Florian Moreau

Jérôme Marrot

Frédéric Banse

Christian Serre

Antoine Tissot

Sequential installation of Fe(II) complexes in MOFs: towards the design of solvatochromic porous solids

Fe(II) coordination complexes have been anchored on robust Metal-Organic Frameworks secondary building units using a sequential installation strategy. The coordination sphere around Fe(II) has been modulated upon sorption of alcohols in the pores of the MOF, leading to a reversible colour change from orange to red. In the case of MFU-4l, a preferential adsorption of MeOH has been observed, evidencing the fact that the host matrix can be used for pre-concentrating a targeted analyte, which may be beneficial for the design of sensing devices.

Introduction

The design of volatile organic compounds (VOCs) chemosensors, i.e. solids which undergo a measurable change of their physical properties upon interaction with a targeted molecule is of great interest for environmental applications. 1,2 Porous switchable compounds are promising candidates for sensing, as their physical properties -for example their color-can be tuned upon guest sorption. Metal-Organic Frameworks (MOFs) are highly tunable porous solids, which can present outstanding sorption properties, combining high uptakes with important selectivities towards specific VOCs. [3][4][5][6] In some cases, their physical properties, such as their color, emissive or magnetic properties can evolve upon guest sorption. [7][8][9][10][11][12][13] However, switchable MOFs often lack of an adequate chemical stability for being used in real chemosensing applications. 14 In order to overcome this issue, an alternative strategy consists of loading molecular species, such as isomerizable or redox/pH responsive organic molecules or switchable coordination complexes, inside the pores of chemically stable MOFs. 15 To this end, spin crossover (SCO) complexes are promising candidates of switchable objects that can undergo a low-spin (LS) ↔ high-spin (HS) conversion upon external stimuli such as temperature, pressure, light or chemisorption, accompanied by a change of magnetic, structural and optical properties. 16 The loading of SCO complexes in porous architectures has been first studied using oxalate-based networks, in which the complexes act as templates for the framework synthesis, resulting in non-porous hybrid structures. [17][18][19] Few examples of SCO complexes loaded in presynthesized MOFs have also been described. [20][21][22] In these cases, the limited MOF stability in presence of water and/or solvents may however prevent studying the effect of solvent or vapor loading on the spin state of the guest complexes. Recently, some of us have reported the loading of Fe III (sal 2 trien) + complexes into the water stable mesoporous aluminum trimesate MIL-100(Al) (MIL stands for Matériaux de l'Institut Lavoisier) using a bottle-in-a-ship strategy. 23 A reversible switching between HS and LS states at room temperature has been evidenced upon several water sorption-desorption cycles, evidencing the potential of such hybrid solids for sensing. However, this strategy does not allow controlling the location of the guest molecules in the pores, particularly when the pore size largely exceeds the dimensions of the complex, making the switching properties of the compound barely predictable. This issue might be circumvented by anchoring, covalently or through coordination bonds, the guest complexes on the host matrix.

The post-synthetic modification of MOFs, either using organic linkers possessing free coordination sites or by anchoring coordinating moieties on secondary building units (SBUs), has been widely explored. 24,25 In addition, some spin crossover complexes can be functionalized with coordinating groups such as amines or carboxylic acids. For example, Fe II (HBPI) 2 (ClO4) 2 (HBPI = 2,6-bis(1H-pyrazol-1-yl)isonicotinic acid) is a molecular complex presenting both a spin transition close to room temperature and two free pending carboxylic acids on the linker. 26,27 This complex is a variation of Fe(BPP) 2 X 2 complexes (BPP=bispyrazolpyridine derivatives) that often present a spin crossover in solution, which can depend on the nature of the solvent, making them appealing for solvent sensing. [28][29][30][31] Therefore, taking into account the ability of Zn(II), Al(III) or Zr(IV)-based MOFs to coordinate carboxylate groups on SBUs possessing labile positions, we have anchored [Fe II (BPI)(HBPI)(ClO 4 ) 2 ] -, name thereafter 1, in water-stable MOFs presenting large enough pores to accommodate the complex, namely MFU-4l and MIL-100(Al) using a sequential installation strategy (Scheme 1). MFU-4l is a Zn II triazolate microporous framework with two types of cubic cages with similar dimensions (15.5 Å sides), their vertices being occupied by Kuratowsky type secondary building units (SBUs) Zn 5 (BTDD) 6 X 4 (X= Cl-, BTDD=bis(1H-1,2,3-triazolo [4,5-b],-[4',5'i])dibenzo [1,4]dioxin) while half of the cages exhibit 8 labile Cl ligands from the SBUs pointing towards their centers (see Figure S1). 32 The framework of MIL-100(Al) is constituted of 2 types of large cages (24 and 27 Å) with microporous windows (5.5 and 8.5 Å), the SBUs being oxocentered trimers of Al octahedra Al 3 O(H 2 O) 2 (OH) with a central µ3-oxo, 6 bridging carboxylates, and terminal labile H 2 O or OH -ligands pointing towards the center of windows (see Figure S2). 33,34 The color change of the functionalized 1MOF compounds dispersed in different solvents at room temperature has then been investigated, which is a preliminary step toward the sensing of VOCs in gas phase. Scheme 1. Sequential installation strategy for pore functionalisation with anchored 1 complexes. 1) BPI.Li, MeOH, RT, 16-36h, 2) Fe(ClO 4 ) 2 .xH 2 O, HBPI, EtOH, RT, 16h.

Results and discussion

 Synthesis and characterization of spin crossover functionalized MOFs

The two MOFs selected in this study present large pores (>15 Å) that allow the sequential installation of coordination complexes 1 within their frameworks while retaining porosity. They also possess labile positions on their SBUs and in both cases, the SBU-terminal -X groups can be substituted by anionic species such as carboxylates. For the synthesis of 1MOF hybrid solids, MOFs were first soaked in a methanolic solution containing BPI.Li. After subsequent washing with MeOH, the amount of anchored BPI-was evaluated using 1 H NMR of the digested solid, indicating an average loading of 0.6 BPI -linkers per SBU for MIL-100(Al) (Al 3 O(OH) 0.4 (H 2 O) 2 (BTC) 2 (BPI) 0.6 ) and 2 BPI -linkers for MFU-4l (Zn 5 (BTDD) 3 Cl 2 (BPI) 2 ) (See ESI). Concomitantly, the amount of chlorine in MFU-4l decreases according to EDS, which indicates that the Cl -anions coordinated to the SBU are replaced by BPI -. In a second step, the solids were redispersed in EtOH in which Fe II (ClO 4 ) 2 and HBPI were sequentially added. The addition of Fe II and HBPI in the dispersion was accompanied with a gradual color change of the solid from white/clear brown to deep red, indicating the formation of [(BPI)Fe(HBPI)] + -type complexes. Finally, the red solids were washed with EtOH and acetone and dried under vacuum at room temperature.

The

formula of the loaded solids is [Al 2.7 Fe 0.3 O(OH) 0.4 (H 2 O) 2 (BTC) 2 ][(BPI) 0.6 Fe 0.6 (BPI) 0.5 (ClO 4 ) 1.2 ]
for 1MIL-100(Al) and Zn 5 (BTDD) 3 Cl 2 [(BPI) 2 Fe 2 (HBPI) 1.5 (ClO 4 ) 4 ] for 1MFU-4l according to 1 H NMR, EDS and TGA analysis (See ESI), which is in good agreement with the amount of linker anchored on the first step. The electroneutrality of the compounds is ensured by the loading of perchlorate anions in the solids. In the case of 1MIL-100(Al), the Fe/Al ratio measured by EDS exceeds the one expected from the loading of the SCO complex into the pores and results probably from a partial substitution of Al III by Fe III in the framework during the formation of the complex (See ESI).

The host MOFs exhibited PXRD patterns and N 2 sorption isotherms in agreement with the reported ones. 32,34 PXRD measurements on the loaded solids indicated that their crystallinity was not affected by the loading process (See Figure 1). Moreover, N 2 sorption isotherms at 77 K evidenced that a partial porosity was retained on the loaded MOFs, with BET surface areas decreasing from 1670 to 1050 m 2 g -1 for MIL-100(Al) and 2730 to 650 m 2 g -1 for MFU-4l (See Figure 2). The much larger drop in porosity for MFU-4l compared to MIL-100(Al), despite similar densities of complex 1 loaded in their pores (0.65 and 0.68 complexes/nm 3 for 1MIL-100(Al) and 1MFU-4l, respectively) can be attributed to differences in cages systems. Indeed in 1MFU-4l half of the -Cl groups from the SBUs have been replaced by anchored 1 complexes thus filling half of the pores of the framework and blocking the communication between the empty pores, whereas in MIL-100(Al), such pore blocking shall occur to a much lesser extent due to the framework topology. In addition to the decrease of sorption capacity, pore size distributions calculated by DFT evidenced in both cases a decrease of the average pore size during the loading process, which is consistent with a partial pore filling by coordination complexes (See Figures S5-6). The sequential installation of the complex 1 in the pores of the MOFs has further been evidenced by vibrational spectroscopy. Infrared spectra of the loaded MOFs unambiguously presented typical vibrations of the molecular complex 1 at 1705, 1625, 1573, 1526, 1404 cm -1 and of ClO 4 -anions at 1080 cm -1 (see Figure 3 and ESI). In particular, the vibration at 1625 cm -1 corresponds to the ν(C=N) vibration band of the HBPI linker coordinated to Fe(II) (ν(C=N)~1610 cm -1 for free HBPI/BPI-). The relative intensity of the vibration band at 1705 cm -1 , which corresponds to the vibration of the ν(COOH) group in 1, was less intense in 1MIL-100(Al) and 1MFU-4l than in Fe(HBPI) 2 (ClO 4 ) 2 . It confirms the partial deprotonation of the HBPI linker in 1MOF, in line with the first step of the installation strategy where BPI-is anchored to the MOFs inorganic building unit. During the washing/drying steps, the color of the samples changed from red to orange depending on the solvation state of the compound (acetone or EtOH), which is a first indication that the spin state and/or the coordination sphere of the loaded complexes can be modulated by solvent sorption. Therefore, UV-Visible spectra of the solids dispersed in several solvents have been measured. 1MOFs compounds exhibited an absorption band located at 450 nm in most solvents, except in MeOH or EtOH where the transition is located at 500-520 nm (see Figure 4). This difference may be explained by either (1) a spin change, as the MLCT transitions of Fe(II) complexes are often shifted towards lower energy upon HS-LS transition or (2) a solvation effect or (3) a change of the coordination environment around Fe(II).

In order to discriminate between these hypotheses, temperature dependent UV-Vis spectra of Fe II (HBPI) 2 (ClO 4 ) 2 in acetone have been measured (see Figure S10). A gradual thermal spin crossover centered around 285 K was observed, which is in line with previous results from the literature. [START_REF] Cook | A Unified Treatment of the Relationship Between Ligand Substituents and Spin State in a Family of Iron(II) Complexes[END_REF] Upon the HS-LS conversion, the Fe(II) MLCT absorption band gained in intensity but its position remained almost unchanged (459 nm at 320 K vs. 464 nm at 210 K). 29 Therefore, the spin crossover phenomenon cannot explain the red shift of the absorption observed in alcohols. This is further supported by magnetization measurements performed on 1MIL-100(Al) and 1MFU-4l dispersed in EtOH that confirms that Fe(II) is HS in both compounds (see Figure S22). In order to check the effect of solvation on the color of [Fe(HBPI) 2 ] 2+ , we prepared these complexes in several solvents by mixing Fe(ClO 4 ) 2 with a large excess of HBPI. In all cases, orange solutions were obtained with a maximum of the MLCT absorption peak located around 460 nm. Large differences in intensities were observed, indicating that the spin state of the complexes might depend on the solvent at RT (see Figure S11). However, this experiment evidenced that the red species observed in the MOF are not likely to correspond to [Fe(HBPI) 2 ] 2+ complexes where Fe II is surrounded by 6 N atoms from the linker in a pseudo-octahedral geometry. Then, we have performed titration experiments by adding progressive amount of 1-BPP or HBPI in a solution of Fe II in acetone or EtOH. With 1-BPP, a progressive increase in the MLCT absorption band intensity is noticed upon addition of the linker, without any shift of the band position. In acetone, a plateau is obtained after the addition of 2 equivalents of 1-BPP, suggesting a quantitative formation of Fe(1-BPP) 2 2+ species (see Figure S12). On the opposite, in EtOH the intensity increases progressively and no saturation is reached after the addition of 4 equivalents of 1-BPP, indicating that the coordination of 1-BPP to Fe II is less favored, probably due to the higher complexing ability of EtOH compared to acetone (see Figure S13). Using HBPI and acetone as solvent, a similar trend is observed, corresponding to quantitative formation of [Fe(HBPI) 2 ] 2+ complexes without any detectable intermediate species (see Figure S14). On the opposite, when using HBPI and EtOH as solvent, a red compound quickly precipitated for molar ratios HBPI/Fe II close to one, as evidenced by the baseline jump in the UV-Visible spectra (see Figure S15). With more diluted solution, no precipitation occurred but a mixture between two species with maximum absorption at 460 nm and 500 nm respectively was observed, depending on the concentration and the molar ratio HBPI/Fe II (see Figure S16). At high dilution and with a large excess of HBPI, a red complex absorbing at 500 nm was quantitatively obtained (see Figure S17). As its absorption spectrum well corresponds to the one of 1MOFs compounds dispersed in alcohols, we can hypothesize that the same species were obtained in both experiments. In order to identify it, the solid was crystallized by slow evaporation of a methanolic solution containing Fe II and HBPI. Dark red crystals were obtained and their structure was determined by single crystal XRD. In this compound, Fe(II) presents a coordination sphere that is different from [Fe II (HBPI) 2 ](ClO 4 ) 2 : each Fe center is coordinated by 3 N atoms from one BPI -linker, one COO -group from a second BPI -linker and 2 MeOH molecules, giving rise to a 1D coordination polymer (2, see ESI). Solid state UV-Vis spectra collected with an integration sphere for [Fe II (1-BPP) 2 ](ClO 4 ) 2 and 2 evidenced that the FeN 6 coordination sphere of HS [Fe II (1-BPP) 2 ](ClO 4 ) 2 corresponds to a transition centered at 430 nm, while the FeN 3 O 4 coordination sphere of 2 corresponds to a transition centered at 510 nm (see Figure S18). In addition, Nikovskiy et al. recently reported a molecular structure presenting a similar coordination sphere, which is also associated to a red color while Fe(II) is in the HS state. 36 In summary, the color change observed upon sorption of alcohols in the MOFs is probably not related to a spin state change or an effect of solvation without modification of the Fe II coordination sphere but is likely to be associated to a coordination switch of one BPI ligand, giving rise to a red HS Fe(II) complex where one BPI linker is coordinated to the Fe II through its 3 N atoms and the other one through the carboxylic acid group, while alcohol molecules can complete the coordination sphere. For all the solvents tested, the solutions remained perfectly clear after centrifugation, evidencing that no leaching of the metal complex occurred during the dispersion process. Moreover, the MOF crystallinity was not affected by the dispersion in solvents. The samples that were first dispersed in alcohols did recover the orange color after drying and redispersing in acetone, indicating that the switching process is overall reversible (see Figure S19). However, a decrease of the MLCT band intensity is observed after several dispersion/drying cycles (ca. 10% per cycle), probably due to the partial leaching of HBPI when the solids are dispersed in alcohols.

UV-Vis spectra have then been recorded using acetone-alcohol mixtures. For these experiments, the sample was stirred for few minutes before the measurements to allow the diffusion of the solvent molecules in the pores up to a stationary state (see Figure S20). Different behaviors could be observed depending on the MOF-alcohol combination. In 1MIL-100(Al), a gradual evolution between the states corresponding to the solid dispersed in pure alcohol and pure acetone was observed (see Figure S21). In contrast, in the case of 1MFU-4l dispersed in MeOH/acetone mixtures, the transition was abrupt and occurred at a composition corresponding to 20% of MeOH in acetone (see Figure 5). This observation suggests that MFU-4l is likely to adsorb preferentially MeOH compared to acetone while in the other cases, the composition of the adsorbed solvent is close to the one of the dispersant.

When considering the water vapor sorption properties of MFU-4l and MIL-100(Al), it is clear from previous studies that MFU-4l is a more hydrophobic material at low humidity level. 37,38 However, the steepness of the water sorption isotherm observed for MFU-4l indicates that the appearance of an Hbond network in the pores is favored once guest molecules reach binding sites as partial pressure increases, possibly thanks to the oxygen atoms on the dioxin core of the linker that can act as H-Bond acceptors. In such a case, methanol, due to its small size and its H-bond donor nature, might be preferentially adsorbed in the pores of MFU-4l at the expense of acetone. On the contrary, MIL-100(Al) exhibits some coordinated water molecules grafted on Al(III) octahedra from the framework. One could suspect therefore acetone to be more easily bound to the Aluminium Lewis acid sites, through its C=O bond, than methanol. 39 These hypotheses will however need to be validated by additional experiments beyond the scope of this study.

Conclusions

In conclusion, this article demonstrates for the first time that Fe(II) coordination complexes can be anchored on robust MOF SBUs using a sequential installation strategy. The coordination sphere around Fe(II) could be modulated upon sorption of alcohols in the pores of the MOF, leading to a reversible colour change from orange to red, while the anchoring of the complex with coordination bonds ensured that no leaching occurred in the dispersion. In the case of MFU-4l, a preferential adsorption of MeOH compared to acetone has been observed, evidencing the fact that the host matrix can be beneficial for pre-concentrating a targeted analyte. The application of this strategy to additional SCO complexes/MOFs compositions is underway, aiming at getting a more systematic understanding on how the properties of the guest complexes can be modulated by solvent sorption.

In the following step, the switching properties of these solids in presence of VOCs vapors will be investigated. UV-Vis: The optical spectra on solids and suspensions have been measured with a Varian Cary 300 Bio UV-Vis spectrometer equipped with an integration sphere in diffuse reflectance mode. In the case of the suspensions, concentration of ca. 1mg of solid/mL has been used. VT measurements in solution were performed with a Varian Cary 60 spectrophotometer equipped with a Hellma immersion probe (1 cm optical pathlength) and a fiber-optic cable. Low temperature was maintained using a Thermo Haake CT90L cryostat.

TGA: TGA were performed on Mettler Toledo TGA/DSC 1, STAR®System apparatus under N 2 atmosphere, at a heating rate of 5 °C.min -1 up to 600 °C.

Materials

All reactants were purchased from commercial suppliers and used without further purification.

Preparation of MFU-4l. MFU-4l was prepared by scaling up a previously reported procedure. 40

Conflicts of interest

There are no conflicts to declare.

Figure 1 :Figure 2 :Figure 3 :

 123 Figure 1: X-ray powder diffraction patterns (Cu Kα radiation) of MFU-4l and 1MFU-4l (left) and of MIL-100(Al) and 1MIL-100(Al) (right)

Figure 4 :Figure 5 :

 45 Figure 4: UV-Vis spectra at room temperature of 1MIL-100(Al); inset: pictures of 1MIL-100(Al) dispersed in acetone (orange) and EtOH (red) (left) and 1MFU-4l (right) dispersed in different solvents; inset: pictures of 1MFU-4l dispersed in acetone (orange) and MeOH (red)

  Considering the increasing number of highly porous water stable diamagnetic MOFs bearing open metal sites, and the diversity of SCO complexes, this creates opportunities for the design of a new generation of VOCs sensing devices.

	Experimental
	PXRD: Powder X-ray Diffraction data were recorded on a high-throughput Bruker D8 Advance
	diffractometer working on transmission mode and equipped with a focusing Göbel mirror producing
	CuKα radiation (λ = 1.5418 Å) and a LynxEye detector.
	SC-XRD: Single crystal X-ray Diffraction data were recorded on a four-circle kappa-axis Bruker D8
	Venture diffractometer equipped with a Mo microsource and photon III detector.
	N 2 adsorption: N 2 sorption measurements were performed at 77 K on Micromeretics Tristar or Triflex
	apparatus. The samples (ca. 40 mg) were activated by heating the powder at 120 °C (MIL-100(Al)) or
	170°C (MFU-4l) under secondary vacuum for 12 h.
	SEM-EDS: SEM images were recorded with FEI Magellan 400 scanning electron microscope equipped
	with and Oxford EDS probe.

  BPI.Li, the lithium salt of HBPI, was prepared by refluxing a solution of HBPI (2.0 g, 7.83 mmol) and LiOH (188 mg, 7.83 mmol) in MeOH (500 mL) for 1 h, followed by evaporation of the solvent under reduced pressure, recrystallization from EtOH/petroleum ether 60-40, filtration and drying in air. BPIMFU-4l. Dried MFU-4l (0.5 g, 0.4 mmol) was suspended in 100 mL of MeOH and a solution of BPI.Li (210 mg, 0.8 mmol) in MeOH (100 mL) was slowly added over 1 day. The resulting mixture was stirred at room temperature for 16 h. The solid was isolated by filtration, washed with MeOH (3*50 mL) and dried in air. BPI loading was determined by liquid 1 H NMR of the acid digested functionalized MOF (see SI) and confirmed by EDS and TGA. Formula: Zn 5 (BTDD) 3 (BPI) 2 Cl 2 BPIMIL-100(Al). Hydrated MIL-100(Al) (0.2 g, 0.26 mmol) and BPI.Li (0.2 g, 0.77 mmol) were added to EtOH (100 mL) and the resulting mixture was refluxed for 16 h. The solid was isolated by filtration, washed with EtOH (3*50 mL) and dried in air. BPI loading was determined by liquid 1 H NMR of the acid digested functionalized MOF (see SI) and confirmed by EDS and TGA. Formula: Al 3 O(OH) 0.4 (H2O) 2 (BTC) 2 (BPI) 0.6 1MFU-4l. MFU-4l_BPI (50 mg, 0.03 mmol) and Fe(ClO 4 ) 2 .xH 2 O (23 mg) were added to EtOH (10 mL) and the suspension was stirred for 30 min, then a solution of BPI (12 mg, 0.05 mmol) in EtOH (10 mL) was added and the suspension was stirred for 16 h. The solid was isolated by centrifugation, then washed with acetone (2*20 mL) and dried in air. BPI and Fe loadings were determined by liquid 1 H NMR of the acid digested functionalized MOF (see SI) and confirmed by EDS and TGA. Formula: Zn 5 (BTDD) 3 Cl 2 [(BPI) 2 Fe 2 (HBPI) 1.5 (ClO 4 ) MIL-100(Al)_BPI (50 mg, 0.08 mmol) and Fe(ClO 4 ) 2 .xH 2 O (20 mg) were added to EtOH (10 mL) and the suspension was stirred for 30 min, then a solution of BPI (12 mg, 0.05 mmol) in EtOH (10 mL) was added and the suspension was stirred for 16 h. The solid was isolated by centrifugation, then washed with acetone (2*20 mL) and dried in air. BPI and Fe loadings were determined by liquid 1 H NMR of the digested functionalized MOF (see SI) and confirmed by EDS and TGA. Formula: Al 2.67 Fe 0.33 O(OH) 0.4 (H 2 O) 2 (BTC) 2 [(BPI) 0.6 Fe 0.6 (HBPI) 0.51 (ClO 4 ) 1.2 ]

	BTDD
	(1g, 3.76 mmol) was dissolved in refluxing DMF. The obtained solution was cooled down to room
	temperature and anhydrous zinc chloride (10.28 g, 75.43 mmol) was added. The resulting suspension
	was stirred under reflux for 16 hours. The precipitate was isolated by filtration while hot and washed
	sequentially with hot DMF (2*100 mL) then MeOH (3*50 mL) and DCM (50 mL) and dried under
	vacuum. Formula: Zn5(BTDD)3Cl4
	Preparation of MIL-100(Al). Al(NO 3 ) 3 .9H 2 O (1.5 g, 4.0 mmol) and trimesic acid (1.26 g, 6.0 mmol) and
	distilled water (25 mL) were added in a 100 mL capacity microwave reactor (Mars) and the resulting
	suspension was stirred at room temperature for 5 minutes. The reactor was sealed and the reaction
	mixture was heated to 210°C (1200 W) within a 1 minute ramp then kept at this temperature for 5
	minutes under strong stirring. Immediately after reaction, the reactor was promptly cooled down
	using a water/ice bath. The white microcrystalline precipitate was isolated by filtration, washed with
	100 mL of H 2 O, then 100 mL of EtOH and dried in air. Finally the solid was resuspended in 125 mL
	EtOH and heated under reflux for 3 hours, then filtered and dried in air. Formula:
	Al 3 O(OH)(H 2 O) 2 (BTC) 2 .12H 2 O

BPI.Li. 2,6-bis(1H-pyrazol-1-yl)isonicotinic acid (HBPI) was prepared according to a previously reported procedure.

41 4

] 1MIL-100(Al).

Acknowledgements

The authors thank the ANR MOFSCO (ANR-18-CE09-0005-01) for funding. P. Chandra, G. Mouchaham, F. Nouar and J. Sun are thanked for their help on MOF synthesis Notes and references


Roberto M Souza
email: robertosouza@usp.br

Michell F C Ordoñez

Sabeur Mezghani

Samuel Crequy

Newton K Fukumasu

Izabel F Machado

Mohamed El Mansori

Orthogonal cut of SPS-sintered composites with ferrous matrix and Fe Mo S particles: Numerical and experimental analysis

Keywords:

des établissements d'enseignement et de recherche français ou étrangers, des laboratoires publics ou privés.

Introduction

In current industrial practice, there is an increasing demand for machining of workpieces of different materials at higher cutting speeds, maintaining high productivity and performance [START_REF] Perez | Effect of of cutting speed on surface integrity integrity face milled 7050-T7451 aluminium workpieces[END_REF][START_REF] Pu | Effect of dynamic recrystallization at tool-chip interface on accelerating tool wear during high-speed cutting of AISI1045 steel[END_REF]. Another trend is the reduction of liquid lubricants commonly used in wet machining, which is desirable from the point of view of decreasing operational cost and reducing environmental damage [START_REF] Diniz | Optimizing the use of dry cutting in rough turning steel operations[END_REF]. However, surface finishing precision and efficiency in dry machining usually impose additional requirements in terms of friction reduction as well as heat dissipation at the cutting tool/chip interface [START_REF] Liu | An investigation of side flow during chip formation in orthogonal cutting[END_REF][START_REF] Sutter | Temperature fields in a chip during high-speed orthogonal cutting-An experimental investigation[END_REF]. The cutting temperature is dependent on the cutting speed as well as on the materials in contact. For higher cutting temperatures, the hardness of some materials decreases significantly, modifying the geometry of the cutting edge and the workpiece surface finishing quality [START_REF] Sreejith | Dry machining: machining of the future[END_REF]. More specifically, for dry-machining of soft materials like austenitic stainless steel and aluminum, hardness reduction may difficult the material flow outside the cutting zone, due to adhesion and built-up edge formation, increasing the tool wear [START_REF] Maurotto | Effects of machining abuse on the surface properties of AISI 316L stainless steel[END_REF][START_REF] Medina-Clavijo | Microstructural aspects of the transition between two regimes in orthogonal cutting of AISI 1045 steel[END_REF][START_REF] Bahi | A new friction law for sticking and sliding contacts in machining[END_REF].

Previous works showed that the presence of low-shear materials at the chip/tool interface may improve the cutting by reducing both the friction coefficient and interfacial temperature [START_REF] Jianxin | Self-lubrication of Al 2 O 3/TiC/CaF 2 ceramic composites in sliding wear tests and in machining processes[END_REF][START_REF] Hao | Quantitative analysis of frictional behavior of cupronickel B10 at the tool-chip interface during dry cutting[END_REF]. These materials, frequently formed as tribofilms, originate from the tool material, are transferred from the workpiece, or from the interaction of the materials with the environment [START_REF] Yuan | Control of tribofilm formation in dry machining of hardened AISI D2 steel by tuning the cutting speed[END_REF]. That interaction usually produces compounds that are continuously generated and removed during the cutting operation. One common way to obtain these tribofilms involves the application of protective coatings onto the tool surface, leading to the presence of low-friction layers. Examples of soft coating used in this category are the solid lubricants, such as molybdenum disulfide (MoS 2 ), which present low shear resistance due to its lamellar structure [START_REF] Singer | Wear behavior of triode-sputtered MoS2 coatings in dry sliding contact with steel and ceramics[END_REF]. Hard coatings like diamond-like carbon or cubic boron nitride may also increase the tool life and reduce friction and wear, being suitable for dry machining operations [START_REF] Dai | The cutting performance of diamond and DLCcoated cutting tools[END_REF][START_REF] Bobzin | High-performance coatings for cutting tools[END_REF]. Although the choice of these coatings offers many advantages, formation of a brittle interface and the high deposition temperature of some processes may be a disadvantage [START_REF] Sreejith | Dry machining: machining of the future[END_REF]. Another method that involves tool modification is the fabrication of a micro-textured pattern produced by sintering or laser methods, for the formation of anti-adhesive cutting surfaces [START_REF] Sugihara | Development of novel cutting tools with dimple textured surfaces for dry machining of aluminum alloys[END_REF]. The addition of a solid lubricant such as graphite into of the micro-pattern generates a tribofilm on the rake face, which reduces friction [START_REF] Feng | Fabrication techniques and cutting performance of micro-textured self-lubricating ceramic cutting tools by insitu forming of Al2O3-TiC[END_REF].

An efficient route towards friction reduction involves sintering of a powder mixture containing the workpiece material and low-friction particles. The fabrication of these self-lubricant composites by advanced methods such as spark plasma sintering (SPS) has expanded its application field, mainly due to the possibility of obtaining compositelike materials with minimal porosity and with a set of tribological features that are not found in conventional materials [START_REF] Ordoñez | Effect of tungsten carbide addition on the tribological behavior of Astaloy 85Mo powder consolidated via spark plasma sintering[END_REF]. The sintering process offers the option of adding different solid lubricants, allowing an improvement in machinability, since tribofilms may be obtained without modification of the tool surface, avoiding or reducing the need of cutting fluids.

Different investigations reported the positive effect of solid lubricant addition in diverse sintered materials, for both friction and wear reduction. Addition of graphite and hexagonal boron nitride in iron matrix composites increased the scuffing resistance during dry sliding tests, due to the formation of a self-replenished tribolayer composed of both solid lubricants [START_REF] Hammes | Effect of hexagonal boron nitride and graphite on mechanical and scuffing resistance of self lubricating iron based composite[END_REF]. For Cu-WS 2 sintered composites, the bending strength and the tribological properties depend on the WS 2 particle size. For smaller particles, a higher interface area increases oxidation and formation of products that increase friction and wear. Moreover, larger particle sizes produced more resistant composites, due to the formation of a more continuous and protective tribofilm that reduced the wear rate [START_REF] Zhou | Effect of WS2particle size on mechanical properties and tribological behaviors of Cu-WS2composites sintered by SPS[END_REF]. Self-lubricant composites with steel matrix and MoS 2 addition produced a microstructure composed of MoS 2 and FeS phases that generated a tribofilm with lower friction and wear, during sliding tests conducted between 150 and 350 °C. However, above 350 °C the presence of MoO 3 and Fe 3 O 4 oxides reduced the MoS 2 lubrication and increased the FeS phase superficial migration, which provided lubrication at this temperature range [START_REF] Liu | Tribological behavior of M50-MoS2self-lubricating composites from 150 to 450 °C[END_REF]. In machining operations that were conducted using tools made of Al 2 O 3 /TiC ceramic matrix with additions of up to 10 vol% CaF 2, it was reported the formation of a low friction layer. This tribofilm formed at the tool-chip interface reduced the friction coefficient and protected the ceramic matrix from severe wear by brittle microfracture [START_REF] Jianxin | Self-lubrication of Al 2 O 3/TiC/CaF 2 ceramic composites in sliding wear tests and in machining processes[END_REF]. Previously mentioned sintering techniques allow to manufacture and obtain metal matrix composites with a range of reinforcement characteristics, such as particle size, morphology and volume fraction, which are difficult to control with conventional manufacturing methods. In the last years, the application of sintered materials was restricted, due to the presence of high porosity levels, which may reduce the mechanical and tribological properties. Nevertheless, the development of advanced sintering techniques such as SPS may lead to samples with higher densification and may expand its application to new fields in engineering. Therefore, it is important to study the mechanical and tribological properties of these SPS sintered composites and to analyze its behavior, for example during machining.

The finite element method (FEM) may be a complementary tool to evaluate the friction forces and the stress distribution at the chip-tool interface for different materials in cutting operations [START_REF] Mahnama | Application of FEM simulation of chip formation to stability analysis in orthogonal cutting process[END_REF][START_REF] Laakso | Using FEM simulations of cutting for evaluating the performance of different johnson cook parameter sets acquired with inverse methods[END_REF]. Overall, the material response depends on the contribution of different zones along this interface. The primary shear zone is located from the cutting edge to the work surface. High shear deformation at this zone is responsible for local heating and material softening [START_REF] Liu | An investigation of side flow during chip formation in orthogonal cutting[END_REF]. The material deformation and sliding along the rake face contributes with temperature rise at the secondary shear zone [START_REF] Sutter | Temperature fields in a chip during high-speed orthogonal cutting-An experimental investigation[END_REF]. The selection of the cutting conditions (i.e cutting speed, feed, depth) influences the temperature and strain distribution at the different zones. For example, FEM analyses of the cutting of pure Ti and Ti6Al4V were useful in predicting the segmented chip formation by adiabatic shear for different cutting speeds [START_REF] Bai | Improved analytical prediction of chip formation in orthogonal cutting of titanium alloy Ti6Al4V[END_REF][START_REF] Melkote | A physically based constitutive model for simulation of segmented chip formation in orthogonal cutting of commercially pure titanium[END_REF]. In analytical-numerical modeling of machining of ductile AISI 1045 steel with a TiAlN coated tool, sliding and sticking zones were defined at the tool-chip contact for determination of cutting forces, temperature and stress distribution, providing good agreement with experimental results [START_REF] Bahi | A new friction law for sticking and sliding contacts in machining[END_REF].

Few works have used FEM analyses to study machining of materials that present low shear phases [START_REF] Odum | Surface formation study using a 3-D explicit finite element model of machining of gray cast iron[END_REF][START_REF] Ljustina | A FE based machining simulation methodology accounting for cast iron microstructure[END_REF]. One example refers to the FEM simulation of the machining of commercial AISI 303 austenitic stainless steel containing manganese sulfide (MnS) particles. Results have indicated that during dry machining the sulfide secondary phase may act as stress concentration points. Thus, secondary phases with more irregular morphology or the presence of a higher volume fraction may increase these stress concentration regions and enhance the cutting efficiency [START_REF] Chagas | Numerical model of machining considering the effect of MnS inclusions in an austenitic stainless steel[END_REF].

This work investigated the cutting performance of SPS sintered Astaloy85Mo steel with additions of 2 wt% or 4 wt% of molybdenum disulfide (MoS 2 ). Analyses were based on orthogonal cut operations in different cutting depths. The effect of the solid lubricant on the microstructure, mechanical properties and on the tribofilm formation was also studied. The experimental results were compared with FEM simulations for the cutting of these materials in terms of friction forces and self-lubricant behavior.

Experimental and numerical procedure

Workpiece material

Low alloy steel Astaloy 85Mo (Fe-0.85 wt%Mo-0.27 wt% C) was selected as the matrix material. powder, with average particle size on the order of 100 μm, was milled with a planetary ball mill (Fritsch Pulverissete 4) using a ball-to-powder weight ratio of 10:1 under argon atmosphere. Each milling procedure was carried out in a total time of 10 h, which corresponds to 60 cycles with 10 min ON and 20 min OFF, to avoid overheating. Different amounts (2 and 4) wt. % of MoS 2 powder (Jet-Lube, Inc) were added to the ferrous powder. These materials were homogenized using a Y-type mixer during a period of 4 h. The mixtures were consolidated by spark plasma sintering (SPS) using a DR. SINTER SPS1050 (Sumitomo Coal Mining Co. Ltd, Japan) equipment. The powders were pressed uniaxially into a graphite mold of 20 mm of diameter with a pressure of 70 MPa and simultaneously heated up to a temperature of 1000 °C with a heating rate of 100 °C. min -1 . These conditions were held for 5 min. Overall, three materials were considered in this work: steel with additions of 0 (unreinforced), 2 wt % and 4 wt % of MoS 2 .

Tool material and cutting test

Dry cutting tests were conducted using a GSP-EL 136-shaper machine with triangular TCMT16T304 WC-Co inserts mounted on a tool holder. Tests were carried out with the same cutting speed, v c (32 m/ min) and with two cutting depths, d (0.15 and 0.20 mm). Three different passes were performed for each condition. During the cutting operation, tangential (F t ), radial (F r ) and axial (F a ) force components were measured by means of a piezoelectric Kistler dynamometer (type 9255B) (Fig. 1b). A denoising filter was later used to improve the signal-to-noise ratio during the cutting stage and the average force was calculated. To analyze the workpiece response for each condition an insitu observation was conducted using a high-speed camera with an acquisition rate of 6000 fps. The experimental procedure for material preparation, in conjunction with the cutting test, is summarized in Fig. 1a.

Material characterization

The apparent density (ρ a ) of each sintered material was measured using the Archimedes principle, according to the ASTM B962-17 standard. The relative density (ρ r ) was obtained by comparison of the composite density determined using the rule of mixtures, which includes each phase density (Astaloy 85Mo: 7.85 g cm -3 and MoS 2 : 5.06 g cm -3 ), with the experimental apparent density. For the metallographic preparation, samples were ground with 320-grade SiC paper and polished with diamond suspensions up to 1 μm. Microstructural observation and chemical composition analysis of each material was carried out by means of scanning electron microscopy (SEM -JEOL-JXA8600) with Electron Probe Microanalyzer (EPMA) using different 99.99% purity standards of Fe, C, Mo, and MoS 2 , for elemental energy calibration. The crystalline phases in the different samples after sintering were identified by means of X-ray diffraction analysis using an XRD-6000 Shimadzu diffractometer with Cu-Kα radiation and Bragg-Brentano (ϴ -2ϴ) geometry. Data was collected using a tube current of 30 mA and a tube voltage of 40 kV, with the angular 2ϴ range from 20°to 100°at a step size of 0.05°and 5s of integration time. XRD patterns were indexed with the aid of the X'Pert High Score Plus Software and the PDF-2 database (The International Centre for Diffraction Data-ICDD).

The bulk hardness of each material was determined using a Buelher 1900-2000 Vickers testing machine under a load of 10 Kgf (98.07 N) and using 10 s of holding time. Individual phase properties of the ferrous matrix and secondary phases were measured with instrumented indentation (nanoindentation) using a Hysitron TI950 triboindenter equipped with a Berkovich diamond tip, also capable of scanning probe microscopy (SPM) analyses. For Astaloy 85Mo with 2 wt% MoS 2 , five measurements were conducted for each phase with a peak load of 2 mN. The hardness and Young's modulus were obtained from load-displacement curves, according to the Oliver and Pharr method. After the cutting tests, the cross-section and the surface of the chips were analyzed using SEM (JEOL-JSM 6010LA) coupled with elemental mapping by energy dispersive X-ray spectrometry (EDS). The effect of the solid lubricant in the roughness and texture was measured with a 3D profilometer (Taylor Hobson -CCI HD).

FEM-finite-element modeling

Two-dimensional (2D) finite element modeling of the orthogonal cut of the metal matrix composites was conducted using the TRIBOCode digital tribology platform developed at the Surface Phenomena Laboratory (LFS-USP) [START_REF] Fukumasu | Numerical evaluation of cohesive and adhesive failure modes during the indentation of coated systems with compliant substrates[END_REF][START_REF] Fukumasu | Local transformation of amorphous hydrogenated carbon coating induced by high contact pressure[END_REF]. The FEM model considered both homogeneous and heterogeneous microstructures, to evaluate the influence of the sulfides during machining. Although literature reports the importance of temperature on the cutting process [START_REF] Liu | An investigation of side flow during chip formation in orthogonal cutting[END_REF][START_REF] Sutter | Temperature fields in a chip during high-speed orthogonal cutting-An experimental investigation[END_REF], models in this work considered isothermal room temperature conditions. This restriction allows the observation of the stress concentration features of the heterogeneous material, but does not take into account the decrease in matrix ductility usually promoted by the heating of the material during the cutting process. The tool was considered as rigid, with a tool cutting radius of 30 μm and a rake angle γ of 6°. The total machining length was 3.3 mm, such that only the initial steps of the cutting operation were simulated. Only the cutting depth d of 0.20 mm was considered.

The bulk geometry was discretized by a mesh with spatial resolution of 1 μm, totalizing 1,117,325 quadrilateral plane strain elements of type CPE4R. The heterogeneous material was modeled based on an SEM image of the composite material microstructure, in which mesh elements were selected to represent the morphology of the sulfides. Fig. 2 presents the heterogeneous mesh, in which the total cutting length was 6 mm and total sample thickness 1 mm. Ideal interfaces between matrix and sulfides were considered.

The material failure model for the matrix and the sulfides considered both ductile and shear modes, which are described in the literature for a large range of materials [START_REF] Hooputra | A comprehensive failure model for crashworthiness simulation of aluminium extrusions[END_REF][START_REF] Hedayat | Prediction of the bolt fracture in shear using finite element method[END_REF][START_REF] Hu | Finite element modeling of multiple transverse impact damage behaviors of 3-D braided composite beams at microstructure level[END_REF]. The failure modes allow the element deletion of the mesh, leading to crack nucleation and propagation throughout the material. The nucleation of cracks was based on initiation failure criteria, given by Eq. (1) and Eq. ( 2). In these equations, q p is the stress triaxiality, + q kp max is the shear stress ratio, f is the fracture strain and A B A B , , ,

d d s
s and C s are model constants. Ma- terial parameters, presented in Table 1, were tuned to reproduce the experimental cutting behavior, including similar cutting force magnitude. Values present physical meaning, but should not be considered as the precise properties of these materials. In the absence of data for Astaloy 85 Mo, the material failure energy, which allows the (1)

= + + + + A q kp B q kp C f s max s max s 2
(2)

Results and discussion

Powder characterization

In this work, ball milling was selected to reduce the Astaloy 85Mo powder particle size. Fig. 3 shows the particle size distribution of asreceived Astaloy 85Mo and molybdenum disulfide (MoS 2 ) powders. Both powders exhibit a monomodal distribution with a significant difference in average particle size, of 100 μm for Astaloy 85Mo and 5 μm for MoS 2 .

After 10 h of milling, the Astaloy 85Mo powder distribution changed to bimodal (Fig. 3) with a lower average particle size of 24 μm. This reduction is a result of continuous cycles of ball impingement, which cause plastic deformation, work-hardening and fracture. Fig. 4 presents SEM-EDS mapping of the powder of Astaloy 85Mo with 4 wt% MoS 2 after mixing in a Ye type mixer. This figure indicates a homogeneous mixture of well-dispersed MoS 2 particles on the surface of flake-like Astaloy 85Mo particles.

Workpiece characterization

Table 2 summarizes the density and hardness of the sintered samples. The selected choice of SPS parameters provided efficient densification for the Astaloy 85Mo, with relative density near 100%. Solid lubricant addition slightly reduced the density and increased the hardness, with a more significant effect on hardness for the sample with the addition of 2 wt% of MoS 2 .

Fig. 5 presents backscattered scanning (BSEM) micrographs of the sintered Astaloy 85 Mo with solid lubricant addition. The addition of 2 wt % of molybdenum disulfide contributed with the formation of a dark-gray phase with irregular morphology. This morphology may introduce stress concentration points due to local mechanical property differences between the sulfide particles and the surrounding matrix. The sample with 4 wt % MoS 2 addition exhibits the dark-gray irregular phase as well as a second lighter phase, which was smaller and with a rounded morphology.

EPMA quantitative analysis helped to determine the elemental composition of each phase, as shown in Fig. 6. The dark-gray phase, termed as phase 1, presented Fe (50.84 at. %) and S (46.85 at. %) as major constituents and corresponds to iron sulfide (FeS). In addition, the lighter and round phase, termed phase 2, with contents of Fe (53.64 at. %) and Mo (39.14 at%.), indicate iron molybdenum (Fe 3 Mo) formation. Note that the variability in atomic content may increase for phase 2 particles with dimensions below 1 μm, since measurements may include the surrounding matrix. Moreover, Fe, C and Mo contents for the ferrous matrix were 96.44 at. %, 1.51 at. % and 2.03 at. %, respectively. Thus, in comparison with the composition of the starting powder, a higher Mo content, of 2.03 at. % (3.48 wt%), was observed near the dark-gray phase boundaries, which possibly indicates the diffusion of the molybdenum from the MoS 2 into the ferritic-pearlitic matrix during sintering.

XRD patterns of the sintered samples are shown in Fig. 7. Analyses of the matrix phase indicated a body centre cubic structure that corresponds to ferrite (Fe-α) (PDF-01-087-0722). The addition of 2 wt% MoS 2 lead to the formation of an additional phase of FeeS (PDF-01-080-1027) and an indication of the presence of the Fe 3 Mo phase (PDF-00-009-0297). Higher (4 wt%) MoS 2 addition resulted in a higher presence of FeeS and Fe 3 Mo phases, as well as an indication of the presence of FeMo 3 S 4 (01-071-1897). The molybdenum content in the As-taloy85Mo matrix may enable the complete phase transformation of MoS 2 into FeeS and Fe 3 Mo phases [START_REF] Furlan | Self-lubricating composites containing MoS2: a review[END_REF]. For higher 4 wt% MoS 2 addition, the presence of an additional phase formation of FeeMoeS is also observed. Previous works on metal-matrix composites with the addition of MoS 2 powder, and sintering at temperatures above 950 °C, reported decomposition of MoS 2 into FeS and FeeMoeS more stable phases [START_REF] Liu | Tribological behavior of M50-MoS2self-lubricating composites from 150 to 450 °C[END_REF][START_REF] Dhanasekaran | Microstructure, strength and tribological behavior of Fe-C-Cu-Ni sintered steels prepared with MoS2 addition[END_REF]. The hardness and reduced Young's modulus of the phases detected in Astaloy 85Mo with the addition of 2 wt% MoS2 were determined by means of instrumented indentation testing. SPM image shows a sample region, indicating measurement positions inside a dark-gray iron sulfide and on the ferrous matrix (Fig. 8). Results revealed higher hardness and lower reduced Young's modulus for the dark-gray sulfide, in comparison with the ferrous matrix. The lower reduced modulus of these sulfides is related to its lamellar structure, which presents weak bonding between its multiple sheets and provides low shear resistance [START_REF] Oviedo | In situ TEM characterization of shear-stress-induced interlayer sliding in the cross section view of molybdenum disulfide[END_REF].

Cutting forces and chip formation

Fig. 9 shows the forces (Ft, Fr, Fa) measured during an orthogonal cut with a speed of 32 m/min and depth of 0.15 mm. The low hardness of Astaloy 85Mo without MoS 2 addition (Fig. 9a) may have increased the plastic deformation and adhesion on the rake face, which difficult the material flow and increases the radial (Fr) and tangential (Ft) forces with the cutting time and length (lc). Large oscillations at the beginning and at the end of the normal and tangential force signal data for the Astaloy 85Mo steel may be a result of tool vibration as well as the resistance of the to flow. The cutting forces were lower the composites (Fig. 9b andc), especially the tangential ones, which were negligible in most portions of the cutting time. In this work, friction coefficient can be calculated by ratio (Ft/Fr) between tangential and radial forces (Fig. 1b), as frequently described in tribological works and considering this particular experiment configuration. However, one must realize that friction coefficient may be calculated differently in publications dedicated to machining operations [START_REF] Son | Effects of the friction coefficient on the minimum cutting thickness in micro cutting[END_REF]. Fig. 10 compares the average values of radial and tangential forces measured for the sintered materials, for the two cutting depths. For Astaloy 85Mo without MoS 2 addition, the tangential and radial forces increased for the highest cutting depth, which may be related to the increment of material removal. Radial force signals were significantly lower for samples with MoS 2 addition at both cutting depths. Fig. 10 also confirms the trend observed in Fig. 9, in terms of the significantly lower tangential forces observed for the composites. These differences may be explained by the microstructural effect brought by the presence of the sulfide phases. Fig. 11 presents the radial and tangential force values obtained with the simulation of the cutting of the homogeneous (unreinforced) microstructure. In this figure, four values of friction coefficient for the contact between the tool and the workpiece were tested (zero, 0.3, 0.6 and 1.0), and provided a range from 600 to 800 N for the radial force and 50-300 N for the tangential force. The higher COF conditions presented similar cutting forces, suggesting that the effect of the friction coefficient saturates after a value close to 0.6. Although the simulated cutting forces are underestimated in comparison to the experimental values, the numerical procedure allows observing that lower friction coefficient values are capable of reducing both the radial and tangential forces throughout the cutting length (lc).

Chip formation is fundamental to analyze the material response under cutting conditions. The efficiency of chip breaking in dry machining operations influences the superficial quality and tool wear [START_REF] Zheng | Effect of cutting parameters on wear behavior of coated tool and surface roughness in high-speed turning of 300M[END_REF]. In-situ high-speed images at different elapsed times were obtained to understand the effect of solid lubricant addition on the efficiency of chip formation (Fig. 12). For Astaloy 85Mo steel, a built-up edge was formed in front of the tool. Higher elapsed times indicate that more material was accumulated (Fig. 12a) at the edge with the movement of the tool. In contrast, the addition of 2 wt% MoS 2 produces a continuous curled chip with higher removal efficiency (Fig. 12b andc). The benefits of the composite microstructure may be related to an ease in the cutting operation as well as the effect of the solid lubricant in facilitating shear at the secondary shear zone.

Fig. 13a shows the cross section of a chip of Astaloy 85Mo with MoS 2 addition. For the addition of 2 wt% (Fig. 13b), the presence of microcracks in front of the sulfide particles indicates a contribution to the fracture and flow stress at the primary shear zone. High plastic strain through the chip thickness produces highly elongated ferrite grains and extrusion of the dark irregular iron sulfides along a direction parallel to the chip-tool interface. These oriented sulfides may decrease the shear stress in the secondary zone [START_REF] Chagas | Numerical model of machining considering the effect of MnS inclusions in an austenitic stainless steel[END_REF]. For Astaloy85Mo with 4 wt% MoS 2 , an additional lighter FeMo 3 phase tends to align in the direction of deformation (Fig. 13c) and contribute to stress concentration, favoring the lower cutting forces obtained with this material. Elongated iron sulfide particles tend to be directed towards the chip surface and may form a transferred layer which also contributes to the friction reduction (lower tangential force). Further evidence is provided in Section 3.4. Numerical simulation of the cutting progress is presented in Fig. 14 for both homogeneous and heterogeneous microstructures. Fig. 14a presents formation of a continuous chip for the homogeneous material, and discontinuous chips for the heterogeneous (Fig. 14b). Fig. 14c shows a detail of a region in the heterogeneous microstructure. In this figure, sulfides were deformed by the cutting process as observed in Fig. 13c. In the simulations, discontinuous chip formation was related to the stress concentration promoted by the deformation of sulfides in the composites prepared with addition of solid lubricant. Despite the cracks in Fig. 13b andc, the same effect was not observed experimentally (Fig. 12b andc). This difference can be attributed to thermal effects responsible for an increase in matrix ductility. Literature reports thermal effects [START_REF] Ljustina | A FE based machining simulation methodology accounting for cast iron microstructure[END_REF] and differences between numerical and experimental results [START_REF] Mohammed | Modeling the effect of compacted graphite iron microstructure on cutting forces and tool wear[END_REF] for the machining of different types of cast iron. These effects may probably limit crack propagation as observed in the experimental portion of this work, but were not observed in the numerical simulations, due to the isothermal approach selected in model formulation. The material models selected in this work can also contribute with the differences between experimental and numerical results.

The positive effect of a lower friction coefficient was also observed in the numerical simulations, as presented in Fig. 15. In this figure, which considers only the homogeneous material, variations on the COF led to differences in the chip morphology and surface quality. The higher COF cases led to a thicker plastically deformed layer, measured from the free surface, which would be associated with higher deformation and residual stresses and lower surface integrity after the cutting process.

Chip surface characterization -secondary shear zone

Fig. 16 presents an SEM image and the corresponding EDS elemental mapping of a chip surface (side in contact with the tool) of Astaloy 85Mo with the addition of 2 wt% MoS 2, after orthogonal cut under a cutting speed (v c ) of 32 m/min and depth (d) of 0.15 mm. The results evidence the formation of a transferred layer of FeeMoeS, which is produced by the deformation and migration of sulfides along the chip surface. This superficial tribolayer reduces the material sticking and the friction force at the secondary zone. The presence of superficial cracks is a result of the large plastic deformation through the chip thickness.

Fig. 17 presents 3D and 2D optical profilometry analyses of the effect of solid lubricant addition on the surface finish after orthogonal cutting. Unreinforced Astaloy 85Mo presents material detachment produced by adhesion, as observed from the 3D surface profile (Fig. 17a). Adhesion induces material transfer and anisotropic texture that increases the 2D profile roughness due to the presence of deeper valleys. The material detached and transferred to the tool surface explains the built-up edge formation and the higher friction for this sintered sample [START_REF] Atlati | Effect of the local friction and contact nature on the Built-Up Edge formation process in machining ductile metals[END_REF]. Astaloy 85Mo with the addition of 2 wt% MoS 2 shows a 3D surface profile with negligible presence of material detachment and texture at the cutting direction (Fig. 17b). The 2D profile, without deep valleys, evidence low roughness and good surface finish due to the sulfide addition. A link between coefficient of friction and surface finish was also observed by the numerical simulations, in which the maximum plastically deformed depth (Fig. 15) was lower for lower COFs.

Table 3 summarizes the roughness parameters of the chip surface for the different sintered samples. Results indicate a significant difference in the roughness parameters of the plain Astaloy 85Mo, in comparison with samples with MoS 2 addition. The material detachment of Astaloy 85Mo increases the maximum valley depth (Sv) and the texture anisotropy. These superficial defects also influence the other roughness parameters that include the peak to valley ratio, i.e. the average mean height (Sa) and the root mean square (Sq). The roughness parameters were significantly lower for both 2 and 4 wt % solid lubricant contents. However, the highest addition, 4 wt % MoS 2 exhibits a slight reduction of the roughness parameters, which is in agreement with the lower cutting forces showed for the cutting depth of 0.20 mm, due to the additional amount of sulfides.

Conclusions

Orthogonal cutting tests were carried out under a speed of 32 m/ min and cutting depths of 0.15 mm and 0.20 mm, to analyze the effect of 2 and 4 wt% MoS 2 addition on the cutting forces and surface finish of SPS sintered Astaloy 85Mo steel. FEM simulation and experimental results show that the addition of solid lubricant is advantageous to the cutting performance:

1. Astaloy 85Mo sintered material exhibits low hardness and high plastic deformation that contribute to built-up edge formation and excessive adhesion and material detachment. This behavior results in high radial and tangential cutting forces. 2. The addition of 2 wt% MoS 2 produces a microstructural modification by the formation of irregular dark iron-sulfides that provide stress intensity points. This microstructure intensifies stresses and reduces the cutting forces. For samples sintered with 4 wt % MoS 2 , an additional lighter and round FeeMo phase is formed and contributes with additional stress concentration and cutting force reduction at the higher cutting depth of 0.20 mm. 3. Dark irregular iron sulfides are extruded by high plastic strain through the chip thickness, which enables shear and movement towards the chip surface. The low shear resistance of these sulfides increases the cutting efficiency and the formation of a well-defined curled chip. 4. The formation of a transferred layer on the chip surface (side in contact with the tool) contributes with the lower tangential friction force at the chip-tool interface, due to sticking reduction at the secondary shear zone. 5. Experimental and numerical analyses were able to link the influence of COF with the quality of surface finishing after machining, indicating that lower COF produces better quality finishing. 6. The sulfides present in the metal matrix acted as stress concentrators, as observed in the numerical and experimental analyses.

In both, cracks nucleated as a result of the sulfides.

Fig. 1 .

 1 Fig. 1. (a) Schematic representation of the samples preparation and cutting test evaluation and (b) photograph of the cutting setup before the first cutting pass.

Fig. 2 .

 2 Fig. 2. Discretization mesh of the cutting geometry. Regions in gray represent the matrix material while red indicates sulfides for the heterogeneous simulation. The detail in yellow shows the mesh discretization of the sulfide morphology. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)

Fig. 3 .

 3 Fig. 3. Particle size distribution of Astaloy85Mo (black) and MoS 2 (red) asreceived. Astaloy 85Mo (blue) distribution after10 h of milling. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)

Fig. 4 .

 4 Fig. 4. SEM-EDS mapping of Astaloy 85Mo with 4 wt% MoS 2 powder mixture.

Fig. 5 .

 5 Fig. 5. BSEM microstructure for Astaloy 85Mo with different MoS 2 additions (a) 2 wt% and (b) 4 wt%.

Fig. 6 .

 6 Fig. 6. EPMA quantitative analyses for constitutive phases of the Astaloy 85Mo with 4 wt% MoS 2 addition.

Fig. 7 .Fig. 8 .

 78 Fig. 7. X-ray diffractograms of unreinforced Astaloy 85Mo and with MoS 2 addition after sintering.

Fig. 9 .Fig. 10 .

 910 Fig. 9. Radial, axial and tangential force signals during orthogonal cutting under speed (v c ) of 32 m/min and depth (d) 0.15 mm: (a) Astaloy 85Mo steel and with addition of (b) 2 wt% MoS 2 and (c) 4 wt% MoS 2 .

Fig. 11 .

 11 Fig. 11. FEM calculation of radial (solid line) and tangential (dashed line) forces along the length (l c ) for the simulation of the homogeneous case with three COF conditions: 0.0 (black), 0.3 (blue), 0.6 (red) and 1.0 (green). (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)

Fig. 12 .Fig. 13 .

 1213 Fig. 12. High-speed sequential images of orthogonal cutting at different elapsed times for (a) Astaloy 85Mo, (b) Astaloy 85Mo -2 wt% MoS 2 and (c) Astaloy 85Mo -4 wt%MoS 2 .

Fig. 14 .

 14 Fig. 14. Evolution of chip formation in the numerical simulations of homogeneous (a) and heterogeneous (b) microstructure. Detail of the microstructure region is shown in (c), in which the sulfides are colored in red. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)

Fig. 15 .

 15 Fig. 15. Influence of the coefficient of friction on the plastically deformed depth after the cutting process for the homogeneous microstructure. Higher COF promotes deeper plastically deformed regions.

Fig. 16 .

 16 Fig. 16. SEM micrograph and EDS mapping of the chip surface showing tribofilm formation for Astaloy 85Mo -2 wt% MoS 2 under speed 32 m/min and depth 0.15 mm.

Fig. 17 .

 17 Fig. 17. 3D optical profilometry on the chip surface with 2D profile and texture isotropy in detail for (a) Astaloy 85Mo and (b) Astaloy 85Mo -2 wt% MoS 2 .

  The as-received (commercial) ferrous

	Nomenclature	v	Poisson's ratio
	ρ a	apparent density [g.cm -3 ]	H ε f	hardness [GPa] equivalent strain to fracture
	ρ r	relative density [%]	ε p	equivalent plastic strain
	FR	relative frequency [%]	σ 1	yield stress [MPa]
	v c γ	cutting speed [m/min] rake angle [ o ]	τ max A d , B d , A s , B s , C s , k workpiece model constants maximum shear stress [MPa]
	l c	cutting length [mm]	S a	arithmetic mean height surface roughness [μm]
	COF	coefficient of friction	S q	root mean square height surface roughness [μm]
	F t , F r , F a tangential, radial and axial force components [N]	S tr	surface texture isotropy percentage [%]
	d	cutting depth [μm]	S v	maximum valley height surface roughness [μm]
	E r	reduced elastic modulus [GPa]

Table 1

 1 Parameters of the material constitutive models.

	Parameters	Materials
		Matrix	Sulfides
	Elastic modulus, E (GPa)	210.0	160.0
	Poisson ratio, v (-)	0.3	0.3
	Yield Strength, σ 1 (MPa)	337.0	776.0
	A d (-)	-1.55	-0.5
	B d (-)	0.89	0.3
	As (-)	14.09	14.09
	Bs (-)	-48.75	-48.75
	Cs (-)	43.63	43.63
	k (-)	0.3

Table 2

 2 Density (apparent and relative) and Vickers microhardness for workpiece sintered materials.

	Material	Apparent density ρ a (g/cm 3 )	Relative density	Hardness
			ρ r (%)	H (HV 10 )
	Astaloy 85Mo	7.81	99.5	133.4 ± 2.0
	2 wt% MoS 2	7.65	98.2	239.2 ± 5.4
	4 wt% MoS 2	7.56	98.1	198.8 ± 4.2

Table 3

 3 Roughness parameters of the chip surface of the different sintered samples. Analysis of the surface in contact with the tool. Arithmetic mean height (Sa), root mean square height (Sq), maximum valley height (Sv) and surface texture isotropy percentage (Str).

	Material	Sa (μm)	Sq (μm)	Sv (μm)	Texture isotropy (%)
	Astaloy 85Mo	1.23	2.12	21.44	42.9
	2 wt% MoS 2	0.36	0.48	3.90	2.27
	4 wt % MoS 2	0.35	0.46	3.08	1.22

Acknowledgments

The authors would like to acknowledge the Brazilian research support agencies: Sao Paulo Research Foundation (FAPESP) Grant number 2016/18014-6 and the National Council for Scientific and Technological Development (CNPq).