Correlation between fracture mechanisms and microstructure in Mn powder metallurgy steels susceptible to intergranular failure

Author’s Accepted Manuscript Correlation between fracture mechanisms and microstructure in Mn powder metallurgy steels susceptible to intergranular failure Simon Gélinas, Ian Baïlon-Poujol, Carl Blais www.elsevier.com/locate/msea

PII: DOI: Reference:

S0921-5093(18)30807-4 https://doi.org/10.1016/j.msea.2018.06.023 MSA36578

To appear in: Materials Science & Engineering A Received date: 9 February 2018 Revised date: 23 May 2018 Accepted date: 6 June 2018 Cite this article as: Simon Gélinas, Ian Baïlon-Poujol and Carl Blais, Correlation between fracture mechanisms and microstructure in Mn powder metallurgy steels susceptible to intergranular failure, Materials Science & Engineering A, https://doi.org/10.1016/j.msea.2018.06.023 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

Correlation between fracture mechanisms and microstructure in Mn powder metallurgy steels susceptible to intergranular failure

Simon Gélinasa, Ian Baïlon-Poujolb, Carl Blaisa

a) Mining, Metallurgical and Materials Engineering, Université Laval, 1065 de la Medicine, Local 1728, Quebec City, QC, Canada, G1V 0A6. b) (At the time of completion of the work) Rio Tinto Metal Powders, 1655 Route Marie Victorin, Sorel-Tracy, QC, Canada, J3R 4R4 Simon Gélinas : [email protected] Ian Baïlon-Poujol : [email protected] Carl Blais (Corresponding author) : [email protected] ; (+1) 418-656-2049

1

Abstract

This study was undertaken to thoroughly characterize the relationship that exists between microstructure, strength and fracture mechanisms of Fe-Mn-C powder metallurgy steels. Indeed, up to now the peculiar fracture behaviour of PM components made from this metallurgical system was not clearly understood by the PM community. A design of experiments (DOE) approach was used to determine the effect of chemical composition on tensile properties of Fe-Mn-C PM steels components and therefore on the fracture mechanisms leading to failure. Quantitative data describing fracture mechanisms as a function of sintered microstructure were obtained and compared against tensile strength measurements. Results led to the identification of microstructural conditions for which Fe-Mn-C PM steels are susceptible to the detrimental effect of Mn-oxide inclusions. It was found that the presence of martensite having a hardness higher than 661 HV exacerbates the detrimental effect of inclusions present in sinter necks, which explains the decreasing tensile strength observed with increasing carbon concentration. The beneficial effect of the divorced eutectoid phase as a secondary microstructural constituent is also discussed. Knowledge generated regarding the fracture behaviour of components made from this family of PM alloys allowed for the optimization of the volume fraction of microstructural constituents to maximize tensile properties. This work provides information that was missing up to now regarding the capabilities and limitations of admixed Mn PM steels in terms of tensile properties and allows the formulation of design guidelines to alleviate the adverse effects of Mn oxidation.

2

Keywords

Embrittlement, intergranular failure, microstructure, mechanical behaviour, powder metallurgy Mn steels.

1. Introduction

Manganese in ferrous powder metallurgy (PM) has been studied for quite some time due to its interesting combination of low cost and beneficial effect on improving mechanical properties of iron and steels. As a matter of fact, the cost of Mn has always been lower than that of the more common alloying elements used in the PM steel industry, such as nickel and copper [1]. Additionally, the solution strengthening effect of manganese as well as its significant influence on increasing hardenability makes it an obvious choice for alloying PM ferrous materials [2, 3]. As already stated in literature, these characteristics clearly identify Mn as a prime candidate for the lean PM approach, where total concentration of alloying elements and cost should be minimized while leaving strength unaffected, or ideally improved [4-7]. However, despite being a primordial element in wrought steel metallurgy, inherent characteristics of manganese are preventing its utilization in high or medium manganese PM steels. As widely reported, manganese belongs to a group of chemical elements quite sensitive to oxidation, along with chromium and silicon [3, 8-13]. This means that special attention must be given to atmosphere purity (dew-point) and reduction potential as well as sintering process design (heating rate and final sintering temperature [13]) to 3

efficiently sinter admixed PM steels containing these elements. To successfully carry their experiments, some authors relied on the most straight forward solution of using atmospheres containing large volume fractions of hydrogen (25 vol.% or more) [10, 14, 15] and even pure hydrogen [16]. Additionally, multiple approaches (from the early use of a getter powder [11] or the semi-closed container technique [14, 17] to a combination of both [18]) have been investigated to circumvent the need for high purity atmospheres and/or utilization of large-volume fractions of hydrogen. These efforts were made to facilitate industrial sintering operations and prevent increases in production costs related to the use of admixed manganese The motivation behind the use of alternative techniques to enhance atmosphere purity (or create a micro-atmosphere with higher reduction potential) is the prevention of oxide formation in sintered necks [8, 19]. These oxide inclusions, present even though atmosphere dew point requirements are met according to thermodynamic equilibrium, were identified as the result of manganese in vapour phase reducing some of the iron oxide intrinsically present at the surface of particles form the base powder. Working with a master alloy containing Mn as the main alloying element, Oro et al. correlated the brittle decohesion failure around former master alloy particles to the weakening effect of Mn-Si oxide inclusions in grain boundaries [13] and described the concept of "internal getter effect" associated with the use of Mn or other elements forming more stable oxides than Fe (such as Cr, Mn and Si) [13, 20]. Strength deterioration of sinter necks by the presence of oxide inclusions is therefore to be expected when working with admixed manganese, regardless of the Mn carrier, due to the inherent oxygen transfer from the Fe base powder residual oxide layer to the more oxygen avid alloying element [8]. That 4

alone would be a highly probable strength limiting factor in those PM steels. However, upon review of the pertinent literature, some peculiarities were noted when analyzing the conditions leading to maximum strength in admixed PM steels containing electrolytic Mn as well as low carbon and high carbon ferromanganese. In some cases, steels seemingly free of oxides inclusions fail in a brittle manner [10, 15, 17, 21]. Therefore, it is believed that for this category of PM steel, the relationship between microstructure and strength is affected by more than solely the presence of detrimental Mn-rich oxides. This work aims at establishing the effect of microstructure modifications due to changes in chemical composition on the fracture behaviour and ultimate tensile strength of admixed Mn lean PM steels subjected to grain boundary contamination by Mn oxidation. Maximum strength of PM steels containing admixed Mn additives

Literature review reveals that PM steels containing admixed Mn, regardless of Mn carrier (electrolytic Mn, low carbon and high carbon ferromanganese) may show premature strength diminution that could not be linked only to the formation Mn oxide inclusions. Among the first extensive research work done on the use of manganese in PM steels, Zapf (1975) investigated the effect of pressing and sintering parameters, as well as chemical composition, on the mechanical properties of a water atomized Fe powder admixed with either electrolytic Mn (particle size < 5µm) or a Fe-85Mn ferroalloy (<63µm) [11]. Fe-xMn alloys and Fe-2Mn-xC steels sintered at 1280°C in dissociated ammonia with a Fe-8Al getter powder present a decrease in tensile strength when the weight fraction of respective alloying elements reaches threshold values: Fe-6Mn alloy and Fe-2Mn-0.8C steel both correspond to an optimum chemical composition characterized by a similar tensile strength of 550±50 MPa. This behaviour was attributed 5

by the author to the increasing presence of martensite in the microstructure that is believed to cause embrittlement. More recently, Salak et al. investigated the effect of admixed manganese content on the mechanical properties of a 3 wt.% Cr and 0.5 wt.% Mo pre-alloyed steel (Astaloy CrM) using different Mn carriers: electrolytic Mn, a medium carbon FerroMn and a high carbon FerroMn [17]. The authors reported a maximum in tensile strength at 2 wt.% Mn for a carbon content of 0.24 wt.% for the assintered and sinter-hardened specimens. The authors concluded that the decrease in strength at 3 wt.% Mn was related to the increasing proportion of martensite having a high micro-hardness (around 900 HV). The results of both Zapf and Salak revealed the presence of a concentration threshold in Fe-Mn-C admixed PM steels beyond which strength decreases, seemingly non-related to the formation of detrimental oxide inclusions. It is worth noting that, in both studies, optimum chemical concentration is independent of the Mn carrier and that only maximum strength differs. Additionally, the findings or the data presented by multiple other authors also point out to the existence of a concentration threshold not only for Mn, but for carbon content as well. Table 1 presents a summary of selected relevant works focussing on the mechanical properties of Fe-Mn-C admixed steels where strength diminution was observed but could not be linked only to the formation Mn oxide inclusions.

6

Table 1 – Selected work highlighting strength deterioration as a function of chemical composition in Fe-Mn-C admixed PM alloys Materials Ref. Mn carrier

Base powder / Additives

[15]

Highcarbon ferroMn (Fe-73Mn6.3C-0.7Si1.3O)

Processing conditions Optimum conditions

Remarks on strength deterioration

Presence of oxide inclusions

Increasing C and Mn decreased TRS Sintering at 1180°C (-60°C) did not compensate for TRS loss due to incr. C

Present for the highest dew point (-30°C)

Sintering

Atmosphere

Höganäs sponge iron (NC 100.24) + Astaloy Mo

1120°C and 1180°C for 1h

Pure H2 (-30°C) 75H2 – 25N2 (-60°C)

3Mn-0.6C sintered at 1120°C (30°C) Lowest C and Mn

[10]

Med.carbon ferroMn (Fe-71Mn1.4C-1.3O)

Höganäs sponge iron (SC100.26)

1180°C for 40 min (furnace cooled and sinterhardened)

25H2 – 75N2 (-42°C) in semi-closed containers with getter powder

4Mn-0.3C sinterhardened Lowest C and highest Mn

[21]

Unspecified Mn source

Unspecified Hoeganaes iron powder + Admixed Mo

1120°C for 15 min and different cooling rates

10H2 – 90N2

Maximum TRS at 2.4Mn0.4C

[6]

Highcarbon ferroMn (Fe-75Mn7C-1.25Si)

ATOMET 4001 + Admixed ferroCr

1120°C, 1170°C and 1200°C for 30 min

10H2 – 90N2

High Mn and intermediate C content.

At 3 wt.% Mn higher C resulted in higher TS. At 4wt.% Mn TS decreases (except for 4Mn-0.3C sinterhardened) and higher C induced more important decrease. Increasing C content to 0.6wt.% reduce optimal Mn content to 1.75wt.% Higher cooling rate reduces C content threshold Intermediate C content (range from 0.3 to 0.7wt.%) result in higher strength Mn content must be maximized (up to 1.2wt.%)

Not systematically detected

Not discussed.

Not discussed.

2. Methodology

In order to study the relationship that exists between microstructure, strength and fracture mechanisms of Fe-Mn-C powder metallurgy steels, a high carbon ferromanganese was selected as the Mn carrier. Although a different source of Mn could have been used, the similarities with the one used in [6] made it possible to use this data for comparisons purposes and to identify regions of interest in terms of chemical composition and processing variables where strength deterioration is observed. Subsequently, a low alloy iron base powder and a water atomized carbon masteralloy [22] were also selected to match the materials used in [6]. Synthetic graphite was used to adjust the concentration of 7

carbon without causing variation in the Si content. Chemical composition and characteristics of the aforementioned materials are presented in Table 2. Table 2 – Chemical composition of the materials used in the preparation of premixes Material

Supplier

Particle size

High-carbon ferroMn

Hengyuan Metal & Alloys Powders Ltd.

ATOMET 4001

Chemical composition (wt.%) Mn

Cr

C

Mo

Si

Milled D50 = 9.2 µm

75 %

---

7%

---

1.25 %

Rio Tinto Metal Powders

D50=78 µm

0.15%

0.05%

---

0.50%

0.01%

Fe-C-Si master alloy

Rio Tinto Metal Powders

D50=87 µm

---

---

2%

---

1%

KS6 synthetic graphite

IMERYS Graphite & Carbon

D50=3.4 µm

---

---

100%

---

---

A factorial DOE having Mn and C content as factors was formulated around the values identified in data from ref. [6] to provide specimens destined to undergo a thorough characterization of fracture surface, microstructure and tensile strength. The simplicity of 2² DOE used in this work, with the addition of center points and extra points (Table 3), allowed to perform a complete regression analysis (significance of main factors, interactions and second order terms) of quantified characteristics in terms of chemical composition. Such mapping of the properties and characteristics, coupled with an exhaustive analysis of the variations and relationships between them, allows the optimization of tensile properties of those PM steel. Moreover, it reveals the effect of the microstructural changes that are responsible for intergranular fracture and eventually to a premature decrease in strength.

Table 3 – Targeted composition for the factorial 2² DOE 8

DOE Run

Manganese (wt.%)

Carbon (wt.%)

Molybdenum (wt.%)

Silicon (wt.%)

1

1.25

0.65

0.40

0.20

2

1.25

0.55

0.40

0.20

3

0.80

0.65

0.40

0.20

4

0.80

0.55

0.40

0.20

Center-point

1.03

0.60

0.40

0.20

5

1.90

0.70

0.40

0.20

6

1.90

0.50

0.40

0.20

Sintering of standard tensile test specimens (MPIF Standard 10 [23]) pressed to a green density of 7.05 g/cm³ was done in a laboratory scale continuous-belt furnace equipped with a forced-convection cooling unit (FCCU). The sintering temperature was 1200°C and specimens were sintered for 30 minutes under an atmosphere of 10 vol.% H2 – 90 vol.% N2. The FCCU was set to obtain an average cooling rate of 2.2°C/s between 600°C and 350°C. Independently of the sintered microstructure, all specimens were tempered at 200°C for 60 minutes. Prior to tensile testing and characterization, chemical composition of sintered specimens was confirmed using ICP-MS analysis for transition metals (Fe, Mn, Si and Mo) and combustion analysis for combined carbon content. For each DOE run, a minimum of 4 specimens were submitted to tensile tests, conducted according to MPIF Standard 10 [23]. Following the confirmation of ultimate tensile strength decrease with increasing content in alloying elements, fracture surfaces and microstructure were characterized using an image analyse routine to obtain quantitative values. To quantify fractographic features resulting from the prevailing rupture modes, fracture surfaces from tensile specimens were characterized in scanning electron microscopy (SEM) using a JEOL 840-A equipped with a PGT Avalon (606K-3GPS) energy dispersive X-Ray spectrometer (EDS). A series of micrographs was acquired at a 9

predetermined magnification, corresponding to an average surface of 82 000µm² per specimen. Image analysis was carried out using Matlab Image Analysis Toolbox to discriminate fractographic features according to three distinct fracture morphologies (illustrated in Figure 1 – A). Apparent surface fraction of each of the aforementioned features was then measured [24]. Note the use of the term apparent surface fraction, since surface roughness is not taken into account in these 2D observations. A similar quantitative approach using image analysis was applied for the characterization of microstructural constituents in optical microscopy (Olympus GX51 with Clemex Vision PE 8.0 software). Volume fractions of each main constituent (Figure 1 – B) were measured on specimens etched with Nital 2%. A surface area of 423 000 µm² was characterized for each specimen. Additionally, Vickers micro-hardness measurements were performed on the different microstructural components (dwell time of 13 seconds with an applied load of 50 gf) to provide an indication of the change in local mechanical properties as a function of Mn and C content. Measurements were performed with a Matsuzawa MMT-X7A micro hardness tester controlled with Clemex CMT.HD v7 software. Linear regression was performed based on the minimization of a least-square criterion via an iterative approach. Goodness of fit was evaluated for each parameter addition based on a two-sided Student's t-test with α=0.05 significance level. A factor that barely failed the t-test could be included in the final models based on theoretical considerations. Global model prediction adequacy was evaluated using the square of the multiple correlation coefficients (R2). Additionally, models presented in this paper were all deemed 10

significant based on an F-test with α=0.05 significance level and summarized by the test p-value [25].

Figure 1 – A) Typical fracture surface features and B) principal microstructural constituents measured by micrograph segmentation procedure using image analysis.

3. Results

Chemical analysis of sintered specimen confirmed that targeted chemical compositions were attained for each DOE run. Tensile test results presented in Table 4, specifically elongation values, indicate that the more highly alloyed specimens are subject to fragile rupture. Moreover, this diminution of ductility is accompanied by a decrease in strength. SEM characterization of fracture surfaces revealed a mixture of cupules, cleavage (river patterns) and intergranular rupture (grain boundary decohesion) present in varying amounts for each DOE run. In can be seen from the values presented in Table 4 that the deterioration of strength and ductility corresponds to an increase in grain boundary decohesion (while ductile rupture components or cupules are decreasing). Fragile rupture is therefore favoured by increasing Mn or C content.

11

Table 4 – Results of the tensile tests of Fe-Mn-C alloys (errors are 95% confidence intervals) and apparent fraction of fracture surfaces constituents measured from SEM micrographs. Intergranular Cleavage Dimples (app. fract.)* (app. fract.)* (app. fract.)*

DOE Run

TS (MPa)

YS (MPa)

Elong. (%)

1

642 ± 16

564 ± 27

0.76 ± 0.05

0.20

0.04

0.12

2

735 ± 25

587 ± 8

1.02 ± 0.14

0.12

0.08

0.20

3

681 ± 20

513 ± 18

1.35 ± 0.18

0.09

0.10

0.22

4

612 ± 7

453 ± 13

1.73 ± 0.34

0.05

0.05

0.26

Center-point

657 ± 21

517 ± 18

1.05 ± 0.09

0.12

0.09

0.17

5

519 ± 20

519 ± 20

0.38 ± 0.06

0.32

0.01

0.10

6

720 ± 13

684 ± 24

0.70 ± 0.02

0.16

0.03

0.15

*±0.02 estimated error based on replicate measurements (95% confidence)

Regression analysis of the tensile properties (Table 5) confirms that both Mn wt.% and C wt.% significantly affect tensile and yield strength, as expected. Given the presence MnC interaction and Mn2 terms, when increasing the content of alloying elements over a given chemical composition, a decrease in strength occurs. This behaviour follows the trend previously identified in literature for many admixed Mn PM steels. Moreover, the interaction between the two main factors (Mn·C term) introduces a variation in the optimum chemical composition for maximization of tensile or yield strength. Regarding the state of the fracture surfaces, mainly cupules and intergranular fracture apparent fraction, Mn wt.% and C wt.% or their interactions are significant factors (Table 5). The result of the regression analysis confirms that with an increase of the content in alloying element comes a decrease in ductile rupture and an increase in fragile rupture by grain boundary decohesion, which eventually leads to a decrease of maximum strength. In the case of the apparent fraction of river patterns, the data available did not allow us to conclude on the significance of one or both factors at 95% confidence level. The model is 12

still presented since the combination term of Mn wt.% and C wt.% was barely excluded based on Student's t-test, but will not be further discussed Table 5 – Regression models describing mechanical properties and fracture behaviour Property

Model

R2

P-value

Tensile strength (MPa)

(

)

( )

(

)

(

)

0.86

1.51e-08

Yield strength (MPa)

(

)

( )

(

)

(

)

0.94

1.80e-12

0.85

4.35e-09

0.71

1.66e-02

0.53

6.42e-02

0.96

2.00e-03

Elongation (%) Cupules (app. fract.) River patterns (app. fract.) Intergranular (app. fract.)

(

)

( )

(

(

)²

) (

) (

)

( )

Typical microstructure for each DOE run can be seen in Figure 2. Each specimen contains a varying volume fraction of divorced eutectoid phase (a common feature of PM steels pre-alloyed with Mo, ref. Section 4.3), pearlite and bainite (not differentiated in this work) and martensite. From Figure 2, it is seen that the change from a divorced eutectoid microstructure (#2,#4,#6 - low carbon) at the center of particles of former base steel particles to finer pearlite or bainite (#1,#3,#5 - high carbon) is visible as carbon increases for a fixed Mn wt.%. Moreover, the volume fraction of martensite significantly increases with increasing Mn content: specimens from runs #3 - 4 are low alloyed and runs #5 - 6 are more highly alloyed specimens (Ref. Table 3). Similarly, decrease in strength and ductility appears to be related to the changes in eutectoid morphology and proportion of martensite with increasing content in alloying elements.

13

Figure 2 – Typical sintered microstructure for the corresponding DOE showing varying proportions of the main microstructural constituents.

Table 6 reports the results from image analysis (volume fractions of each phase) and corresponding micro-hardness. Micro-hardness measurements revealed hardness values between 673 - 749 HV for martensite, between 248 - 288 HV for the divorced eutectoid and between 315 - 354HV for the clusters of fine pearlite and bainite. As expected, a finer eutectoid structure (pearlite) possesses a higher hardness than the coarser divorced eutectoid. It is also normal to observe a higher martensite hardness for specimens with a superior C content [26]. Again, this change of local mechanical properties with chemical composition goes along the fact that admixed Mn steels containing an relatively high amount of alloying element are subject to grain boundary decohesion.

14

Table 6 – Volume fraction and Vickers micro-hardness (errors are 95% confidence intervals) of the different microstructural components found in the Fe-Mn-C alloys DOE Run

Pearlite and bainite

Divorced eutectoid

Martensite

Vol. fract.*

HV

Vol. fract.*

HV

Vol. fract.*

HV

1

0.15

354 ± 18

0.12

280 ± 14

0.59

724 ± 17

2

0.12

328 ± 13

0.22

263 ± 6

0.51

674 ± 14

3

0.30

352 ± 13

0.36

288 ± 11

0.22

700 ± 33

4

0.10

315 ± 11

0.61

258 ± 9

0.12

681 ± 26

Center-point

0.16

339 ± 12

0.37

269 ± 11

0.29

687 ± 19

5

0.07

344 ± 16

0.04

278 ± 11

0.75

749 ± 18

6

0.03

--

0.09

248 ± 11

0.74

673 ± 19

*±0.02 estimated error based on replicate measurements (95% confidence)

Regression models in Table 7 describe respectively the volume fraction and hardness of a mixture of fine pearlite and bainite, divorced eutectoid and martensite. The volume fraction of martensite is statistically dependent on Mn wt.% only and eutectoid morphology is dependent of both Mn wt.% and C wt.%, which corroborates metallographic observations discussed earlier (visible in Figure 2), It is important to note that the cooling rate was kept constant for all series of specimens and also that carbon has no significant effect on the formation of martensite in the concentrations range covered by the DOE. Regression analysis also confirmed that the amount of C influenced the hardness for the observed microstructures and that Mn has no significant effect.

Table 7 – Regression models of microstructure volume fractions and micro-hardness. 15

Microstructure

Model (

Pearlite and bainite

)

( )

(

( ) (

Divorced eutectoid

)

( ) (

Martensite ( )

)

(

)

)

R2

P-value

0.94

2.30e-02

0.70

3.72e-02

0.78

8.74e-03

0.86

2.51e-03

0.99

1.39e-04

0.78

8.89e-03

Finally, SEM observations of fracture surfaces reveals the presence of Mn oxides as well as combined Mn and Si oxides for all the chemical compositions considered in the DOE (typical size of inclusion visible in Figure 3). The composition of these inclusions, found in both cupules and in ruptured grain boundaries (intergranular fracture), was confirmed by EDS (Figure 3).

Figure 3 – Typical SEM micrographs of fracture surfaces showing spherical oxide in cupules (left) and in ruptured grain boundaries (right), with typical EDS X-Ray spectrum of such inclusions present in both ruptured components.

16

However, a number of intergranular fracture areas were marginally contaminated by oxides, with some of them free from any inclusions (Figure 4). This suggests that strength deterioration could not be attributed to Mn oxidation alone, although incoherent inclusions are recognized as detrimental to fracture resistance (especially in PM materials when located in sinter necks) [27].

Figure 4 – Intergranular fracture areas showing minimal contamination (Top) and no visible oxide inclusions are found in fracture surfaces throughout the DOE specimens (Bottom)

From experimental results, it is clear that changes in phase proportions and mechanical properties (i.e. micro-hardness) play an important role in determining the susceptibility to intergranular fracture brought about by the presence of Mn-rich oxides. Using the regression models established in the above section, a global portrait of those admixed Mn 17

PM steels can be laid out to interpret the observed premature decrease in strength according to modification of the microstructure and fracture behaviour.

4. Discussion 4.1 Optimum tensile strength conditions

Failure analysis of multiple PM steels revealed that pores, especially with sharp angles, act as stress raisers where cracks can nucleate more easily [27, 28]. Additional defects in the vicinity or directly in grain boundaries, such as Mn oxides inclusions, would strongly accentuate the susceptibly of those grain boundaries/pores combinations to act as the preferred fracture path [27]. Given the nature of the oxides identified, it is logical to assume that the number of these inclusions would increase with an increasing Mn concentration, resulting in more severe decrease in strength. For constant carbon content, the previous analysis revealed that this is indeed occurring (ref. Table 5). However, it also revealed that for fixed Mn content an increase in carbon concentration results in a diminution of the tensile strength. This ambivalent relationship is a strong evidence that the microstructure surrounding the defects produced by Mn oxidation is of primordial importance for explaining the complex mechanical behaviour of those Fe-Mn-C PM steels. Using the regression models describing mechanical properties, apparent surface fraction of rupture surface constituents and volume fraction of microstructural components, it is possible to analyze the behaviour of these Fe-Mn-C alloys in the composition range of 0.75 to 1.88 wt.% Mn and 0.50 to 0.67 wt.% C. As expected from literature, these PM alloys exhibit an optimum tensile strength that is function of their content in alloying 18

elements, both Mn and C, with optimal Mn content being inversely dependent of C content and vice versa.

Figure 5 – Selected models prediction as a function of increasing Mn content for fixed C concentrations (shaded areas represent 95% confidence bounds)

In Figure 5, for a fixed carbon content of 0.5 wt.%, one can see that maximum tensile strength (750 MPa) is reached at 1.56 wt.% Mn. For this given chemical composition, the microstructure contains approximately 68 vol.% of martensite having a hardness of 661HV. Increasing carbon content (Figure 5) induces a shift of maximum strength towards lower Mn content, accompanied by an increase in martensite hardness and a decrease in critical martensite volume fraction (martensite volume fraction corresponding to maximum strength for a given carbon content). The point where strength starts decreasing corresponds, independently of carbon content, to the change in dominant fracture mode, i.e. apparent fraction of intergranular fracture starts to prevail over the 19

ductile features (cupules). One of the key factors influencing maximum strength as a function of C wt.% appears to be the mechanical properties of martensite, represented here by micro-hardness, rather than martensite volume fraction itself (independent from carbon content). Hardenability of conventional wrought steel (i.e. the depth to which an alloy can transform into martensite upon rapid cooling, or quenching) is in fact known to increase with increasing content of alloying elements, whereas carbon mostly influences hardness of quenched microstructures [26]. As seen from the correlation matrix presented in Table 8, a negative linear relationship indeed exists between tensile strength and martensite micro-hardness while there is no apparent linear correlation between tensile strength and martensite volume fraction. Table 8 – Correlation matrix between selected factors and properties related to the decrease in strength (according to Pearson's product-moment correlation coefficient [29]) Inter. app. fract.

P-B vol. fract.

Eutec. vol. fract.

Mart. vol. fract.

TS (MPa)

EL (%)

Mart. HV

1

-0.32

-0.83

0.82

-0.59

-0.97

0.82

1

0.29

-0.57

0.15

0.40

0.11

1

-0.95

0.05

0.91

-0.48

1

-0.09

-0.90

0.38

1

0.41

-0.81

TS (MPa)

1

-0.67

EL (%)

1

Mart. HV

Inter. app. fract. P-B vol. fract. Eutec. vol. fract. Mart. vol. fract.

As expected, the volume fraction of microstructural constituents, beside martensite, also has an effect on tensile strength and such relationship is made evident in Figure 6. With the exception in the lower end of the Mn concentration range investigated (where martensite is not the dominant phase), the highest strength is systematically obtained at 20

low carbon content where the divorced eutectoid microstructure dominates over fine pearlite or bainite. Accordingly, correlation coefficients between measured data (ref. Table 8) indicates that a strong inverse linear relationship exists between the volume fraction of divorced eutectoid and the apparent fraction of intergranular rupture, which is in turn inversely correlated to tensile strength.

Figure 6 – Volume fractions of microstructural constituents and properties prediction as a function of C content for fixed Mn concentrations (shaded areas represent 95% confidence bounds) 4.2 Oxides and microsture contribution to diminishing strength

Martensite tempered at low temperature (200°C and below) has proven to provide high tensile strength values. On the other hand, the latter becomes highly susceptible to intergranular fracture when carbon content is above 0.5 wt.% due to increased strainhardening [30]. A large difference in strength with former austenitic grain boundaries, especially when impurities are present [27, 30], is therefore expected. Data (ref. Table 6) 21

and regression model of martensite micro-hardness (ref. Table 7) are indeed suggesting higher martensite strength at higher carbon content. In other words, the detrimental effect of Mn oxides is aggravated for higher martensite strength (at higher carbon content) since the strength difference between grains and their boundaries is exacerbated. This results in dominance of intergranular fracture (illustrated in Figure 7 – B), which translates in a decrease in strength. Figure 7 – A represents the compromise that must be reached for FeMn-C PM steels to deal with the hardly avoidable formation of oxide inclusion during sintering operations. Thus, empirical evidence suggests that low-carbon martensite forms an ideal phase that reduces the strength differences between particles or grain cores and contaminated necks or grain boundaries. To the authors’ knowledge, literature related to PM steels does not contain detailed information about characteristics and mechanical properties of the Mo-induced divorced eutectoid microstructure. However, parallels can be made with conventional metallurgy. In hypoeutectoid steels, the divorced eutectoid transformation (DET) requires the presence of undissolved carbides in austenite to occur. Alloying elements that stabilize carbides, such as Mo, retard carbide dissolution upon reaching the austenitizing temperature [26]. These undissolved carbides, or even “carbide remnants” creating nonhomogeneous areas in austenite, act as precipitation or nucleation sites. Such sites allow the growth of disconnected cementite particles directly from the austenite phase, compared to spheroidization treatments performed after the formation of lamellar pearlite or martensite [31]. However, at room temperature both microstructures are composed of cementite (carbides) particles (or small aggregates) dispersed in a ferrite matrix. It is known that conventional spheroidization annealing produce a softer and more ductile 22

microstructure than lamellar pearlite [26]. Micro-hardness measurements (ref. Table 6) and regression model (ref. Table 7) indicate that the divorced eutectoid microstructure induced by the presence of pre-alloyed Mo is indeed softer than conventional lamellar pearlite. Therefore, like its spheroidized counterpart, it possesses lower strength, but higher ductility [26, 32]. This combination of mechanical properties possibly explains the reduced susceptibility of the divorced eutectoid phase to intergranular fracture induced by grain boundary contamination, compared to harder fine lamellar pearlite, in a similar way than low micro-hardness martensite is preferable over its higher micro-hardness counterpart. In their analysis of the decrease in strength typical of high-Mn PM steels, Salak et al. also linked this behaviour to the formation of martensite (high hardness) at higher Mn content, replacing softer microstructures (pearlite and bainite) and surpassing the optimal ratio of hard over soft microstructural components [17]. In the current work, by varying the carbon concentration, it was observed that the critical volume fraction of martensite decreased with increasing carbon content, meaning that a smaller volume fraction of inclusion-susceptible martensite (i.e. high hardness martensite) is required for premature fragile failure (ref. Figure 5).

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Figure 7 – Contaminated particles necks and fracture path relationship for A) a microstructure with intermediate crack propagation resistance (i.e. low hardness martensite) and B) for a microstructure highly resistant to crack propagation (i.e. high hardness martensite).

However, considering the presence of some intergranular rupture seemingly free of oxide inclusions (Figure 4) combined with literature report of intergranular failure with no visible contamination, another source of grain boundary embrittlement is to be considered [33]. Segregation of other chemical species such as sulphur, phosphorus [34] or even oxygen and carbon recognized for having a negative effect on mechanical properties in steel is a possibility [27, 30, 33]. Moreover, the concentration gradient inherent to admixed PM materials clearly creates region with higher Mn concentration [17]. These areas are discernable through microstructure heterogeneity as seen in Figure 2 where martensite formation is clearly promoted by local high Mn concentration. Furthermore, these martensitic areas are more likely to be located around grain boundaries representing sinter necks. According to Kwon et al., a difference in strength, measured by a nanoindentation approach (room temperature), was observed between grain core and grain 24

boundaries in Fe-8Mn alloys (such high concentration of Mn is highly probable in admixed Fe-Mn-C alloys) due a difference in hardness between grain core and boundaries, making grain boundaries the preferred fracture path [35]. Moreover, they showed that carbon exacerbated this phenomenon.

Figure 8 – Summary of strength conditions in relation to alloying element content

This difference exists even though no Mn segregation in grain boundaries was detected (although carbon segregation was detected) and was leading to intergranular fracture in tensile tests at cryogenic temperatures. In light of the previous discussion, the proposed explanation for strength deterioration upon reaching concentration thresholds for the FeMn-C PM steels studied in this work is summarized in Figure 8. 5. Conclusion

To understand the complex relationship between microstructure, strength and fracture mechanism in Fe-Mn-C PM steels, a complete characterization of multiple material properties and characteristics was performed. This work combined detailed characterization of specimens with varying chemical composition (following a factorial 25

DOE) with regression analysis to provide a new comprehensive view of the peculiarities of admixed Fe-Mn-C steels made from ferromanganese additives. The main findings of our research, investigating strength decrease upon reaching a threshold chemical concentration, may be summarized as follows: 1.

Fe-Mn-C PM steels made from admixed ferromanganese are subject to grain boundary contamination (complex Mn and Si oxide inclusions were identified during fractographic observations), but the extent to which strength deterioration occurs is dictated by the microstructure surrounding these contaminated areas.

2.

Critical Mn concentration for maximum strength is shown to be a function of C content (inversely correlated) and the decrease in strength corresponds to the change of the dominant fracture mode, i.e. apparent fraction of intergranular fracture starts prevailing over the ductile features (regardless of C content).

3.

Martensite is the principal phase present in high strength specimens as well as in materials showing brittle behaviour and intergranular rupture. Micro-hardness of this phase determines if contaminated grain boundaries will act as the preferred fracture path (eventually leading to prevalence of intergranular fracture).

4.

A threshold volume fraction of martensite exists for the range of chemistries considered in this study and it decreases with increasing C. The fact that a fully martensitic microstructure is not the optimal microstructure, in terms of tensile strength, is peculiar to powder metallurgy materials and especially to this admixed Fe-Mn-C family (because of manganese oxidation and ensuing contamination). 26

5.

The results presented in this study indicate that a larger volume fraction of divorced eutectoid in the remaining microstructure (rather than pearlite or bainite) is linked to higher strength materials.

The latter conclusions allow us to provide a new insight on the possibilities and limitations of admixed Fe-Mn-C lean PM steels. A simple alloy design guideline should be observed when dealing with this material and possibly similar materials leading to oxide inclusions of comparable size in sinter necks: high Mn content to obtain the highest possible vol. fraction of martensite and the lowest C content still promoting martensite formation (assuring low hardness) in the cooling conditions found in PM processing. 6. Acknowledgements

This work was supported by the National Research Council of Canada – Network of Centers of Excellence Auto21, project C502-CPM. 7. References [1] Infomine, http://www.infomine.com/investment/ (accessed 10/01/2018). [2] G.E. Totten, Steel heat treatment metallurgy and technologies, Taylor & Francis, Boca Raton, FL, 2006. [3] A. Šalak, M. Selecká, Manganese in powder metallurgy steels, Cambridge International Science Publishing Ltd, Cambridge, UK, 2012. [4] F. Chagnon, L. Aguirre, A new approach to lean alloy PM steels, Int. Conf. Powder Metall. Part. Mater., MPIF, Chicago, IL, 2013, pp. (10) 78-87. [5] R. Oro, M. Campos, J.M. Torralba, C. Capdevila, Lean alloys in PM: From design to sintering performance, Powder Metall. 55(4) (2012) 294-301. [6] F. Chagnon, J. Campbell-Tremblay, C. Blais, Effect of Carbon, Manganese and Chromium Concentrations, Densities, Sintering Temperatures and Post-Sintering Cooling Rates on Properties of Lean Alloy PM Steels, Int. Conf. Powder Metall. Part. Mater., MPIF, Orlando, FL, 2014, pp. (10) 162-176. [7] P. Sokolowski, B. Lindsley, Leaner alloys for the PM industry, Powder Metall. 55(2) (2012) 84-87.

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