15 to 8000 mN. Four systems were investigated, as follows: (i) AISI 1045 steel; (ii) plasma nitrided AISI 1045 steel; (iii) PVD Cr1xNx coating on AISI 1045 steel (non-duplex) and (iv) PVD Cr1xNx coating on plasma nitrided AISI 1045 steel (duplex). Relations between measured hardness and depth of cavitation were established for these systems. This procedure allowed for greater comprehension concerning the erosion cavitation process.
2. Experimental procedure Polished, annealed AISI 1045 steel was used as the substrate for duplex and non-duplex Cr1xNx systems. A Tecvac IP70 L PAPVD reactor was utilized to deposit 3 mm thick Cr1xNx coatings by electron beam evaporation. For the duplex system, the plasma nitriding treatment and Cr1xNx coating were implemented as successive operations in the PAPVD reactor. The AISI 1045 was plasma nitrided at 430–450 1C for 2 h using a triode glow discharge possessing 60%Ar+40%N2 gas composition. A nonduplex Cr1xNx system (i.e., without plasma nitriding) and plasma nitrided steel (i.e., without the Cr1xNx coating) were also produced for comparison. For steels, duplex and non-duplex Cr1xNx systems, the indentation hardness (HIT) and indentation elastic modulus (EIT) were determined by nanoindentation at the NanoMechanical Properties Laboratory, Ceramic Division, National Institute of Standards and Technology (NIST/USA). Determination of these parameters was made according to the ISO 14577-1: 2002 Standard [1]. Measurements were applied using the Nanoindenter XP equipment (MTS Systems Corporation, Oak Ridge, TN, USA) and a Berkovich indenter. The average indentation hardness and indentation elastic modulus were obtained from between 20 and 25 measurements under the loads from 15 to 8000 mN. Indentation hardness was obtained by HIT ¼
F max Ap ðhc Þ
(1)
where Fmax is the maximum applied force and, Ap(hc) is the projected (cross-sectional) area of contact between the indenter
and the test piece determined from the force–displacement curve and a knowledge of the area function of the indenter. This definition of hardness is based on the contact area under load. A first approximation to the projected area, Ap(hc), is given by the theoretical shape of the indenter. For a perfect Vickers indenter, Ap ¼ 24.5h2c , where hc is the depth of the contact of the indenter with the test piece at Fmax given by hc ¼ ht0.75(FmaxCs) where ht is the total penetration depth into the specimen at Fmax and Cs is the compliance of the contact. And, for a Berkovich indenter, elastic modulus was obtained using the expressions [1]: EIT ¼
1 ðns Þ2 1 1 ðui Þ2 Ei Er
,
(2)
where ns, ni are the Poisson’s ratios of the test piece and of the indenter, respectively; Er is the reduced modulus of the indentation contact; Ei, is the modulus of the indenter; pffiffiffiffi p pffiffiffiffiffiffi (3) Er ¼ 2C s Ap where Cs is the compliance of the contact, i.e., dh/dF of the test force removal curve evaluated at maximum test force (reciprocal of the contact stiffness) and Ap ¼ 24.5h2c , where hc is the depth of the contact of the indenter with the test piece at Fmax. A FUTURE TECH FM-1 microhardness tester was used to evaluate the Knoop hardness of cross-sectioned nitrided surfaces. Knoop hardness measurements, as a function of the depth from the cross-section surface, were performed under an applied force of 0.49 N. In this way, the hardness of the nitrided layer as a function of depth as well as its thickness could be determined. The vibratory cavitation erosion (VCE) resistance of steels and coated systems was evaluated by tests performed according to the ASTM G-32-03 standard [4]. The test specimen size was established by this standard. They were ground and polished to a surface finish of 1 mm before coating deposition. During the test, specimens were attached to a titanium sonotrode, immersed in distilled water and subjected to a vibratory frequency of 20 kHz and amplitude of 45 mm. The test temperature was established at (2075) 1C and the sample weight was periodically assessed to estimate the mass loss. The test was interrupted every 5 or 10 min; in these situations the specimens were removed from the sonotrode, dried and weighed on an analytical scale possessing the readability of 0.1 mg. For each stopping time of the test, quantitative 3D surface measurements were completed using a Hommelwerke T4000 stylus-based profilometer. A TKU 300 pickup (stylus tip radius ¼ 5 mm; cone angle ¼ 901) was employed and the size of the sampling area equalled 8 mm 8 mm. The sampling interval and scanning speed were established at 163 mm and 0.50 mm s1, respectively. A Gaussian filter with a cut-off length of 0.8 mm was utilized in all measurements. Amplitude parameters obtained from 3D roughness profile were used to describe the surface texture [5]. The 3D profilometry was also utilized to quantify the cavitation craters and evaluate the depth of cavitation damage. The 3D surface topography was manipulated and inverted upside down so that the cavitation craters appeared as summits. Histograms containing the number of peaks per mm2 versus amplitude were obtained for each cavitation erosion time. Cumulative mass loss plots were obtained as a function of cavitation time and, from these plots, the mass erosion rate was then estimated. Scanning electron microscopy (SEM) was also employed to monitor the evolution of the cavitation process. In order to estimate the cavitation erosion rates, statistical piecewise linear regression analyses were used [6], and different stages were
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found to occur for all tested samples during the cavitation erosion process. Piecewise linear regression models are useful tools that allow the experimental data to be split into linear segments (splines). Such statistical analyses were employed to predict the knot points associated with changes in the erosion stage, based on the experimental data obtained. Statistica software for Windows was used to perform the statistical analyses [7]. Results obtained from statistical analyses were compared to those obtained from the 3D surface topography analyses and in addition to those obtained from nanoindentation tests.
3. Results and discussion The experimental data obtained from VCE tests and fitted statistical models (piecewise linear regressions) are shown in Fig. 1. Table 1 summarizes the quantitative measures obtained with this statistical procedure. The duration of each stage as the erosion rates for each stage are described. For the AISI 1045 steel (Fig. 1a), only two erosion stages occurred: the incubation stage (first period) and an accelerated stage (second period). Three erosion periods were found to occur in non-duplex system (Cr1xNx coating onto the AISI 1045 steel): the incubation period, coating erosion period and substrate erosion period. For cavitation erosion times in excess of 6.0 h, the cavitation erosion rate (8.4 mg h1) was similar to that of the accelerated stage of the AISI 1045 steel (Fig. 1c). A lower erosion rate was recorded for the intermediate stage, confirming that the deposition of Cr1xNx coating improved the cavitation erosion resistance of the AISI 1045 steel. Statistical analyses also showed the existence of three erosion stages on the plasma nitrided steel (Fig. 1b). The first stage corresponded to the incubation period and the second erosion stage was associated with the nitrided layer (nitriding layer erosion period), which promoted a reduction in the cavitation erosion rate (Table 1). During the third stage (accelerated stage), the erosion rate determined for this period (5.2 mg h1) was lower
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than that of the AISI 1045 steel (8.4 mg h1). This result was interpreted in the previous paper [2] as possibly occurring because the cavitation erosion was still within the nitrided layer (up to cavitation times of 20 h). For the duplex Cr1xNx system, the incubation period was divided into two stages. The total incubation period (two stages) lasted for 2.14 h in the duplex system (Fig. 1d). The third stage was associated with the cavitation erosion exhibited by the Cr1xNx coating (coating cavitation erosion period). A lower cavitation erosion rate was found in the ‘‘nitriding layer erosion period,’’ compared to the same period in the non duplex system, indicating that the plasma nitriding treatment prior to PAPVD coating was effective in reducing the cavitation erosion rate of the Cr1xNx film. Similarly to the plasma nitrided steel, the duplex system also displayed a ‘‘nitriding layer erosion period.’’ The cavitation erosion rate determined for this stage (2.2 mg h1) was similar to that obtained for the plasma nitrided steel (Table 1), but this period extended up to 12.18 h for the duplex Cr1xNx system. An accelerated erosion stage (fourth period) also occurred in the duplex coating, being characterized by a lower erosion rate (3.9 mg h1) than the accelerated erosion period found in the AISI 1045 steel and non-duplex coating (8.470.1 mg h1). Fig. 2 shows hardness and final indentation depth (hf) versus force plots. According to the ISO 14577-4 International Standard, the hardness of hard coating can only be determined with a sharp (small tip radius) indenter that causes yielding within the coating. In general, this will only occur when hc/tco0.5 (ratio of maximum contact depth to coating thickness). For a spherical indenter the relation between the contact depth variable (hc) with final indentation depth (hf) is equal to [1]: hc ¼
hmax þ hf 2
(4)
We can conclude that hc4hf. Coating thickness was evaluated by SEM analyses. For the non-duplex Cr–N system, a thickness equal to 3 mm was determined; for the duplex system, a 5 mm value was estimated. Considering hf values obtained in our tests and the relation hc/tco0.5, for determining final coating hardness,
Fig. 1. Plots showing experimental data and fitted continuous piecewise linear regression models demonstrating the different wear periods: (a) AISI 1045 steel, (b) plasma nitrided AISI 1045 steel (c) non-duplex Cr1xNx system and (d) duplex Cr1xNx system.
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Table 1 Cavitation erosion rates and 95% confidence intervals obtained for each wear regime from statistical analyses (piecewise linear regressions). Specimen
AISI 1045 steel Plasma nitrided AISI 1045 steel Non-duplex CrxN1x Duplex CrxN1x
Erosion rate (mg h1) Incubation period (first regime)
Second regime
Third regime
Fourth regime
2.8270.06 (0.00–1.23 h) 0.1270.05 (0.00–3.40 h) 0.4470.04 (0.00–3.56 h) 1.870.1 (0.45–2.14 h)
8.470.1 (1.23–8.00 h) 2.070.3 (3.40–6.40 h) 4.470.2 (3.56–6.05 h) 0.970.2 (2.14–9.24 h)
5.270.6 (6.40–20.0 h) 8.470.3 (6.05–15.50 h) 2.270.2 (9.24–12.18 h)
3.970.3 (12.18–20.50 h)
Nonexistent regimes.
Table 2 Amplitude parameters for 3D surface roughness. Specimen
Sa (nm)
CI (nm)a
Sq (nm)
CI (nm)a
AISI 1045 steel Plasma nitrided AISI 1045 steel Non-duplex CrxN1x system Duplex CrxN1x system
64.6 119.0 69.0 111.5
61.9–67.2 112.6–125.4 63.5–74.5 107.5–115.5
95.0 158.5 102.0 158.5
91.8–98.2 152.9–164.1 98.5–106.5 156.1–160.9
a
C.I. 95%
Fig. 2. (a) Hardness and (b) final penetration depth data from nanoindentation measurements carried out as a function of applied force.
the indentation depth must be lower than 1.5 mm (1500 nm) for the non-duplex system and 2.5 mm (2500 nm) for the duplex system. Moreover, the recommended Ra value shall be 5% of the maximum indentation depth, i.e., hmax420 Ra [1]. In our 3D studies, the parameters Sa (roughness average, i.e., arithmetic mean deviation of the surface from the mean plane) and Sq (rootmean-square deviation of the surface, RMS) were used for
Fig. 3. Plots showing transversal Knoop hardness in (a) plasma nitrided AISI 1045 steel and (b) duplex system. Case depth was estimated about 200 mm.
characterising the amplitude property of surfaces (Table 2). AISI 1045 steel and the non-duplex system showed lower roughness than plasma nitrided AISI 1045 steel and the duplex system.
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Using instrumented indentation tests, the two conditions imposed for evaluation coating hardness, without any significant influence of the substrate, resulted in: for Cr–N duplex system, hc/tco0.5 (hmaxo2500 nm) and hmax420 Sa (hmax42180 nm) and for Cr–N non-duplex system, hc/tco0.5 (hmaxo1500 nm) and hmax420 Sa (hmax41380 nm). We can see in Fig. 2, for the duplex system these conditions are reached only for the force equal to 1000 mN. The Cr–N coating hardness was evaluated as (9.470.59) GPa and the elasticity modulus as (23679.34) GPa. For Cr–N non-duplex system, the force equal to 200 mN satisfied the two imposed conditions and gave a hardness of (1870.69) GPa and an elasticity modulus of (27877.20) GPa. However, considering that the coating hardness value as the value that does not depend upon the force applied in small depths, as shown in Fig. 2, Cr–N coating hardness can be determined as an average value for forces smaller than 200 mN
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for both systems. This procedure gives a hardness equal to (23.073.59) GPa in the duplex system and equal to (24.874.02) GPa for the non-duplex system. We can see an inconsistency between this procedure and those determined by the ISO 14577-4:2002 International Standard. The high roughness found in the duplex system led to that inconsistency. We concluded that a very smooth surface is fundamental to estimate the actual hardness by nanoindentation. According to Bobji et al. [8], as the roughness of a surface is increased, the hardness measured at depths comparable with the roughness scale deviates increasingly from the actual hardness. Hardness of all systems (coating+plasma nitrided layer+steel substrate) was obtained using a considerable force. Considering F ¼ 4000 mN, we have determined for AISI 1045 steel a hardness value of (3.0170.03) GPa; for plasma nitrided AISI 1045 steel, a hardness of (4.170.3) GPa; for Cr–N non-duplex, hardness of (4.0570.07) GPa and for the duplex system a hardness equal to (5.770.3) GPa. Surface modifications increased hardness for all systems; even when using high force (with significant influence of the substrate) we experienced modifications in the final hardness value. The duplex system showed the maximum hardness value for all forces applied. Knoop hardness measurements, as a function of the depth from the surface, were also performed under a force of 9.81 102 N. Fig. 3 shows the dependence of hardness of the transversal profile. The hardness profile of the nitrided layer as well as its thickness was determined for Nitrided AISI 1045 steel system and the duplex system. A hardening depth equal to 200 mm (2.0 105 nm) was estimated. The maximum indentation depth reached in instrumented indentation tests, when the hardness was measured, was in the order of 7.75 103 nm. The dependence between depth of cavitation damage and hardening depth was also investigated. The 3D surface topography was manipulated and inverted upside down so that the cavitation craters appeared as summits for evaluating the depth of cavitation damage (Fig. 4). Prior to invert the surface, a procedure called ‘‘truncation’’ was employed. Truncation can be understood as the removal of the material above a given plane that is parallel to the mean plane. It may be used to reveal features below certain levels that are otherwise invisible. The truncation level used in our analyses was equal to Sr1 (upper bearing surface) (Fig. 5). Sr1 is the fraction of the surface that consists of small peaks above the main plateau [5,9]. The objective was to remove discrepant peaks present in the original surface. After this procedure, histograms
Sk parameters, gaussian filter, 0.8 mm. Spk = 0.52 µm
Sk = 1.44 µm
0 20 Sr1 = 6.92% Sa1 = 18 µm2/mm Fig. 4. Inverted isometric plots displaying the damages produced by cavitation erosion in the final time studied in each system.
Svk = 4.04 µm
40
60
80 100% Sr2 = 76.5% Sa2 = 476 µm2/mm
Fig. 5. Plot of bearing area ratio illustrating the determination of Sr1 (upper bearing surface).
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Fig. 6. Height distribution histograms for each erosion period determined by the statistical piecewise linear regression analyses.
containing the number of peaks (craters) per mm2 versus amplitude were obtained for each final erosion time specific to each cavitation erosion period (Fig. 6). Fig. 4 shows the inverted isometric plots for the cavitation erosion time equal to approximately 15 h, excluding AISI 1045 steel. The left scale in this figure corresponds to the surface total amplitude. The systems where the plasma nitriding process was
carried out presented shallower craters. In the non-duplex system, the cavitation erosion was characterized by deeper craters. Once the damage reaches the substrate, no obstacle hinders its propagation. Fig. 6 shows the histograms containing the number of ‘‘peaks’’ (cavitation craters) per mm2 versus amplitude, for each stage determined by the statistical piecewise linear regression analyses,
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in the final period associated to the specific stage. The following aspects can be observed: (i) The main aspects of the systems before the cavitation erosion process are: the nitrided systems show the value of amplitude that occurs most frequently (mode, statistical parameter) between 0.40 and 0.45 mm. Similar values can be observed for Cr–N duplex system (0.40–0.48 mm). These amplitudes are higher than those identified for the Cr–N non-duplex system (0.33–0.36 mm), confirming the higher roughness for plasma nitriding systems. (ii) In the final erosion time of the ‘‘incubation period,’’ the highest amplitudes (deepest craters) are present in the nonduplex system. Erosion cavitation was more pronounced in it, and plasma nitriding process appears again as the most effective method to prevent the formation of deep craters. (iii) The deposition of a coating under a hardening case is a means to guarantee shallower craters: this is illustrated in the ‘‘coating erosion period.’’ In fact, SEM images confirmed a completely different aspect of the cavitation erosion mechanism; in the non-duplex system the erosion occurs by localized and profound craters, on the other hand, in duplex systems the mass loss occurs by the slow decohesion of the coating. The amplitude mode was between 13 and 14 mm for the nonduplex system and between 6 and 6.5 mm for the duplex system. (iv) The maximum amplitude presented in the histogram (the maximum value for the x-axis), corresponding to the ‘‘nitrided layer erosion period,’’ was 24.7 mm for nitrided steel system and equal to 40.3 mm for the duplex system. These values are significantly lower than that of the case depth determined, i.e., 200 mm. This result suggests that the cavitation erosion in this stage probably corresponds to the cavitation erosion of the compound layer. Fig. 7 indicates a value for this layer thickness equal to approximately 5 mm. (v) In the ‘‘substrate erosion period,’’ the maximum amplitude depends on the erosion time. The AISI 1045 steel was tested for only 8 h, because severe damage was already observed in this time. plasma nitrided AISI 1045 steel system presented a maximum crater amplitude equal to 54 mm, after a erosion time of 12.5 h. For the same erosion time, Cr–N non-duplex system presented the maximum amplitude equal to 155 mm. Thus we concluded that the Cr–N non-duplex system undergoes a much more accentuated change with the cavitation
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erosion time. On the other hand, Cr–N duplex system, after an erosion time of 15.5 h, presented maximum crater amplitude equal to 73.5 mm, whereas in 12 h, this amplitude was equal to 40.3 mm. The duplex system, similarly to the nitrided plasma steel system, experiences minor damages with the cavitation erosion process. (vi) Even in the erosion times equal to 20 h, no plasma nitriding system presents damage with extension greater than the case depth, estimated equal to 200 mm. For these systems, the maximum amplitude of craters was still restricted within the case depth. Indeed surface hardening by the nitriding process offers a more effective barrier to cavitation erosion. The junction coating and plasma nitriding was the most effective system, according to profilometric analyses. We can conclude that significant improvements in the cavitation erosion resistance can be achieved when thick nitrided cases are present. If it is true, nanohardness, evaluated when the indentation depth is small, does not correlate effectively with the performance in cavitation erosion. For verifying which hardness (nano, micro or macrohardness) should be evaluated in cavitation erosion, plots of H2/E, for two forces used in instrumented indentation tests, were constructed. The following ratio— indentation hardness to the square ðHIT Þ2 and indentation elasticity modulus ðEIT Þ—was employed due to the known relation between resilience and cavitation erosion. The modulus of resilience is defined as the strain energy per unit volume required to stress the material from zero stress to the yield stress given by [10] Ur ¼
S20 2E
(5)
where S0 is the yield stress and E is the modulus of elasticity. Considering the proportionality between indentation pressure (hardness) and yield stress we have decided to correlate cavitation rate and H2/E, instead of S2/E. Local determination of yield stress in different penetration depths is almost impossible to be measured and on the other hand, the nanoindentation instrument allows the measurement of the hardness in different penetrations using different forces, even in very small depths. In coated systems, it is a powerful tool for determining hardness for different penetration depths. EIT values, determined as function of different forces, are showed in Table 3. Fig. 8 was constructed from two cavitation erosion rates: the erosion rate determined for the ‘‘substrate erosion period’’ and those determined for the ‘‘incubation period’’ (Table 1). This first value of rates was related to hardness values which were measured with forces that caused the greater penetration. In our study the force applied was equal to 8000 mN for the depth equivalent to between 9000 and 10 000 nm. For the ‘‘incubation period’’ the force used was equal to 15 mN corresponding to between 185 and 400 nm of indentation depth.
Table 3 Indentation Modulus (EIT) and indentation hardness (HIT) for 8000 and 15 mN. Systems
force (mN)
EIT (GPa)
s
HIT (GPa)
s
Cr–N non-duplex
8000 15 8000 15 8000 15 4000 15
222 381 272 374 192 204 230 254
11 11 16 49 5 13 13 10
3.73 30.2 4.53 26.9 3.51 7.50 3.01 4.10
2 102 2 100 1 101 6 100 1 101 1 100 3 102 6 101
Cr–N duplex Nitrided AISI 1045 steel AISI 1045 steel Fig. 7. SEM photomicrograph of the duplex system indicating the presence of the CrxN1x coating, the compound layer and iron nitrides in plasma nitrided layer.
s, Standard deviation.
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or macrohardness should be evaluated rather than nanohardness, to successfully correlate the material properties to its performance in cavitation erosion.
Acknowledgments The authors are greatly indebted to CNPq-Conselho Nacional de Desenvolvimento Tecnolo´gico, Brazil, and to FAPEMIG-Fundaca˜o de Amparo a` Pesquisa de Minas Gerais, Brazil, for the financial support. They would also like to thank Mr. Renato Machado for nanoindentation measurements at the National Institute of Standards and Technology, NIST, USA. References [1] International Organization for Standardization, Geneva. ISO 14577/1-4 Metallic materials-Instrumented indentation test for hardness and materials parameters. Geneva, 2002. [2] Godoy C, Mancosu RD, Lima MM, Branda˜o D, Housden J, Avelar-Batista JC. Influence of plasma nitriding and PAPVD Cr1x Nx coating on the cavitation erosion resistance of an AISI 1045 steel. Surf Coat Technol 2006;200:5370–8. [3] Child HC. Surface hardening of steel. Engineering design guides. Oxford: Oxford University Press; 1980. [4] American Society for Testing and Materials Standards, ASTM G 32-03. Standard test method for cavitation erosion using vibratory apparatus, 2003. [5] Stout KJ, Blunt L. Three dimensional surface topography. 2nd ed. London: Penton Press; 1994. [6] Montgomery DC, Peck EA. Introduction to linear regression analysis. 2nd ed. New York: Wiley; 1992. [7] Statistica DX. Software for Windows, Version 7, StatSoft South America, 2005. [8] Bobji MS, Biswas SK. Estimation of hardness by nanoindentation of rough surfaces. J Mater Res 1998;13(11). [9] Stout KJ. Development of methods for the characterisation of roughness in three dimensions. London: Penton Press; 2000. [10] Dieter GE. Mechanical metallurgy. McGraw-Hill Kogakusha, Ltd.; 1976.
Fig. 8. Plots of cavitation erosion resistance and resilience for two forces used in nanoindentation tests: (a) 15 mN and (b) 8000 mN.
In Fig. 8, we can see a better relation between cavitation erosion resistance and resilience when hardness was calculated using greater penetrations (forces). This result suggests that micro
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Which hardness (nano or macrohardness) should be evaluated in cavitation? C. Godoy a,, R.D. Mancosu a, R.R. Machado b, P.J. Modenesi a, J.C. Avelar-Batista c a Department of Metallurgical and Materials Engineering, School of Engineering, Universidade Federal de Minas Gerais, Rua Espı´rito Santo 35, Belo Horizonte-MG – 30160-030, Brazil b National Institute of Metrology-Brazil, Standardization and Industrial Quality, Inmetro, Brazil c Tecvac Ltd., Buckingway Business Park, Swavesey, Cambridge CB4 5UG, UK
a r t i c l e in fo
abstract
Article history: Received 25 November 2006 Received in revised form 22 August 2008 Accepted 2 September 2008 Available online 10 October 2008
This paper is concerned with the analysis of data obtained from instrumented depth-sensing nanoindentation testing performed on materials whose hardness varies with depth from surface. The objective is to determine which hardness (nano or macrohardness) should be evaluated in functionally graded materials to correlate the material properties to its performance in cavitation. A linear relationship between H2/E and cavitation erosion resistance could be established. Significant improvements in the cavitation erosion resistance can be achieved when deep nitrided cases are present. & 2008 Published by Elsevier Ltd.
Keywords: Instrumented indentation testing Cavitation erosion resistance Stylus-based profilometer
1. Introduction Cavitation erosion is often defined as the rapid formation and collapse of vapor bubbles in liquids as a result of strong pressure fluctuations. As bubbles generated in the liquid implode on a solid surface, stresses of several gigapascals with the duration of a few microseconds are induced, resulting in the formation of cavities and therefore material wear. Water and sewage pumps, propellers and impellers are typical examples of parts that experience cavitation wear. It is well known that cavitation erosion rates are often influenced by yield properties (hardness and rate of strainhardening), elastic properties (elastic modulus, resilience, superelasticity) and surface topography. Materials exhibiting compressive residual stresses and surface texture characterized by a low number of nucleation sites for the implosion of bubbles can also enhance the cavitation erosion resistance. Therefore, surface treatments such as plasma nitriding and coatings produced by plasma-assisted physical vapor deposition (PAPVD), which promote a significant increase in surface hardness, have potential to improve cavitation wear. PAPVD coatings can also be tailored to exhibit high hardness and low elastic modulus, which translates into a higher elastic strain to failure (i.e., higher resilience). Moreover, compressive residual stresses that often result from both treatments are expected to play a positive role in enhancing the cavitation erosion resistance. Bearing in mind all benefits
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E-mail addresses: [email protected], [email protected] (C. Godoy). 0301-679X/$ - see front matter & 2008 Published by Elsevier Ltd. doi:10.1016/j.triboint.2008.09.007
described above, it is envisaged that PAPVD and plasma nitriding treatments, either isolated or as a combination, could be successfully used in sewage systems and water pumps to combat cavitation erosion. The hardness and elastic modulus of a ceramic coating are critical factors determining the performance of the coated product. Indeed many coatings are specifically developed to provide wear resistance which is usually conferred by their elevated hardness. Measurement of coating hardness is often used as a quality control check. ISO 14577-4:2002 International Standard describes a method of measuring hardness and modulus of elasticity of coatings by means of instrumented indentation testing (IIT), using instruments capable of measuring force and displacement as a function of time during the indentation process [1]. It is relatively uncomplicated to determine the hardness of bulk materials using instrumented indentation. However, since these measurements are made normal to a coated surface, depending on the force applied and the thickness of the coating, the substrate properties influence the result. In this case, the mechanical properties are depth-dependent and hardness becomes a function of the applied force. Under very low applied force conditions, the indentation depth is small and hardness will be evaluated in a micro- or nano-scale reflecting, predominantly, the coating hardness. In contrast to this, if we use intermediate applied force conditions, we will have more influence of the substrate and the measured hardness relates to the superficial substrate hardness. If superficial modifications such as plasma nitriding were carried out, we obtain hardness values dependent on depth in all nitriding case depths. Nevertheless, it is not clear which hardness (nano, micro or macrohardness) should be
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evaluated in a particular mechanical application, to successfully correlate the material properties to its performance. Conversely the cavitation erosion kinetics is influenced by the mechanical properties and surface conditions of the material. Our previous results in cavitation erosion studies [2] indicated that the cavitation erosion resistance of steels can be further improved by two treatments: plasma nitriding and PAPVD coating. Although the plasma nitriding treatment increased the surface roughness, the cavitation erosion resistance was not adversely affected. Moreover, the results clearly showed that the deposition of thin PAPVD coatings provided only a limited improvement in the cavitation performance of the steel substrate. Plasma nitriding was more effective in further enhancing the cavitation erosion resistance, demonstrating that not only hard but sufficiently thick (i.e., deep) layers are required to reduce wear produced by the cavitation erosion process. There is however little information available to the engineer regarding the depth required for specific purposes [3]. In order to understand the variables involved in cavitation erosion process, in this study we have investigated the relationship between depth of cavitation damage and hardness at various depths. 3D profilometry was utilized to quantify the cavitation craters and evaluate the depth of cavitation damage. A contact profilometer, using a lightly-loaded stylus with appropriate geometry, was employed. Hardness measurements were performed using a Nanoindenter XP (MTS Systems Corporation, Oak Ridge, TN, USA). Forces applied varied from 15 to 8000 mN. Four systems were investigated, as follows: (i) AISI 1045 steel; (ii) plasma nitrided AISI 1045 steel; (iii) PVD Cr1xNx coating on AISI 1045 steel (non-duplex) and (iv) PVD Cr1xNx coating on plasma nitrided AISI 1045 steel (duplex). Relations between measured hardness and depth of cavitation were established for these systems. This procedure allowed for greater comprehension concerning the erosion cavitation process.
2. Experimental procedure Polished, annealed AISI 1045 steel was used as the substrate for duplex and non-duplex Cr1xNx systems. A Tecvac IP70 L PAPVD reactor was utilized to deposit 3 mm thick Cr1xNx coatings by electron beam evaporation. For the duplex system, the plasma nitriding treatment and Cr1xNx coating were implemented as successive operations in the PAPVD reactor. The AISI 1045 was plasma nitrided at 430–450 1C for 2 h using a triode glow discharge possessing 60%Ar+40%N2 gas composition. A nonduplex Cr1xNx system (i.e., without plasma nitriding) and plasma nitrided steel (i.e., without the Cr1xNx coating) were also produced for comparison. For steels, duplex and non-duplex Cr1xNx systems, the indentation hardness (HIT) and indentation elastic modulus (EIT) were determined by nanoindentation at the NanoMechanical Properties Laboratory, Ceramic Division, National Institute of Standards and Technology (NIST/USA). Determination of these parameters was made according to the ISO 14577-1: 2002 Standard [1]. Measurements were applied using the Nanoindenter XP equipment (MTS Systems Corporation, Oak Ridge, TN, USA) and a Berkovich indenter. The average indentation hardness and indentation elastic modulus were obtained from between 20 and 25 measurements under the loads from 15 to 8000 mN. Indentation hardness was obtained by HIT ¼
F max Ap ðhc Þ
(1)
where Fmax is the maximum applied force and, Ap(hc) is the projected (cross-sectional) area of contact between the indenter
and the test piece determined from the force–displacement curve and a knowledge of the area function of the indenter. This definition of hardness is based on the contact area under load. A first approximation to the projected area, Ap(hc), is given by the theoretical shape of the indenter. For a perfect Vickers indenter, Ap ¼ 24.5h2c , where hc is the depth of the contact of the indenter with the test piece at Fmax given by hc ¼ ht0.75(FmaxCs) where ht is the total penetration depth into the specimen at Fmax and Cs is the compliance of the contact. And, for a Berkovich indenter, elastic modulus was obtained using the expressions [1]: EIT ¼
1 ðns Þ2 1 1 ðui Þ2 Ei Er
,
(2)
where ns, ni are the Poisson’s ratios of the test piece and of the indenter, respectively; Er is the reduced modulus of the indentation contact; Ei, is the modulus of the indenter; pffiffiffiffi p pffiffiffiffiffiffi (3) Er ¼ 2C s Ap where Cs is the compliance of the contact, i.e., dh/dF of the test force removal curve evaluated at maximum test force (reciprocal of the contact stiffness) and Ap ¼ 24.5h2c , where hc is the depth of the contact of the indenter with the test piece at Fmax. A FUTURE TECH FM-1 microhardness tester was used to evaluate the Knoop hardness of cross-sectioned nitrided surfaces. Knoop hardness measurements, as a function of the depth from the cross-section surface, were performed under an applied force of 0.49 N. In this way, the hardness of the nitrided layer as a function of depth as well as its thickness could be determined. The vibratory cavitation erosion (VCE) resistance of steels and coated systems was evaluated by tests performed according to the ASTM G-32-03 standard [4]. The test specimen size was established by this standard. They were ground and polished to a surface finish of 1 mm before coating deposition. During the test, specimens were attached to a titanium sonotrode, immersed in distilled water and subjected to a vibratory frequency of 20 kHz and amplitude of 45 mm. The test temperature was established at (2075) 1C and the sample weight was periodically assessed to estimate the mass loss. The test was interrupted every 5 or 10 min; in these situations the specimens were removed from the sonotrode, dried and weighed on an analytical scale possessing the readability of 0.1 mg. For each stopping time of the test, quantitative 3D surface measurements were completed using a Hommelwerke T4000 stylus-based profilometer. A TKU 300 pickup (stylus tip radius ¼ 5 mm; cone angle ¼ 901) was employed and the size of the sampling area equalled 8 mm 8 mm. The sampling interval and scanning speed were established at 163 mm and 0.50 mm s1, respectively. A Gaussian filter with a cut-off length of 0.8 mm was utilized in all measurements. Amplitude parameters obtained from 3D roughness profile were used to describe the surface texture [5]. The 3D profilometry was also utilized to quantify the cavitation craters and evaluate the depth of cavitation damage. The 3D surface topography was manipulated and inverted upside down so that the cavitation craters appeared as summits. Histograms containing the number of peaks per mm2 versus amplitude were obtained for each cavitation erosion time. Cumulative mass loss plots were obtained as a function of cavitation time and, from these plots, the mass erosion rate was then estimated. Scanning electron microscopy (SEM) was also employed to monitor the evolution of the cavitation process. In order to estimate the cavitation erosion rates, statistical piecewise linear regression analyses were used [6], and different stages were
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found to occur for all tested samples during the cavitation erosion process. Piecewise linear regression models are useful tools that allow the experimental data to be split into linear segments (splines). Such statistical analyses were employed to predict the knot points associated with changes in the erosion stage, based on the experimental data obtained. Statistica software for Windows was used to perform the statistical analyses [7]. Results obtained from statistical analyses were compared to those obtained from the 3D surface topography analyses and in addition to those obtained from nanoindentation tests.
3. Results and discussion The experimental data obtained from VCE tests and fitted statistical models (piecewise linear regressions) are shown in Fig. 1. Table 1 summarizes the quantitative measures obtained with this statistical procedure. The duration of each stage as the erosion rates for each stage are described. For the AISI 1045 steel (Fig. 1a), only two erosion stages occurred: the incubation stage (first period) and an accelerated stage (second period). Three erosion periods were found to occur in non-duplex system (Cr1xNx coating onto the AISI 1045 steel): the incubation period, coating erosion period and substrate erosion period. For cavitation erosion times in excess of 6.0 h, the cavitation erosion rate (8.4 mg h1) was similar to that of the accelerated stage of the AISI 1045 steel (Fig. 1c). A lower erosion rate was recorded for the intermediate stage, confirming that the deposition of Cr1xNx coating improved the cavitation erosion resistance of the AISI 1045 steel. Statistical analyses also showed the existence of three erosion stages on the plasma nitrided steel (Fig. 1b). The first stage corresponded to the incubation period and the second erosion stage was associated with the nitrided layer (nitriding layer erosion period), which promoted a reduction in the cavitation erosion rate (Table 1). During the third stage (accelerated stage), the erosion rate determined for this period (5.2 mg h1) was lower
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than that of the AISI 1045 steel (8.4 mg h1). This result was interpreted in the previous paper [2] as possibly occurring because the cavitation erosion was still within the nitrided layer (up to cavitation times of 20 h). For the duplex Cr1xNx system, the incubation period was divided into two stages. The total incubation period (two stages) lasted for 2.14 h in the duplex system (Fig. 1d). The third stage was associated with the cavitation erosion exhibited by the Cr1xNx coating (coating cavitation erosion period). A lower cavitation erosion rate was found in the ‘‘nitriding layer erosion period,’’ compared to the same period in the non duplex system, indicating that the plasma nitriding treatment prior to PAPVD coating was effective in reducing the cavitation erosion rate of the Cr1xNx film. Similarly to the plasma nitrided steel, the duplex system also displayed a ‘‘nitriding layer erosion period.’’ The cavitation erosion rate determined for this stage (2.2 mg h1) was similar to that obtained for the plasma nitrided steel (Table 1), but this period extended up to 12.18 h for the duplex Cr1xNx system. An accelerated erosion stage (fourth period) also occurred in the duplex coating, being characterized by a lower erosion rate (3.9 mg h1) than the accelerated erosion period found in the AISI 1045 steel and non-duplex coating (8.470.1 mg h1). Fig. 2 shows hardness and final indentation depth (hf) versus force plots. According to the ISO 14577-4 International Standard, the hardness of hard coating can only be determined with a sharp (small tip radius) indenter that causes yielding within the coating. In general, this will only occur when hc/tco0.5 (ratio of maximum contact depth to coating thickness). For a spherical indenter the relation between the contact depth variable (hc) with final indentation depth (hf) is equal to [1]: hc ¼
hmax þ hf 2
(4)
We can conclude that hc4hf. Coating thickness was evaluated by SEM analyses. For the non-duplex Cr–N system, a thickness equal to 3 mm was determined; for the duplex system, a 5 mm value was estimated. Considering hf values obtained in our tests and the relation hc/tco0.5, for determining final coating hardness,
Fig. 1. Plots showing experimental data and fitted continuous piecewise linear regression models demonstrating the different wear periods: (a) AISI 1045 steel, (b) plasma nitrided AISI 1045 steel (c) non-duplex Cr1xNx system and (d) duplex Cr1xNx system.
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Table 1 Cavitation erosion rates and 95% confidence intervals obtained for each wear regime from statistical analyses (piecewise linear regressions). Specimen
AISI 1045 steel Plasma nitrided AISI 1045 steel Non-duplex CrxN1x Duplex CrxN1x
Erosion rate (mg h1) Incubation period (first regime)
Second regime
Third regime
Fourth regime
2.8270.06 (0.00–1.23 h) 0.1270.05 (0.00–3.40 h) 0.4470.04 (0.00–3.56 h) 1.870.1 (0.45–2.14 h)
8.470.1 (1.23–8.00 h) 2.070.3 (3.40–6.40 h) 4.470.2 (3.56–6.05 h) 0.970.2 (2.14–9.24 h)
5.270.6 (6.40–20.0 h) 8.470.3 (6.05–15.50 h) 2.270.2 (9.24–12.18 h)
3.970.3 (12.18–20.50 h)
Nonexistent regimes.
Table 2 Amplitude parameters for 3D surface roughness. Specimen
Sa (nm)
CI (nm)a
Sq (nm)
CI (nm)a
AISI 1045 steel Plasma nitrided AISI 1045 steel Non-duplex CrxN1x system Duplex CrxN1x system
64.6 119.0 69.0 111.5
61.9–67.2 112.6–125.4 63.5–74.5 107.5–115.5
95.0 158.5 102.0 158.5
91.8–98.2 152.9–164.1 98.5–106.5 156.1–160.9
a
C.I. 95%
Fig. 2. (a) Hardness and (b) final penetration depth data from nanoindentation measurements carried out as a function of applied force.
the indentation depth must be lower than 1.5 mm (1500 nm) for the non-duplex system and 2.5 mm (2500 nm) for the duplex system. Moreover, the recommended Ra value shall be 5% of the maximum indentation depth, i.e., hmax420 Ra [1]. In our 3D studies, the parameters Sa (roughness average, i.e., arithmetic mean deviation of the surface from the mean plane) and Sq (rootmean-square deviation of the surface, RMS) were used for
Fig. 3. Plots showing transversal Knoop hardness in (a) plasma nitrided AISI 1045 steel and (b) duplex system. Case depth was estimated about 200 mm.
characterising the amplitude property of surfaces (Table 2). AISI 1045 steel and the non-duplex system showed lower roughness than plasma nitrided AISI 1045 steel and the duplex system.
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Using instrumented indentation tests, the two conditions imposed for evaluation coating hardness, without any significant influence of the substrate, resulted in: for Cr–N duplex system, hc/tco0.5 (hmaxo2500 nm) and hmax420 Sa (hmax42180 nm) and for Cr–N non-duplex system, hc/tco0.5 (hmaxo1500 nm) and hmax420 Sa (hmax41380 nm). We can see in Fig. 2, for the duplex system these conditions are reached only for the force equal to 1000 mN. The Cr–N coating hardness was evaluated as (9.470.59) GPa and the elasticity modulus as (23679.34) GPa. For Cr–N non-duplex system, the force equal to 200 mN satisfied the two imposed conditions and gave a hardness of (1870.69) GPa and an elasticity modulus of (27877.20) GPa. However, considering that the coating hardness value as the value that does not depend upon the force applied in small depths, as shown in Fig. 2, Cr–N coating hardness can be determined as an average value for forces smaller than 200 mN
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for both systems. This procedure gives a hardness equal to (23.073.59) GPa in the duplex system and equal to (24.874.02) GPa for the non-duplex system. We can see an inconsistency between this procedure and those determined by the ISO 14577-4:2002 International Standard. The high roughness found in the duplex system led to that inconsistency. We concluded that a very smooth surface is fundamental to estimate the actual hardness by nanoindentation. According to Bobji et al. [8], as the roughness of a surface is increased, the hardness measured at depths comparable with the roughness scale deviates increasingly from the actual hardness. Hardness of all systems (coating+plasma nitrided layer+steel substrate) was obtained using a considerable force. Considering F ¼ 4000 mN, we have determined for AISI 1045 steel a hardness value of (3.0170.03) GPa; for plasma nitrided AISI 1045 steel, a hardness of (4.170.3) GPa; for Cr–N non-duplex, hardness of (4.0570.07) GPa and for the duplex system a hardness equal to (5.770.3) GPa. Surface modifications increased hardness for all systems; even when using high force (with significant influence of the substrate) we experienced modifications in the final hardness value. The duplex system showed the maximum hardness value for all forces applied. Knoop hardness measurements, as a function of the depth from the surface, were also performed under a force of 9.81 102 N. Fig. 3 shows the dependence of hardness of the transversal profile. The hardness profile of the nitrided layer as well as its thickness was determined for Nitrided AISI 1045 steel system and the duplex system. A hardening depth equal to 200 mm (2.0 105 nm) was estimated. The maximum indentation depth reached in instrumented indentation tests, when the hardness was measured, was in the order of 7.75 103 nm. The dependence between depth of cavitation damage and hardening depth was also investigated. The 3D surface topography was manipulated and inverted upside down so that the cavitation craters appeared as summits for evaluating the depth of cavitation damage (Fig. 4). Prior to invert the surface, a procedure called ‘‘truncation’’ was employed. Truncation can be understood as the removal of the material above a given plane that is parallel to the mean plane. It may be used to reveal features below certain levels that are otherwise invisible. The truncation level used in our analyses was equal to Sr1 (upper bearing surface) (Fig. 5). Sr1 is the fraction of the surface that consists of small peaks above the main plateau [5,9]. The objective was to remove discrepant peaks present in the original surface. After this procedure, histograms
Sk parameters, gaussian filter, 0.8 mm. Spk = 0.52 µm
Sk = 1.44 µm
0 20 Sr1 = 6.92% Sa1 = 18 µm2/mm Fig. 4. Inverted isometric plots displaying the damages produced by cavitation erosion in the final time studied in each system.
Svk = 4.04 µm
40
60
80 100% Sr2 = 76.5% Sa2 = 476 µm2/mm
Fig. 5. Plot of bearing area ratio illustrating the determination of Sr1 (upper bearing surface).
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Fig. 6. Height distribution histograms for each erosion period determined by the statistical piecewise linear regression analyses.
containing the number of peaks (craters) per mm2 versus amplitude were obtained for each final erosion time specific to each cavitation erosion period (Fig. 6). Fig. 4 shows the inverted isometric plots for the cavitation erosion time equal to approximately 15 h, excluding AISI 1045 steel. The left scale in this figure corresponds to the surface total amplitude. The systems where the plasma nitriding process was
carried out presented shallower craters. In the non-duplex system, the cavitation erosion was characterized by deeper craters. Once the damage reaches the substrate, no obstacle hinders its propagation. Fig. 6 shows the histograms containing the number of ‘‘peaks’’ (cavitation craters) per mm2 versus amplitude, for each stage determined by the statistical piecewise linear regression analyses,
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in the final period associated to the specific stage. The following aspects can be observed: (i) The main aspects of the systems before the cavitation erosion process are: the nitrided systems show the value of amplitude that occurs most frequently (mode, statistical parameter) between 0.40 and 0.45 mm. Similar values can be observed for Cr–N duplex system (0.40–0.48 mm). These amplitudes are higher than those identified for the Cr–N non-duplex system (0.33–0.36 mm), confirming the higher roughness for plasma nitriding systems. (ii) In the final erosion time of the ‘‘incubation period,’’ the highest amplitudes (deepest craters) are present in the nonduplex system. Erosion cavitation was more pronounced in it, and plasma nitriding process appears again as the most effective method to prevent the formation of deep craters. (iii) The deposition of a coating under a hardening case is a means to guarantee shallower craters: this is illustrated in the ‘‘coating erosion period.’’ In fact, SEM images confirmed a completely different aspect of the cavitation erosion mechanism; in the non-duplex system the erosion occurs by localized and profound craters, on the other hand, in duplex systems the mass loss occurs by the slow decohesion of the coating. The amplitude mode was between 13 and 14 mm for the nonduplex system and between 6 and 6.5 mm for the duplex system. (iv) The maximum amplitude presented in the histogram (the maximum value for the x-axis), corresponding to the ‘‘nitrided layer erosion period,’’ was 24.7 mm for nitrided steel system and equal to 40.3 mm for the duplex system. These values are significantly lower than that of the case depth determined, i.e., 200 mm. This result suggests that the cavitation erosion in this stage probably corresponds to the cavitation erosion of the compound layer. Fig. 7 indicates a value for this layer thickness equal to approximately 5 mm. (v) In the ‘‘substrate erosion period,’’ the maximum amplitude depends on the erosion time. The AISI 1045 steel was tested for only 8 h, because severe damage was already observed in this time. plasma nitrided AISI 1045 steel system presented a maximum crater amplitude equal to 54 mm, after a erosion time of 12.5 h. For the same erosion time, Cr–N non-duplex system presented the maximum amplitude equal to 155 mm. Thus we concluded that the Cr–N non-duplex system undergoes a much more accentuated change with the cavitation
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erosion time. On the other hand, Cr–N duplex system, after an erosion time of 15.5 h, presented maximum crater amplitude equal to 73.5 mm, whereas in 12 h, this amplitude was equal to 40.3 mm. The duplex system, similarly to the nitrided plasma steel system, experiences minor damages with the cavitation erosion process. (vi) Even in the erosion times equal to 20 h, no plasma nitriding system presents damage with extension greater than the case depth, estimated equal to 200 mm. For these systems, the maximum amplitude of craters was still restricted within the case depth. Indeed surface hardening by the nitriding process offers a more effective barrier to cavitation erosion. The junction coating and plasma nitriding was the most effective system, according to profilometric analyses. We can conclude that significant improvements in the cavitation erosion resistance can be achieved when thick nitrided cases are present. If it is true, nanohardness, evaluated when the indentation depth is small, does not correlate effectively with the performance in cavitation erosion. For verifying which hardness (nano, micro or macrohardness) should be evaluated in cavitation erosion, plots of H2/E, for two forces used in instrumented indentation tests, were constructed. The following ratio— indentation hardness to the square ðHIT Þ2 and indentation elasticity modulus ðEIT Þ—was employed due to the known relation between resilience and cavitation erosion. The modulus of resilience is defined as the strain energy per unit volume required to stress the material from zero stress to the yield stress given by [10] Ur ¼
S20 2E
(5)
where S0 is the yield stress and E is the modulus of elasticity. Considering the proportionality between indentation pressure (hardness) and yield stress we have decided to correlate cavitation rate and H2/E, instead of S2/E. Local determination of yield stress in different penetration depths is almost impossible to be measured and on the other hand, the nanoindentation instrument allows the measurement of the hardness in different penetrations using different forces, even in very small depths. In coated systems, it is a powerful tool for determining hardness for different penetration depths. EIT values, determined as function of different forces, are showed in Table 3. Fig. 8 was constructed from two cavitation erosion rates: the erosion rate determined for the ‘‘substrate erosion period’’ and those determined for the ‘‘incubation period’’ (Table 1). This first value of rates was related to hardness values which were measured with forces that caused the greater penetration. In our study the force applied was equal to 8000 mN for the depth equivalent to between 9000 and 10 000 nm. For the ‘‘incubation period’’ the force used was equal to 15 mN corresponding to between 185 and 400 nm of indentation depth.
Table 3 Indentation Modulus (EIT) and indentation hardness (HIT) for 8000 and 15 mN. Systems
force (mN)
EIT (GPa)
s
HIT (GPa)
s
Cr–N non-duplex
8000 15 8000 15 8000 15 4000 15
222 381 272 374 192 204 230 254
11 11 16 49 5 13 13 10
3.73 30.2 4.53 26.9 3.51 7.50 3.01 4.10
2 102 2 100 1 101 6 100 1 101 1 100 3 102 6 101
Cr–N duplex Nitrided AISI 1045 steel AISI 1045 steel Fig. 7. SEM photomicrograph of the duplex system indicating the presence of the CrxN1x coating, the compound layer and iron nitrides in plasma nitrided layer.
s, Standard deviation.
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or macrohardness should be evaluated rather than nanohardness, to successfully correlate the material properties to its performance in cavitation erosion.
Acknowledgments The authors are greatly indebted to CNPq-Conselho Nacional de Desenvolvimento Tecnolo´gico, Brazil, and to FAPEMIG-Fundaca˜o de Amparo a` Pesquisa de Minas Gerais, Brazil, for the financial support. They would also like to thank Mr. Renato Machado for nanoindentation measurements at the National Institute of Standards and Technology, NIST, USA. References [1] International Organization for Standardization, Geneva. ISO 14577/1-4 Metallic materials-Instrumented indentation test for hardness and materials parameters. Geneva, 2002. [2] Godoy C, Mancosu RD, Lima MM, Branda˜o D, Housden J, Avelar-Batista JC. Influence of plasma nitriding and PAPVD Cr1x Nx coating on the cavitation erosion resistance of an AISI 1045 steel. Surf Coat Technol 2006;200:5370–8. [3] Child HC. Surface hardening of steel. Engineering design guides. Oxford: Oxford University Press; 1980. [4] American Society for Testing and Materials Standards, ASTM G 32-03. Standard test method for cavitation erosion using vibratory apparatus, 2003. [5] Stout KJ, Blunt L. Three dimensional surface topography. 2nd ed. London: Penton Press; 1994. [6] Montgomery DC, Peck EA. Introduction to linear regression analysis. 2nd ed. New York: Wiley; 1992. [7] Statistica DX. Software for Windows, Version 7, StatSoft South America, 2005. [8] Bobji MS, Biswas SK. Estimation of hardness by nanoindentation of rough surfaces. J Mater Res 1998;13(11). [9] Stout KJ. Development of methods for the characterisation of roughness in three dimensions. London: Penton Press; 2000. [10] Dieter GE. Mechanical metallurgy. McGraw-Hill Kogakusha, Ltd.; 1976.
Fig. 8. Plots of cavitation erosion resistance and resilience for two forces used in nanoindentation tests: (a) 15 mN and (b) 8000 mN.
In Fig. 8, we can see a better relation between cavitation erosion resistance and resilience when hardness was calculated using greater penetrations (forces). This result suggests that micro