The impact of cold work and hard phases on cavitation and corrosion resistance of high interstitial austenitic FeCrMnMoCN stainless steels

Wear 376-377 (2017) 1009–1020

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The impact of cold work and hard phases on cavitation and corrosion resistance of high interstitial austenitic FeCrMnMoCN stainless steels P. Niederhofer n, S. Huth, W. Theisen Ruhr-Universität Bochum, Chair of Materials Technology, 44780 Bochum, Germany

art ic l e i nf o

a b s t r a c t

Article history: Received 1 September 2016 Received in revised form 28 December 2016 Accepted 4 January 2017

High interstitial austenitic stainless steels have been shown to exhibit superior mechanical properties, which include a unique combination of high strength and high toughness, due to the positive effects of the combination of carbon and nitrogen in solid solution. By the addition of molybdenum, a significant improvement of their localized corrosion resistance has been achieved. Further strengthening is necessary in the case of application in environments featuring both high mechanical loads and corrosive attack, e.g. in bearings in sea water. It can be induced by cold work hardening as well as the addition of hard phases, which in turn can affect the wear and corrosion resistance. In this study, the impact of cold work and the precipitation of niobium carbonitrides on the resistance to cavitation erosion, localized, and general corrosion has been investigated. Two newly developed high interstitial FeCrMnMoCN steels were analyzed by vibratory cavitation testing in distilled water and potentiodynamic polarization measurements in sodium chloride solution and sulfuric acid after cold work strengthening. The latter was induced by cold rolling to different degrees. Microstructural characterization was performed by hardness testing, optical, and scanning electron microscopy. The results show improved strength but decreased cavitation erosion resistance caused by the hard phases. In contrast, neither the localized nor the general corrosion resistance seem severely affected. The cold rolling leads to intense work hardening and enhanced cavitation erosion resistance, while the corrosion behavior is not significantly influenced. In the case of cavitation erosion, the improved resistance of the cold work hardened steel matrix seems to dominate the negative effect of the hard phases. The combination of high wear and high corrosion resistance, even in severely cold work strengthened condition, makes the FeCrMnMoCN austenitic stainless steels promising candidates for application in harsh environments. & 2017 Elsevier B.V. All rights reserved.

Keywords: High interstitial Austenite FeCrMnMoCN Cavitation erosion Corrosion Cold work

1. Introduction The evolution of materials featuring a major resistance to cavitation erosion (CE) is highly interesting, since this kind of wear can cause severe functional drawbacks in different applications, e.g. rotating components in hydraulics or different fluids such as sea water [1,2]. Depending on the liquid's characteristics, different types of corrosion can occur in such applications as well, e.g. localized pitting corrosion in the case of media containing chloride ions. On the other hand, in acids like sulfuric acid, general surface corrosion appears [3]. In view of corrosion resistance, single-phase stainless steels like conventional FeCrNi austenites seem appropriate. However, their mechanical properties, primarily strength, are rather low, and n

Corresponding author. E-mail address: [email protected] (P. Niederhofer).

http://dx.doi.org/10.1016/j.wear.2017.01.016 0043-1648/& 2017 Elsevier B.V. All rights reserved.

consequently, such is their resistance to CE. Interstitial strengthening by nitrogen has been shown to effectively improve mechanical properties of both FeCrNi as well as FeCrMn austenites, amongst others strength, work hardening ability, and fatigue life, while toughness and ductility are slightly decreased. The effect is even larger if nickel is substituted by manganese, because the latter promotes nitrogen solubility and thus, higher amounts can be introduced (high nitrogen steels, HNS). However, alloying with very large contents of nitrogen requires expensive pressurized metallurgical methods [4]. Several studies revealed enhanced CE resistance of both FeCrNiN [5–12] and FeCrMnN [13–16] austenitic stainless steels or layers. The work hardening ability can be used to further strengthen austenites, in particular HNS [4], and consequently, cold deformation prior to cavitation testing was found to improve CE resistance [12,14,17,18]. More recent research revealed that the joint alloying of carbon and nitrogen (high interstitial steels, HIS) exhibits several advantages over high nitrogen steels. For example, the mechanical properties of FeCrMnCN alloys are

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even improved compared to FeCrMnN steels and they can be cast at ambient pressure, as summarized in [19]. Cavitation erosion resistance of high interstitial austenitic stainless steels is superior as well and can be correlated with their increased ability to absorb impact energy elastically, which in turn means a higher resistance to plastic deformation. Furthermore it was indicated that CE resistance can still be improved by cold work strengthening [20–22]. With regard to surface corrosion, high interstitial austenitic stainless steels alloyed with manganese are known to exhibit lower resistance compared to FeCrNi alloys. On the other hand, the interstitial elements carbon and especially nitrogen induce improved pitting corrosion behavior, if they are entirely dissolved in the austenitic matrix [23–25]. A potential reason for the positive influence of the interstitial elements may be the high density of free electrons, which in addition to improved mechanical properties causes short range atomic order and thus reduces clustering of chromium. Consequently, the chromium atoms are more homogeneously distributed, which may result in increased stability of the passive layer [26]. Molybdenum is known to improve pitting corrosion resistance as well, however, the reasons are not yet fully understood. The combination of nitrogen and molybdenum is linked with enhanced repassivation potential in the case of localized corrosion [27,28]. Generally, nonmetallic inclusions, especially sulfides but also carbides or nitrides as well as dislocations, may locally lead to weakness of the passive layer, resulting in preferential sites for the initiation of pitting corrosion [29,30]. While cold working of austenitic stainless steels improves the mechanical properties, as already mentioned, it is considered as detrimental for both surface as well as pitting corrosion resistance. In the case of surface corrosion, small differences in resistance were reported after cold working of both FeCrNi as well as FeCrMnMoN steels. Increasing current density in the passive region in a FeCrNi alloy was attributed to deformation-induced martensite as well as changes in texture. For the FeCrMnMoN steel, generally a high stability of the passive layer even in highly cold worked condition was claimed. However, no direct correlation of passive current densities with degree of deformation could be found [31,32]. The impact of cold working on pitting corrosion resistance of austenitic stainless steels is not generally clarified and thus, different potential mechanisms are discussed. On the one hand, strain-induced martensite is held responsible for decreasing pitting potentials with rising degree of deformation in some studies concerning FeCrNi steels [31,33]. On the other hand, lower pitting resistance is frequently attributed to higher dislocation densities inducing localized degradation of the passive layer for FeCrNi(N), FeCrMnNiN as well as FeCrMnMoN steels [34– 38]. In the latter case, changes in the composition of the passive layer (less oxides but more hydroxides with increasing degree of deformation) were additionally reported [32]. However, neither the amount nor the reason of a detrimental effect of cold work on pitting corrosion resistance is yet fully understood. In a previous study, using laboratory melts of high interstitial austenitic stainless steels, a clear positive influence of molybdenum on pitting corrosion resistance was found. Furthermore, a positive impact of cold work on CE resistance was indicated. However, a distinct influence of prestraining by tensile deformation became not obvious [22].

For the current study, specimens of high interstitial austenitic stainless steel containing varying amounts of molybdenum and niobium (to induce the precipitation of hard phases), were investigated with focus on their resistance to cavitation erosion and pitting, as well as surface corrosion. Furthermore, the influence of severe cold work strengthening on both wear and corrosion behavior was studied systematically.

2. Experimental procedure 2.1. Materials Two different newly developed high interstitial FeCrMnMoCN austenitic stainless steels were investigated within this study, each nominally containing around 18 mass% of chromium and manganese, 2.5 mass% of molybdenum, between 0.35 and 0.45 mass% of carbon and 0.61 to 0.7 mass% of nitrogen. One alloy additionally comprises 1 mass% of niobium. For comparison purposes, a similar alloy without addition of molybdenum and niobium was investigated as well. The actual chemical composition, measured by optical emission spectroscopy, is listed in Table 1. Also given are values derived from tensile testing. All three high interstitial austenites had been produced in industrial scale by induction melting, electroslag-remelting and hot-working. Prior to examination, they were solution annealed at temperatures between 1125 and 1175 °C for 30 min followed by rapid quenching in water, which led to fully austenitic microstructures free of undesired precipitations. For naming convention, the alloys are designated with C and N, and a number, indicating the sum of carbon and nitrogen in mass%. Additions of molybdenum and niobium are provided as well. 2.2. Cold working and hardness testing After solution annealing, cold rolling was performed to different logarithmic deformation degrees (φ) up to a value of approximately 1.2. The effect of cold work strengthening was investigated by means of Vickers macrohardness measurements according to DIN EN ISO 6507 [39] with a KB30S hardness tester made by KB Prüftechnik and using a characterization force of 98.7 N (HV10). All given values are averaged from at least three single measurements. 2.3. Specimen preparation and metallographic characterization The specimens for the CE test (surface area of 20
20 mm, and 2–5 mm in thickness) were prepared by cutting, followed by mechanical grinding with grits of 54 and 18 μm. Final polishing was performed with grain sizes of 3 and 1 μm. The specimens were examined prior to cavitation testing as well as during the weighing interruptions by a conventional Olympus BX60M light microscope and by SEM using a GEMINI LEO 1530 VP featuring a working distance of 10 mm and an acceleration voltage of 20 kV.

Table 1 Chemical composition of examined materials in mass% as well as yield (Rp0.2) and tensile strength (Rm) in MPa, and elongation at fracture (A) in %. Alloy

C

N

Cr

Mn

Mo

Nb

Ni

Si

Rp0.2

Rm

A

CN1.12MoNb CN1.1Mo CN1.07

0.45 0.42 0.49

0.67 0.68 0.58

16.7 17.8 18.8

18.3 19.1 18.9

2.65 2.52 0.07

0.94 – –

0.24 0.14 0.40

0.36 0.28 0.43

610 595 604

1140 1095 1075

59 70 74

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2.4. Cavitation erosion tests

3. Results

Vibratory cavitation erosion testing was performed by use of an ultrasonic cavitation horn UIP2000hd made by Hielscher. It was operated in distilled water, which has kept at room temperature, at a frequency of 20 kHz and an amplitude of 40 μm. The test setup in general is according to G32-10 [40], but deviant from the standard, an indirect method with a distance of 0.5 mm between the specimen and the sonotrode tip were used. Tests were interrupted in specific intervals, and the specimens were cleaned, dried, and weighed on a Sartorius Mechatronics balance providing an accuracy of 0.01 mg. Cumulative mass loss was calculated and plotted versus time. From the resulting mass loss curves, length of incubation period (IP) and maximum erosion rate (MER) were derived by means of the tangent method. In the case of CN1.12MoNb and CN1.07, the experiments were repeated and consequently, mass loss is an average value from two specimens. The error bars in the corresponding curves indicate the highest and the lowest mass loss values.

3.1. Cold work strengthening

2.5. Corrosion tests Surface as well as pitting corrosion resistance were investigated using potentiodynamic polarization curves according to G5-13 [41] in 0.5 mol sulfuric acid (H2SO4) as well as 3% sodium chloride (NaCl) solution. A corrosion test cell featuring a three electrode configuration with the specimen being the working electrode, a platinum sheet serving as counter electrode and a saturated calomel- (Hg2Cl2-) reference electrode (SCE, with a potential of þ244 mV referred to the standard hydrogen electrode (SHE)) was used. Control and measurement of potential and current was performed with different potentiostats. In the case of surface corrosion testing, a type Vertex potentiostat made by Ivium was utilized, while pitting corrosion testing was conducted by means of a PGZ301 potentiostat manufactured by Radiometer Analytical. The preparation for the corrosion tests was executed by establishing an electrical contact to the backside of the specimens with a spot-welded wire. Subsequently, the specimens were embedded in polymer resin and ground with SiC paper of 1000 mesh. Remaining gaps between embedding resin and the specimen's surface as well as the edges of the latter were sealed using lacquer in order to prevent crevice corrosion. The residual surface was measured by use of a flat bed scanner and image analysis software. The electrolyte was rinsed with N2 gas for a duration of 30 min prior to the experiment in order to provide a constant content of oxygen. Subsequently, a comparable initial state was generated by performing cathodic polarization at 2244 mV for 60 s prior to the actual corrosion test. The latter consisted of the measurement of the open circuit potential (OCP) followed by recording of the current density-potential-curve, which began 10 mV below OCP and was operated using a polarization rate of 0.167 mV/s. All potentials within this study are measured and quoted compared to that of the calomel reference electrode. At surface corrosion testing, the current densities during passivation and especially in the passive region are evaluated as most important values. In contrast, during pitting corrosion testing, the significant value is the breakdown or pitting potential, which designates the initiation of stable growth of pittings resulting in failure of the passive layer [3]. According to G150-13 [42], it is technically defined as the potential, where the current density constantly transcends a value of 100 μA/cm2.

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As described in Section 2.2, strengthening of specimens of the high interstitial austenitic stainless steels was performed by means of cold rolling after solution annealing and quenching. This resulted in an increase in hardness, which is shown in Fig. 1 compared to the initial state. Without prior cold rolling, CN1.12MoNb (320 HV10) exhibits slightly higher hardness than CN1.1Mo (280 HV10) and CN1.07 (260 HV10). Increasing degrees of deformation correlate with rising hardness values. At φ ¼0.9, both alloys containing molybdenum show a hardness of approximately 610 HV10, while the value of CN1.07 is lower, at 570 HV10. One specimen of CN1.12MoNb was rolled up to a final deformation degree of approximately 1.2, resulting in a hardness of 660 HV10. 3.2. Cavitation erosion testing Cavitation erosion tests were performed on specimens which were exposed to different degrees of cold rolling, as described in Section 2.4. The resulting mass loss curves are shown in Fig. 2 for each steel in different degrees of deformation. Exposure to cavitation almost immediately results in marginal mass loss, which indicates the absence of a real incubation period. The calculated values of maximum erosion rate and length of the incubation period are listed in Table 2. An increasing degree of deformation leads to a significantly higher resistance to cavitation erosion. This can be seen for all investigated alloys by comparing the cumulative mass loss after 80,000 s, which decreases with increasing amount of cold work. This is caused by lower maximum erosion rates, while the length of incubation periods do not exhibit distinct tendencies. The comparison of the mass loss curves of the different alloys reveals an obviously detrimental influence of the niobium carbonitrides. In the case of CN1.12MoNb, mass loss is higher compared to CN1.1Mo. This holds true for all degrees of deformation. However, CN1.07 shows the highest mass loss among the investigated alloys. 3.2.1. Surface changes The evolution of cavitation erosion damage was investigated in the case of CN1.12MoNb and CN1.07 for all degrees of deformation during the weighing interruptions. Selected images are shown in Figs. 3 to 10. In the case of CN1.1Mo, SEM images of the surfaces were only taken prior to and after 80,000 s of cavitation testing. This is displayed in Fig. 11. Regarding the surfaces prior to cavitation testing in the case of

Fig. 1. Hardness depending on degree of deformation.

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Fig. 2. Cumulative mass loss induced by cavitation erosion testing of the investigated alloys: (a) CN1.12MoNb, (b) CN1.1Mo, and (c) CN1.07.

Table 2 Maximum erosion rate (MER) and length of incubation period (IP). Alloy

φ

MER in 10  5 mg/s

IP in s

CN1.12MoNb

0 0.3 0.6 0.9

8.29 6.30 4.44 1.95

30,166 34,762 33,577 30,000

CN1.1Mo

0 0.3 0.6 0.9

7.42 5.00 1.45 0.53

35,445 36,400 31,034 38,095

CN1.07

0 0.3 0.6 0.9

13.0 6.78 6.08 2.11

31,538 30,185 36,214 36,450

CN1.12MoNb (Figs. 3(a) to 6(a)) it becomes obvious, that the Nb(C, N) are affected by cold rolling. With rising degree of deformation, an increasing number of niobium carbonitrides (Nb(C,N)) exhibits either cracks or partial fractures. After cavitation erosion testing for 10,000s in undeformed condition (Fig. 3(b)), erosion mainly initiates both at grain/twin and phase boundaries. With rising degree of deformation, the origin of erosion is shifted to phase boundaries and leads to fracture or entire removal of the niobium carbonitrides, resulting in the formation of pits, whereas the austenitic matrix seems less affected (Figs. 4(b) to 6(b)). This holds true after longer duration of cavitation testing as well (Figs. 3 (b) and (c) to 6(b) and (c)). Along with an increasing degree of

deformation, the austenitic matrix is less affected by cavitation erosion, while the pit formation originating from the Nb(C,N) becomes more severe. In the case of the single-phase austenitic steels without prior cold work, damage arises from extrusion at grain/twin boundaries and slip lines, which can be seen for CN1.07 after 10,000 s of cavitation testing in Fig. 7(a). After 80,000 s (Fig. 7(b)), a rather uniform surface damage can be observed. Despite the missing hard phases, the effect of cold rolling prior to cavitation testing is quite similar. The number of visible extrusions arising from slip lines and of slip lines themselves decreases with rising degree of deformation, whereas the number of localized pits increases. This can be seen when comparing the different degrees of deformation after 10,000 s of cavitation testing in Figs. 7(a) to 10(a). In consequence, after 80,000 s, an increasing fraction of surface areas shows higher resistance to cavitation erosion with rising degree of cold work (Figs. 7(b) to 10(b)). This fact further indicates an influence of crystallographic orientation on the CE resistance. Fig. 11 shows the surfaces of specimens of CN1.1Mo in different degrees of deformation after 80,000 s of cavitation testing. Compared to the corresponding images of CN1.07 (Figs. 7(b) to 10(b)) a similar influence of cold work on the CE resistance becomes obvious, which includes evidence of an impact of crystallographic orientation. Furthermore, the CE resistance seems to be higher compared to CN1.07, which is particularly apparent in the case of the highest degree of deformation (φ ¼0.9, Fig. 11(d)), where only rather small amounts of cavitation erosion damage can be seen.

P. Niederhofer et al. / Wear 376-377 (2017) 1009–1020

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10000 s

a)

100 µm

100 µm

80000 s

40000 s

c)

b)

100 µm

d)

100 µm

Fig. 3. Surface changes induced by cavitation erosion of CN1.12MoNb without cold work: (a) prior to, (b) after 10,000 s, (c) after 40,000 s, and (d) after 80,000 s of cavitation testing.

3.3. Corrosion testing Fig. 12 shows the current density-potential-curves of the investigated alloys without prestraining and after cold rolling to a degree of deformation of 0.9, the latter only in the case of CN1.1Mo and CN1.12MoNb. The data of CN1.07 is shown for comparison purposes and was taken from previous work [43]. Unlike

cavitation erosion behavior, in the case of generalized corrosion, no significant influence of cold work on the resistance can be derived. The current density-potential-curves of the undeformed and the cold worked specimens are approximately similar. The main difference can be seen when comparing the curves of the steels containing molybdenum to CN1.07. The latter exhibits a higher passivation current density of approximately 70 compared

10000 s

a)

100 µm

100 µm

80000 s

40000 s

c)

b)

100 µm

d)

100 µm

Fig. 4. Surface changes induced by cavitation erosion of CN1.12MoNb, cold rolled to φ¼ 0.3: (a) prior to, (b) after 10,000 s, (c) after 40,000 s, and (d) after 80,000 s of cavitation testing.

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10000 s

a)

100 µm

100 µm

80000 s

40000 s

c)

b)

100 µm

d)

100 µm

Fig. 5. Surface changes induced by cavitation erosion of CN1.12MoNb, cold rolled to φ¼0.6: (a) prior to, (b) after 10,000 s, (c) after 40,000 s, and (d) after 80,000 s of cavitation testing.

to values of <10 μA/cm2 in the case of CN1.1Mo and CN1.12MoNb. However, the similar passive current density (2 to 3 μA/cm2 in any case) seems to be rather independent of the content of molybdenum as well as the slight differences in other alloying elements. The current density-potential-curves in 3% sodium chloride solution are shown in Fig. 13. Comparable to generalized corrosion,

no significant difference between specimens with and without cold rolling can be derived. Apart from that, a clear influence of alloying system becomes obvious by the distinctly lower pitting potential of CN1.07 (around 375 mV) compared to CN1.1Mo and CN1.12MoNb, while the latter two show comparable behavior (approximately 1100 mV).

10000 s

a)

100 µm

100 µm

80000 s

40000 s

c)

b)

100 µm

d)

100 µm

Fig. 6. Surface changes induced by cavitation erosion of CN1.12MoNb, cold rolled to φ¼0.9: (a) prior to, (b) after 10,000 s, (c) after 40,000 s, and (d) after 80,000 s of cavitation testing.

P. Niederhofer et al. / Wear 376-377 (2017) 1009–1020

10000 s

a)

1015

80000 s

100 µm

b)

100 µm

Fig. 7. Surface changes induced by cavitation erosion of CN1.07 without cold work: (a) after 10,000 s, and (b) after 80,000 s of cavitation testing.

4. Discussion The superior mechanical properties of high interstitial austenitic stainless steels, including a remarkable combination of strength and ductility [19] and a high fatigue resistance [44–46], are caused by the positive effects of alloying with both carbon and nitrogen. These are a low stacking fault energy, a high density of free electrons, and a large potential for cold work strengthening (Section 1). Their CE resistance has been shown to be superior [20]. The positive impact of cold work hardening on cavitation erosion behavior is already known from conventional austenites. A previous study indicated a similar effect in the case of FeCrMnCN steels as well [22], which in the present contribution has been investigated systematically. Besides, the influence of hard phases on both cold work strengthening and cavitation erosion behavior was analyzed in detail, given that the intentional precipitation of carbides or nitrides is a rather unconventional approach in the case of stainless austenitic steels. Corrosion behavior of high interstitial austenitic stainless steels has been found to be influenced by the interstitial elements carbon and nitrogen as well as molybdenum [22,26,43]. Due to the possible impact of cold work hardening reported in [22], this study includes the investigation of the impact of severe cold rolling and hard phases on both pitting and surface corrosion resistance. As described in Section 3.1 and shown in Fig. 1, rising degrees of deformation induced by cold rolling lead to significantly increasing hardness values. The cold work strengthening is accompanied by the formation and increasing density of twins as well as planar dislocation gliding [19,26]. The slightly higher hardness of CN1.12MoNb may be attributed to the lower grain size induced by the Nb(C,N) as well as their own strengthening effect. Furthermore, the substitutional element molybdenum supposedly also contributes to strengthening. However, the differences are not significantly pronounced. Regarding potential applications which

10000 s

a)

require high strength, a bearing ring made from CN1.07 has been manufactured, exhibiting a hardness of 700 HV in the area which was exposed to the highest degree of deformation [47]. In view of additional demand by pitting corrosion, e.g. in bearings in sea water applications, both CN1.1Mo and CN1.12MoNb were developed [21]. The latter was intentionally alloyed with niobium in order to precipitate hard phases, which is typical in conventional bearing steels, and may enhance resistance to abrasive or adhesive wear. In terms of strengthening, the comparison of hardness induced by cold rolling in Fig. 1 reveals a similar or even improved strengthening potential for both CN1.12MoNb and CN1.1Mo compared to CN1.07. 4.1. Cavitation erosion resistance As described in Section 3.2, the mass loss curves obtained from cavitation erosion testing of the different high interstitial austenitic stainless steels revealed a positive influence of prior cold work strengthening as well as a detrimental effect of the Nb(C,N). Both will be discussed in the following Sections considering the differences in evolution of surface damage. 4.1.1. Influence of cold work strengthening The positive influence of increasing amounts of cold rolling becomes obvious from Fig. 2 for all investigated alloys. However, when regarding the data listed in Table 2, it can be derived, that the improvement is governed by lowering of the maximum erosion rates, while the lengths of the incubation periods do not exhibit distinct tendencies. Prior cold working was already shown to increase CE resistance for different austenitic steels [12,14,17,18]. In contrast to the present study, both increasing length of incubation period as well as lower maximum erosion rates were reported. This may be a consequence of the rather short durations used in CE testing (80,000 s, Section 3.2). However, this was

80000 s

100 µm

b)

100 µm

Fig. 8. Surface changes induced by cavitation erosion of CN1.07, cold rolled to φ¼ 0.3: (a) after 10,000 s, and (b) after 80,000 s of cavitation testing.

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10000 s

a)

80000 s

100 µm

b)

100 µm

Fig. 9. Surface changes induced by cavitation erosion of CN1.07, cold rolled to φ¼ 0.6: (a) after 10,000 s, and (b) after 80,000 s of cavitation testing.

sufficient to reveal significant differences, as described in Section 3.2. Mechanical properties of austenitic stainless steels are also affected by cold work strengthening. For example, not only the quasi-static [26] but also the cyclic yield strengths [48] of high interstitial austenites have been reported to be improved by cold deformation. As has been shown in prior investigations, CE resistance can be explained by a material's ability to absorb impact energy elastically, which, by trend, can be correlated with respective micromechanical properties derived from instrumented indentation testing [8,18,20]. Cold work strengthening of austenitic steels increases the amounts of energy, which can be absorbed elastically, as was also shown in the case of a fully austenitic layer on a duplex stainless steel induced by high temperature gas nitriding [12]. Similar observations were made in a FeCrMnMoCN alloy in laboratory scale, which was prestrained by tensile deformation. Investigation by means of instrumented indentation testing showed both increasing indentation hardness as well as higher amounts of elastic fraction of the indentation energy with rising degree of deformation [21]. In consequence, this means an enhanced resistance to plastic deformation. Both has been shown to be, by trend, influenced by the amount of interstitial elements in FeCrMnCN steels [20,43]. This may contribute to the higher mass loss and thus lower CE resistance of CN1.07 compared CN1.1Mo and CN1.12MoNb (Fig. 2). Furthermore, regarding CE as a low cycle fatigue phenomenon [49,50], correlation with fatigue-derived properties seems suitable as well. Given that fatigue testing in general is elaborate and tests are often performed strain controlled, while during cavitation, the dominant factor seems to be the stress induced by microjets and shock waves, appropriate literature data is rare. However, the investigation of the influence of prior cold deformation on fatigue testing of a high interstitial austenitic FeCrMnCN stainless steel revealed that both cold work strengthening as well as higher amounts of interstitial elements increase the endurance limit [51],

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a)

which confirms the results of the current contribution. The impact of cold working on CE resistance is supported by the evolution of surface damage (Section 3.2.1). Generally, in the case of single-phase metals, surface damage induced by cavitation originates from extrusions of material at grain boundaries and former slip lines [50]. This was confirmed in the case of austenitic FeCrNi(N) and FeCrMnN austenites [8,14,49] and FeCrMnCN steels in particular [20]. The positive effect of cold working can be seen by larger areas of rather unaffected surface with increasing degree of deformation in all investigated austenites after CE testing (Figs. 3 to 6(d), 7 to 10(b), and 11), indicating the increasing resistance to plastic deformation, as already discussed. The similar effect can be observed already within the incubation periods (after 10,000 s) in the case of CN1.12MoNb (Figs. 3 to 6(b)) as well as CN1.07 (Figs. 7 to 10(a)). This is revealed by significantly less evidence of plastic deformation by dislocation movement, which are appearance of slip lines and formation of extrusions arising from them as well as grain and twin boundaries, with increasing degree of deformation. Furthermore, differences between CE resistances of individual grains become obvious, as described in Section 3.2.1. Influences of grain orientation on CE resistance of austenitic steels have been reported in different studies [9–12,52] and were attributed to deviations in resistance to plastic deformation due to orientation-dependent number of active slip systems. 4.1.2. Influence of hard phases In the case of CN1.12MoNb, as described in Section 3.2, a detrimental influence of the hard phases can be seen, when comparing the mass loss curves to that of CN1.1Mo (Figs. 2(a) and (b)). Both steels exhibit comparable chemical composition apart from niobium, which results in similar amount of alloying elements within the austenite. The different grain size may have an influence, but as shown in a previous study [20], in the case of high interstitial austenites, it is either insignificant or dominated by other effects. However, within this study, the alloy featuring the

80000 s

100 µm

b)

100 µm

Fig. 10. Surface changes induced by cavitation erosion of CN1.07, cold rolled to φ¼0.9: (a) after 10,000 s, and (b) after 80,000 s of cavitation testing.

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ϕ = 0.3

ϕ=0

a)

100 µm

ϕ = 0.6

c)

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b)

100 µm

ϕ = 0.9

100 µm

d)

100 µm

Fig. 11. Surface changes induced by cavitation erosion of CN1.1Mo, after 80,000 s of cavitation testing: (a) without prior cold work, (b) cold rolled to φ¼0.3, (c) cold rolled to φ¼ 0.6, and (d) cold rolled to φ¼ 0.9.

smaller grain size (CN1.12MoNb) exhibits larger mass loss, which is in contradiction to literature data, summarized e.g. in [20]. From the evolution of surface damage (Section 3.2.1), the

impact of the niobium carbonitrides on CE resistance becomes visible as well. While in the condition without prior cold working (Fig. 3), the appearance of erosion within the incubation period

Fig. 12. Current density-potential curves in 0.5 mol sulfuric acid of the investigated alloys: (a) CN1.12MoNb, (b) CN1.1Mo, and (c) CN1.07, the latter is taken from [43].

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Fig. 13. Current density-potential curves in 3% sodium chloride solution of the investigated alloys: (a) CN1.12MoNb, (b) CN1.1Mo, and (c) CN1.07, the latter is taken from [43].

(Fig. 3(b)) is not significantly different from single phase austenites, further cavitation testing results in formation of deep cavities originating from the Nb(C,N). This becomes even more obvious with increasing degree of deformation (Figs. 4 to 6), which is promoted by the fact, that increasing amounts of cold rolling induce a rising number of Nb(C,N) showing fracture already prior to cavitation testing. With increasing degree of deformation, the pits originating from the hard phases become deeper, while the austenitic matrix is less affected. In different multiphase materials, cavitation erosion damage arising from phase boundaries has been observed [53–55]. Furthermore, the formation of pits within the microstructure of steels (e.g. caused by fracture of hard phases) is claimed to increase load by further cavitation and thus to result in higher mass loss, as has been reported in different studies [18,43,56–58]. However, as can be seen from the mass loss curves, the overall CE resistance of CN1.12MoNb increases with rising amount of cold work despite its negative effect on the integrity of the carbonitrides. This underlines the dominant effect of the CE resistance of the austenitic matrix, which has already been discussed, at least in the case of relatively small contents of Nb(C,N). Similar observations have been made in the case of carbide-rich martensitic stainless steels manufactured by powder metallurgy [21,58], where the large chromium-rich carbides were preferentially attacked by cavitation, which resulted in partial fracture or entire removal. When comparing a specimen in fully martensitic condition to a specimen exhibiting high amounts of metastable retained austenite, the higher CE resistance of the metallic matrix of the latter (caused by strain-induced phase transformation to martensite) led to lower mass loss. This confirms the high importance of the CE resistance of the metallic matrix in the case of multi-phase materials.

4.2. Corrosion resistance As described in Section 1, both pitting and surface corrosion resistance of high interstitial austenitic stainless steels can be influenced by different factors. Among these are the chemical composition, nonmetallic inclusions, and cold work strengthening, which in many cases is required in order to improve mechanical properties. 4.2.1. Influence of molybdenum As already mentioned (Section 1), the improvement of localized corrosion resistance by molybdenum (expressed by the comparison of the pitting potentials of CN1.07 with those of the alloys containing molybdenum in Fig. 13) may be caused by enhanced repassivation behavior and suppression of initiation of pitting, which is especially pronounced in combination with nitrogen, as stated in literature [27,28]. The results of this study confirm those of a previous work [22] on similar alloys, but manufactured in laboratory scale. However, in the case of general corrosion, the effect of molybdenum is limited to a decrease of the passivation current density (Fig. 12), which leads to simplification of passivation and thus contributes to a higher corrosion resistance, while the passive current density remains unaffected. Similar observations have been described in literature for FeCrNi, FeCrMnN, and FeCrMnCN austenitic stainless steels [21,26,59]. In consequence, significant impact of alloying with molybdenum is obviously limited to localized corrosion behavior. 4.2.2. Influence of cold work As can be seen in Figs. 12 and 13, neither localized nor general corrosion of CN1.12MoNb and CN1.1Mo seem to be severely affected by prior cold rolling. In the case of CN1.07, no cold rolled

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specimen could be investigated. However, prior examination of a similar alloy prestrained by tensile deformation revealed comparable results to the present study [21]. If strain-induced martensitic transformation (formation of local corrosion elements at the boundaries between austenite and martensite) is considered as a potential reason for the claimed detrimental effect of prestraining on corrosion resistance (Section 1), this can be neglected for the FeCrMn(MoNb)CN steels due to their very high austenite stability [26,60]. Even at very large amounts of cold work, no formation of α-martensite has been observed. Prestraining of high interstitial austenites, if at all, was accompanied by only very small contents of ϵ-martensite due to the low stacking fault energy [26]. Compared to α-, the pitting corrosion behavior of ϵ-martensite is assumed not to deviate significantly from that of the austenitic phase due to the hexagonal structure [60]. Cold working of high interstitial austenites is connected with increasing density of dislocations [19,26], which may act as weaknesses in passive layers in general (Section 1). However, the rather unaffected localized corrosion resistance shown in the current contribution may be caused by a very high stability of the passive layers, probably supported by high amounts of molybdenum and nitrogen and the high density of free electrons, which promotes repassivation and leads to a fine dispersion of alloying elements by hindering clustering of carbon and chromium. These effects may counteract a detrimental impact of prestraining. Another possible explanation may be, that only parts of pitting corrosion behavior are affected by cold deformation, which cannot be investigated in isolation by potentiodynamic measurements, as stated in literature reviews [35,37]. For example, low degrees of deformation were found to induce an increasing number of metastable pittings, while higher amounts of cold work led to degraded repassivation behavior. By means of current density-potential-curves, such differences cannot be detected. However, as already mentioned, some studies reported of decreasing breakdown potentials of heavily cold deformed FeCrMnMoN steels in pitting corrosion testing [32,36]. This may serve as another indication for a potential impact of the density of free electrons, which has been shown to be significantly increased in the case of joint addition of carbon and nitrogen compared to exclusive nitrogen alloying [19,26]. As described in Section 3.3 and visible from Fig. 12, no significant influence of cold work on the passivation behavior in sulfuric acid was found. As a possible explanation may once again serve the high stability of the passive layer caused by short range atomic order (Section 1). Furthermore, surface corrosion is less susceptible to local inconsistencies of the passive layer than pitting behavior. Thus, a more detailed investigation of the passive layers of austenitic stainless steels, especially in severely cold deformed condition, may be of interest in the future. 4.2.3. Influence of hard phases Like mentioned in Section 1, nonmetallic inclusions such as manganese sulfides, but also carbides or nitrides have been found to act as defects in passive layers. Similar effects are attributed to dislocations. Such imperfections of passive layers often serve as preferential sites for the initiation of localized corrosion, as already stated. In the case of FeCrMnMoCN austenitic stainless steels manufactured by powder metallurgy (PM), a detrimental influence of oxides has been discovered [43]. In contrast, as described in Section 3.3, in the present study neither localized nor generalized corrosion resistance was significantly affected by the Nb(C,N). On the one hand, a possible explanation may be the high purity of niobium carbonitrides, which do not dissolve large amounts of chromium [61] thus avoiding chromium-depleted areas, which can occur if chromium-rich precipitations were present. On the other hand, referring to the PM alloys investigated in [43], the size

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of the Nb(C,N) is rather small compared to the oxides, which may serve as an explanation for their lower impact even on localized corrosion resistance, if a detrimental effect of second phase on the passive layer is considered. In the case of highly cold worked specimens, which showed fracture or even removal of Nb(C,N), no significant effect on corrosion resistance is observed, which could be an additional indicator for a highly stable passive layer on the investigated steels CN1.1Mo and CN1.12MoNb. A similar explanation was given in the case of an FeCrMnMoN steel [32] Furthermore, compared to the PM austenites investigated in [43], the alloys analyzed in the current contribution contain larger amounts of molybdenum, which has been shown to have a superior influence on resistance to pitting corrosion (Section 4.2.1). Given the lower susceptibility of general corrosion to defects of the passive layer, the low impact of hard phases seems feasible.

5. Summary and conclusions In this study, three different high interstitial austenitic stainless steels were investigated by means of cavitation erosion as well as localized and general corrosion testing with special focus on the effects of prestraining and hard phases. The following conclusions can be drawn: 1. Work hardening induced by cold rolling enhances the mechanical properties such as hardness of high interstitial austenitic stainless steels. This effect is slightly more pronounced if niobium carbonitrides are present. 2. The very high cavitation erosion resistance of high interstitial austenitic stainless steels can even be improved by severe cold work strengthening. This is attributed to their further increased resistance to plastic deformation. 3. The presence of Nb(C,N) deteriorates CE resistance. However, if cold rolling was applied, the increasing resistance of the austenitic matrix dominated the negative effect of the hard phases. 4. Localized corrosion resistance is strongly increased by molybdenum, while neither severe cold work strengthening, nor the niobium carbonitrides induced significant reduction of pitting potential under the conditions investigated. This may be due to a highly stable passive layer induced by a high density of free electrons as well as enhanced repassivation behavior. 5. In the case of surface corrosion, the effect of molybdenum is less significant, while the influences of cold working and hard phases are not severe, probably due to lower susceptibility to localized defects of the passive layer. The high resistance to cavitation erosion and corrosion of the high interstitial austenites, especially in severely cold work strengthened condition, makes them promising candidates for application in harsh environments. Still, further potential seems feasible, e.g. application of CN1.12MoNb or a comparable alloy featuring higher amounts of hard phases under additional demand by abrasive wear may be of interest for upcoming investigations.

Acknowledgments The authors gratefully acknowledge financial support by the European Union and the State of Northrhine-Westphalia in the context of the Competence Centre for Hydraulic Fluid Flow Systems at Ruhr-Universität Bochum (reference no. IV B 3-43-02/ 2005-WFBO-011Z) as well as the Federal Ministry for Economic Affairs and Energy of Germany within the framework of the joint research project POSEIDON (reference no. 03ET1072D). Special

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thanks to Adam Kazuch, M.Sc., who performed most of the experimental work and electron microscopy.

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