Jérôme Andrieux
email: jerome.andrieux@univ-lyon1.fr

Bruno Gardiola

Olivier Dezellus

Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X-ray diffraction and modeling

Keywords: Metal matrix composites (MMCs), Solid state reaction, Interface Interdiffusion, Synchrotron diffraction, Theory and modeling (kinetics, transport, diffusion)

published or not. The documents may come

Synthesis of Ti matrix composites reinforced with TiC

particles: in situ synchrotron X-ray diffraction and modeling

Introduction

The high specific mechanical properties of Metal Matrix Composite (MMC) lead to much considerations over the past decades for structural lightening in aerospace applications or wear resistance in car breaks. Among the different materials used as reinforcement for Titanium based MMC, titanium carbide (TiC) was selected for its excellent chemical compatibility with the matrix [START_REF] Clyne | An introduction to Metal Matrix Composites[END_REF][START_REF] Vk | Recent advances in metal matrix composites[END_REF][START_REF] Miracle | Metal matrix composites -From science to technological significance[END_REF][START_REF] Huang | In situ TiC particles reinforced Ti6Al4V matrix composite with a network reinforcement architecture[END_REF]. The powder metallurgy route, that is widely used to produced Ti-based MMC [START_REF] Liu | Design of powder metallurgy titanium alloys and composites[END_REF], leads to the preparation of green compacts where Ti and stoichiometric TiC are not in thermodynamic equilibrium before the consolidation of the MMC. During the heating step, an interfacial reaction between Ti and TiC will occur as already shown and studied in the literature [START_REF] Quinn | Solid-State Reaction Between Titanium Carbide and Titanium Metal[END_REF][START_REF] Wanjara | Evidence for stable stoichiometric Ti2C at the interface in TiC particulate reinforced Ti alloy composites[END_REF]. More precisely, a preceding paper gave details of the general scenario of chemical interaction between the Ti matrix and the TiC particles during the consolidation step at high temperature [START_REF] Roger | Synthesis of Ti matrix composites reinforced with TiC particles: thermodynamic equilibrium and change in microstructure[END_REF]. Four key steps were proposed starting with dissolution of the smallest particles to reach C saturation in the Ti matrix, followed by a change in the TiC stoichiometry in order to reach thermodynamic equilibrium between matrix and reinforcement following reaction (1): y TiC + (1-y) Ti -> TiC y [START_REF] Clyne | An introduction to Metal Matrix Composites[END_REF] According to the assessment of the C-Ti binary system by Dumitrescu et al., the phase equilibrium is characterized by a y value ranging from 0.55 to 0.595 at 1000°C and 600°C

respectively [START_REF] Dumitrescu | A reassessment of Ti-C-N based on a critical review of available assessments of Ti-N and Ti-C[END_REF]. Reaction 1 corresponds to incorporation of Ti from the metallic matrix to the To cite this paper : J. Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

3 carbide phase. The reaction is therefore associated with two major phenomena: an increase in the total amount of reinforcement in the composite material and an increase in the mean particle radius by a factor of 1.19 [START_REF] Roger | Synthesis of Ti matrix composites reinforced with TiC particles: thermodynamic equilibrium and change in microstructure[END_REF]. As a consequence, the average distance between the particles decreases during the course of the reaction and contact between individual particles occurs in the most reinforced domain of MMC. This initiates an aggregation phenomenon that is responsible for the high growth rate of particles observed for heat treatment times lower than 1h. Finally, Ostwald ripening is responsible for the change in particles for longer heat treatment times [START_REF] Roger | Synthesis of Ti matrix composites reinforced with TiC particles: thermodynamic equilibrium and change in microstructure[END_REF].

The present work is focused on the two first steps, i.e. C saturation of the Ti matrix and the change in TiC stoichiometry resulting in the Ti-TiC composite material tending towards its thermodynamic equilibrium. The main objectives are an experimental determination of the kinetics of these two first steps, supported by a modeling of the diffusion phenomena occurring at the interface between a particle and the matrix. For this purpose, the reaction was investigated by in-situ synchrotron X-ray Diffraction (XRD) with a high-energy monochromatic Xray beam at the European Synchrotron Radiation Facility (ESRF -ID15B). Nowadays, it is quite common in metallurgy to perform simulations of phase formation and transformations by using software packages such as DICTRA [START_REF] Andersson | Thermo-Calc & DICTRA, computational tools for materials science[END_REF]. Such calculations require the use of two databases.

The first is a thermodynamic database that can provide a good description of the system; from the point of view of both phase equilibria, and energetic description of the different phases of the system. The second is a mobility database that is needed to calculate the diffusion coefficient inside the main phases of the system. To the authors' knowledge, despite the great success of such calculations in describing metallurgical processes, the same approach has not been used for the synthesis of Metal Matrix Composites, despite the fact that the two types of materials can be considered as very similar. In the present study, the change in the Ti matrix and its TiC reinforcement over several heat treatment periods is simulated with the DICTRA package in order to compare the results with the experimental study by in-situ X-ray diffraction.

As usual, if the simulations are able to reproduce experimental results obtained in wellcontrolled conditions, then they can also be used to extrapolate and predict microstructure changes in much more severe conditions.

To cite this paper : J.

Experimental procedures

Sample preparation

The starting materials used to prepare powder compacts were commercial TiC 0.96 and Ti powders (see Tables 1 to 3 for their compositions The small cylinders of powder compacts were inserted under argon in a graphite crucible previously filled with a layer of Ti powder that was used as a getter for gaseous impurities such as oxygen or nitrogen. The graphite crucible was heated by an induction coil under vacuum (~10 -4 mbar) using an induction furnace available at ESRF through the Sample Environment Support Service (SESS). The temperature was measured and controlled using a K-type thermocouple placed in the crucible and connected to a 2408i Eurotherm controller. In step mode, the heating rate was measured at 60K.s -1 with temperature stabilization of ±1K at 1273K (1000°C). The starting time of the experiment, t=0, corresponds to the beginning of the heating step.

Synchrotron XRD

Transmission X-Ray Diffraction (XRD) experiments were performed on beamline ID15B at ESRF (Grenoble) using a high-energy X-ray monochromatic beam (300x300 µm

2

) with a wavelength of 0.14370 Å (~86.3 keV) and a ∆E/E~10 -4 . To follow small variations in the diffraction peak position and improve the angular resolution (∆(2θ)=0.003 °(2θ)), the sampledetector distance was increased to 4.358 meters, leading to a cell parameter resolution of ∆a=0.002 Å. In addition, the detector was off-centered by 200 mm compared to the direct beam to select a 2θ range from 2.8 to 5°(2θ) and to focus the acquisition on the main TiC diffraction peaks, i.e. the (111)-3.295°2θ and the (200)-3.806°2θ peaks. Continuous acquisitions were performed using a MAR345 detector with an exposure time of 90s. Thus, diffraction rings were radially integrated using Fit2D software [START_REF] Hammersley | Two-dimensional detector software: From real detector to idealised image or two-theta scan[END_REF]. An example of a diffraction pattern collected on the sample at room temperature with peak identification is given in Figure 1.

To cite this paper : J. The structural study was performed by sequential Rietveld refinement using the FullProf Suite Software with the WinPlotr graphical interface [START_REF] Roisnel | WinPLOTR: A windows tool for powder diffraction pattern analysis[END_REF][START_REF] Rodríguez-Carvajal | Recent advances in magnetic structure determination by neutron powder diffraction[END_REF]. The profile function 7 in Full Prof, a

Pseudo-Voigt function with axial divergence correction, was used. The instrument resolution file was defined using LaB 6 as standard. Following the reduced 2θ range of acquisition due to the experimental configuration and the aim of the present paper, the Rietveld refinement was only performed on TiC diffraction peaks. For each refinement process, the refinement weighting factors were found at maximum R p <15%, R wp <20%, R exp <10%, χ 2 <5. More details in the sequential Rietveld refinement procedure are given in Section 3.1.

Experimental Results

Rietveld refinement

Titanium carbide has a halite structure (two face-centered cubic (fcc) sublattices with Ti and C, respectively). It displays deviations in stoichiometry over a wide range of homogeneity, from TiC 0.47 to TiC 0.99 , because of the presence of vacancies in the carbon sublattice [START_REF] Seifert | Thermodynamic optimization of the Ti-C system[END_REF].

Consequently, this phase is usually modeled in Calphad assessments by using the compound energy formalism with two sublattices having the same number of sites, one totally filled by Ti atoms, the other filled by a random mixture of C atoms and vacancies [START_REF] Dumitrescu | A reassessment of Ti-C-N based on a critical review of available assessments of Ti-N and Ti-C[END_REF][START_REF] Jonsson | Assessment of the Ti-C system[END_REF][START_REF] Frisk | A revised thermodynamic description of the Ti-C system[END_REF] leading to a chemical formula that can be expressed as TiC y in order to illustrate both the stoichiometry range and the location of vacancies on the C sublattice. Experimentally, the number of vacancies has a conspicuous effect on the lattice parameter. In a review of the Group 4a carbides, Storms [START_REF] Storms | A critical review of refractories. Part I. Selected properties of Group 4a, 5a and 6a carbides[END_REF] reported that the lattice parameter expands for 0.99 > y > 0.85 and shrinks for 0.85 > y > 0.5, the total change in lattice parameter being less than 1%, from 4.330nm to 4.298nm respectively [START_REF] Bittner | Magnetische untersuchungen der carbide TiC, ZrC, HfC, VC, NbC und TaC[END_REF][START_REF] Norton | Properties of non-stoichiometric metallic carbides[END_REF][START_REF] Rudy | Ternary phase equilibria in transition metal-boroncarbon-silicon systems[END_REF][START_REF] Storms | The refractory carbides[END_REF][START_REF] Ramqvist | Variation of Lattice Parameter and Hardness with Carbon Content of Group 4 B Metal Carbides[END_REF][START_REF] Vicens | Contribution to study of system Titanium-Carbon-Oxygen[END_REF][START_REF] Kiparisov | Effect of titanium carbide composition on the properties of titanium carbide-steel materials[END_REF][START_REF] Frage | Iron-titanium-carbon system. II. Microstructure of titanium carbide (TiCx) of various stoichiometries infiltrated with iron-carbon alloy[END_REF][START_REF] Fernandes | Characterisation of solar-synthesised TiCx (x=0.50, 0.625, 0.75, 0.85, 0.90 and 1.0by X-ray diffraction, density and Vickers microhardness[END_REF]. The unusual presence of a maximum in the lattice parameter vs. composition curve, corresponding to anomalous volume behavior of TiC at small vacancy concentration, was recently explained by Hugosson as resulting from an effect of local relaxation of the atoms surrounding the vacancy sites [START_REF] Hugosson | Phase stabilities and structural relaxations in substoichiometric TiC1-x[END_REF]. Consequently, changes in y of the TiC y phase are associated with a slight modification in both the lattice parameter (about 1%)

and the peak intensities because of variations in the number of vacancies in the C sublattice.

Thanks to the improved angular resolution of the present in-situ XRD study (∆a=0.002 Å), these variations were captured and analyzed by Rietveld refinement to follow in situ the pathway to equilibrium during heat treatment of a Ti-TiC composite material.

The TiC 0.96 particles used as starting material present a continuous particle size distribution ranging over three decades from 15-20 nm for the smallest to 7-10 µm for the biggest [START_REF] Roger | Synthesis of Ti matrix composites reinforced with TiC particles: thermodynamic equilibrium and change in microstructure[END_REF],

leading to different crystallite sizes, with constant starting stoichiometry (Table 2) and lattice parameters. Diffraction peaks of such a population of particles are characterized by a convolution of different contributions due to the different crystallite sizes at the same 2θ position: a sharp and intense peak due to the biggest particles (i.e. a mean diameter value of about 1 µm) and a broader peak due to the smallest particles (a few tenths of nm), leading to a broadening at the bottom of the TiC diffraction peak, as evidenced on Figure 2.a. This specific peak shape was fitted during Rietveld refinement by using two populations of TiC particles: the first, labeled "TiC 0.96 _SC", presents a small crystallite size and is related to the smallest particles whereas the second, labeled "TiC 0.96 _BC", is associated with the biggest particles. An example of a Rietveld refinement result of (200) TiC peak is given in Rietveld refinement, the chemical composition of these two populations corresponding to the starting material (i.e. C and Ti site occupancies) was fixed and kept constant whereas the intensity scale factor, the peak shape profile parameters and the lattice parameters were refined independently. In terms of temperature, the diffraction peaks associated with TiC 0.96 _BC and TiC 0.96 _SC shifted to lower 2θ values due to thermal expansion. In addition, new diffraction peaks were observed for a higher value of 2θ and associated with the formation of the substoichiometric TiC y phase, labeled "TiC y ". During the sequential Rietveld refinement, the composition, the intensity scale factor, the peak shape profile parameters and the lattice parameters of the TiC y phase were refined. An example of Rietveld refinement of (200) TiC diffraction peaks acquired at 900°C is given in Figure 2.b.

Fig. To cite this paper : J. Andrieux The time-dependent change in cell parameter is given on Figure 4. Note that t=0 on Figure 4 corresponds to the beginning of the heating step. First of all, this figure confirms the presence during the whole process of only two populations of TiC, characterized by two distinct values of lattice parameters. From a lattice parameter a RT =4.330Å at room temperature, the initial population corresponding to TiC 0.96 has a parameter of a 800°C =4.353 Å at 800°C and this cell parameter remains stable during the course of reaction 1. It may be deduced that the cell expansion for the TiC 0.96 due to the temperature increase is ∆a=0.023 Å. Concerning the TiC y population that forms at high temperature, Figure 4 shows an almost constant lattice parameter of a TiCy@800°C =4.336Å, the variation being less than 0.02% during the experiment. Assuming that the thermal expansion is the same for TiC 0.96 and TiC y , the lattice parameter of the TiC y population at RT can be estimated to be equal to a TiCy@RT =4.314Å. Following the correlation between the cell parameter and the stoichiometry of TiC y reported by Storms [START_REF] Storms | A critical review of refractories. Part I. Selected properties of Group 4a, 5a and 6a carbides[END_REF],

the composition of the TiC y population is found to be TiC 0.57 . This experimental determination of the TiC y composition that forms at high temperature is in good agreement with the value

To cite this paper : J. Andrieux expected when thermodynamic equilibrium is reached between the carbide and Ti phase at 1073K (800°C) [START_REF] Roger | Synthesis of Ti matrix composites reinforced with TiC particles: thermodynamic equilibrium and change in microstructure[END_REF][START_REF] Dumitrescu | A reassessment of Ti-C-N based on a critical review of available assessments of Ti-N and Ti-C[END_REF].

The formation of a sub-stoichiometric TiC y phase is confirmed at the interface between TiC 0.96 particles and the Titanium matrix as illustrated on STEM micrograph of a selected zone of the MMC microstructure (Figure 5). It leads to a core-shell structure of the particles, with a stoichiometric TiC core and a TiC y shell, in accordance with bright field mode contrasts. To cite this paper : J. Andrieux

13

The total mass fraction of reinforced TiC in the composite, calculated using equation ( 1), based on the experimental results obtained during the isothermal experiment at 1073K (800°C), is presented in Figure 7 (filled circles). The calculated mass fraction of reinforced TiC in the composite is of major concern as it is a key parameter governing the expected mechanical properties of the MMC.

As the in-situ experiment time was set to 3600s, final equilibrium was not reached but extrapolation based on total conversion of the initial stoichiometric population to the substoichiometric composition determined by XRD provides a means of estimating the equilibrium mass fraction of TiC that is found to be equal to 0.2303 (dashed-dotted horizontal line on Figure 7). This value is in good agreement with that expected after thermodynamic equilibrium is reached (0.2306). To cite this paper : J. Andrieux

Modeling

Simulation conditions

In order to model the changes that take place in the Ti-TiC metal matrix composite during high temperature heat treatment, the DICTRA software package [START_REF] Andersson | Computer simulation of multicomponent diffusional transformations in steel[END_REF] was used. DICTRA stands for diffusion-controlled transformation and is based on a numerical solution of the multicomponent diffusion equations and local thermodynamic equilibrium at the phase interfaces. The program is suitable for treating phenomena such as growth [START_REF] Borgenstam | DICTRA, a tool for simulation of diffusional transformations in alloys[END_REF], dissolution [START_REF] Agren | Kinetics of Carbide Dissolution[END_REF] and coarsening of particles in a matrix phase [START_REF] Gustafson | Coarsening of TiC in austenitic stainless steel -experiments and simulations in comparison[END_REF]. The DICTRA software package is built with databases for thermodynamics and diffusion. The present simulations were performed using the MOB2 mobility database provided by the ThermoCalc company and the thermodynamic assessment of the C-Ti system published by Dumitrescu et al. [START_REF] Dumitrescu | A reassessment of Ti-C-N based on a critical review of available assessments of Ti-N and Ti-C[END_REF].

The simulations were performed considering a spherical geometry, representative of reinforcing particles embedded in a metallic matrix. The carbide particle was defined as being spherical with an initial radius R TiC at the center of a spherical cell of radius R Mat defined by the surrounding metallic matrix (see Figure 8). the composite powder after the ball-milling step, before any high temperature treatment, is estimated to be D TiC = 128 nm. However, the residual presence of a tail corresponding to micron-sized particles must be taken into account. Note that the volume of matter corresponding to a 1µm particle is equivalent to 1000 particles or 8000 particles of respectively 100 and 50nm in size. It is difficult to obtain the real size distribution by counting particles from SEM observations because the quantity of biggest particles will be systematically overestimated. Therefore, the size distribution will be used as a guide giving the main trends, but exact numerical values cannot be used for modeling.

Fig. 9 Diameter distribution of TiC0.96 particles removed from the composite powder by selective acid etching after a ball-milling step and before any high temperature treatment.

As a consequence, the initial size of TiC 0.96 particles in the model was estimated from Figure 6 and more precisely from the volume ratio, r, between untransformed and transformed populations of TiC, respectively TiC 0.96 and TiC 0.57 populations. During the heat treatment, each particle can be considered as a core-shell structure with the residual TiC 0.96 composition as a core of radius R and the modified TiC 0.57 stoichiometry as a shell of thickness e. Given that the transformation occurs by a solid state diffusion process through the external shell, its thickness e depends only on the time of interaction and can be estimated by equation ( 2) regardless of the initial particle size (D inter is the interdiffusion coefficient in TiC 0.57 ).

To cite this paper : J. Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

16

Eq. 2

The ratio r of the core to shell volumes for a spherical particle, corresponding to the advancement of the reaction transforming the particle, is then given by equation 3:

Eq. 3

The radius of the residual inner TiC 0.96 core can then be estimated for a given diffusion time t and advancement of reaction r following equation ( 4):

Eq. 4

This equation can be used to estimate, for a given interaction time t, the radius of the residual TiC 0.96 core, leading to a given advancement of reaction r. According to Figure 6, the residual amount of the initial TiC 0.96 population is about 50% (r = 0.5) after t = 400s at 1073K (800°C).

The interdiffusion coefficient reported by van Loo et al. at this temperature for TiC 0.57 is equal to

.s -1 [START_REF] Van Loo | On the diffusion of carbon in titanium carbide[END_REF]. Therefore, numerical applications lead to a shell thickness e = 10nm and a core radius R = 23 nm.

From these simple calculations, it may be concluded that, considering a single particle size distribution, the very fast transformation rate of the initial TiC 0.96 composition can only be obtained for particles with an initial radius smaller than the typical value of about 30 nm. As a consequence, an initial TiC particle radius of 30nm is used for modeling (diameter equal to 60nm).

In the model, the thickness of the metallic shell in which the carbide particle is embedded is fixed in order to obtain a massive amount of reinforcement equal to 16.23 mass%, i.e. a value corresponding to the composite material experimentally studied. In the case of carbide particles with a diameter of 60 nm, this consideration leads to a Ti matrix shell thickness of 29 nm. To model the isothermal treatment at 1073K (800°C), the matrix and particle domains are considered as being respectively the HCP_A3 phase with an initial C mass fraction of 10 -6 (see table 2), and the carbide TiC 0.96 phase with a starting C composition of 19.74 mass%C (Table 3).

To cite this paper : J. Andrieux (800°C) of about 1 s. This result should be compared with Figure 6 where the Rietveld refinement of diffraction peaks shows the disappearance of the population of smallest particles after about 100 s. Obviously, some discrepancies are to be expected for such short times as the experimental heating rate is finite (i.e. 60K.s -1

), even if a step function toward 1073K (800°C)

was used for the induction heater and for calculations. Therefore, the typical times of the experimental results are expected to be slightly higher than those obtained from the kinetics calculations. More precisely, a careful analysis of Figure 4 reveals that the thermal expansion of the initial population of stoichiometric TiC 0.96 particles is observed experimentally in the center of the sample a few hundreds of seconds after starting the induction heater. Considering these limitations in the comparison of the shortest characteristic times for experiments and calculations, the agreement is relatively good and illustrates that chemical exchanges and solid state diffusion of C in the Ti matrix are extremely fast.

The solid line in Figure 7 presents the calculated change in the TiC mass fraction as a function of treatment time at 800°C in the case of TiC 0.96 particles with a unique size of 60 nm diameter.

The experimental mass fraction of TiC determined from the in-situ experiments is also reported (filled circles). The model correctly reproduces the fast increase in the TiC mass fraction observed by XRD in-situ measurements at the beginning of heat treatment. However, with a unique size of 60 nm, the final equilibrium is almost reached after only 1000 s whereas, experimentally, equilibrium is still not obtained after 3600s. As previously noted, the presence after milling of residual micron-sized particles is responsible for the experimentally observed two-step change: fast conversion of the smallest particles and much slower conversion of the biggest particles.

Given that diffusion processes control the change in the amount of TiC particles, this change is therefore limited by the presence of the few big particles with the initial TiC 0.96 composition remaining in the TiC population after milling. Therefore, in order to improve the modeling process and to illustrate the drastic influence of the residual presence of big particles, the same calculation was performed with a 1µm size particle and associated with the preceding results by

To cite this paper : J. Andrieux 7 with a dashed line. As expected, the presence of one micron-sized particle, in association with 6000 particles with a diameter of 60nm, allows the capture of both the initial transformation rate and the asymptotic trend towards the equilibrium state. Note that just one micron-sized particle represents 38% of the total volume of particles in the sample.

Finally, to illustrate the potential of thermokinetic modeling, a third class is added with an intermediate size of 160nm in diameter and the results are presented in Figure 7 with a dotted line (population in numbers are 6000, 200 and 2 for 60, 160 and 1000nm sized respectively).

Discussion

The present results highlight the fast kinetics of the trend towards thermodynamic equilibrium of a Ti-based matrix composite reinforced by TiC 0.96 particles. During isothermal treatment at 1073K (800°C), the dissolution of the smallest TiC 0.96 particles to reach saturation of the Ti matrix is obtained after a few tenths of a second, while the formation of the equilibrium composition of the carbide phase, TiC 0.57 , concomitantly increases sharply. Moreover, after only 6min of isothermal treatment at 1073K (800°C), 50% of the conversion of TiC particles from their initial TiC 0.96 composition towards the equilibrium value, TiC 0.57 , is achieved.

It is to be expected that such high reaction rates will have major consequences on the MMC synthesis process. First of all, reaching the C saturation of the Ti matrix induces the dissolution of about 10% of the initial TiC 0.96 particles (Figure 10.b). This dissolution process affects the smallest particles, which means that most of the effort made during milling to decrease the size of the initial particles is cancelled out. Next, the change in initial TiC composition (TiC 0.96 ) towards the equilibrium value (TiC 0.57 ) is achieved by partial conversion of the Ti matrix into carbide phase. This process therefore leads to an increase in the total amount of carbide in the composite from 16mass% to about 19mass% after 6min at 1073K (800°C) (Figure 7) and to an increase in particle size [START_REF] Roger | Synthesis of Ti matrix composites reinforced with TiC particles: thermodynamic equilibrium and change in microstructure[END_REF]. These changes induce a trend towards the formation of particle clusters that might have detrimental effects on the mechanical properties because of local embrittlement.

To cite this paper : J.

20

These two main processes, i.e. dissolution of the smallest particles and increase in total number of particles, occur during the industrial high temperature consolidation step of a Ti-based matrix composite reinforced by TiC particles. It is difficult to avoid them for two reasons: the driving force is thermodynamic and their kinetics is very fast. As a consequence, they have to be considered upstream in the process design: for example by reducing the initial quantity of TiC particles.

The reliability of thermokinetic modeling to simulate the change in TiC reinforcement in Ti-based composites has been demonstrated for the case of isothermal treatment (Figure 7). It can thus now be used to predict the change in mass fraction of reinforcement during the heat treatment that is performed prior to the consolidation step. Figure 11 presents a typical industrial heat treatment that involves inserting a billet, containing the cold compacted composite powder, inside a furnace heated to 1173 (900°C) for 1h. According to temperature measurements inside the consolidation furnace used for the experiment, the duration of the heating step is considered to be 10 min while the isothermal holding time is 50 min. The calculated change in TiC mass fraction, considering three particle size classes (as in Figure 7), is reported in Figure 11, as well as the experimental value determined just after the consolidation step. Once again, this figure highlights the fast kinetics of the reaction occurring inside the MMC as most of the transformation of the TiC stoichiometry, associated with the increase in particle mass fraction, is already achieved after a holding time of 10 min.

Fig. 11 Change in the calculated TiC mass fraction during heat treatment of a composite powder compact prior to the consolidation step (solid line). Temperature vs time is also reported (dashed line) as well as the experimental mass fraction determined after heat treatment and the consolidation step (filled circle).

To cite this paper : J. Andrieux

Conclusion

The reaction tending towards thermodynamic equilibrium during the synthesis of Ti/TiC MMC prepared by the powder metallurgy route was studied by in-situ synchrotron X-ray diffraction. It was found that the carbide composition changes rapidly from its initial stoichiometric composition TiC 0.96 towards a sub-stoichiometric value (TiC 0.57 ) corresponding to the thermodynamic equilibrium with the C saturated Ti matrix. The reaction is almost complete after only a few minutes at 800°C for the smallest particles, whereas the rate-limiting step is the particle size. Modeling of the diffusion processes in MMCs isothermal heat treatment at 800°C was performed using three particles size classes. First dissolution of the smallest particles (about 10% of the initial TiC 0.96 particles) is expected to be achieved after only 1s at 800°C.

Second the change in TiC composition lead to an increase in the total amount of carbide in the composite from 16 mass% to 19 mass%. The consequences on the industrial process of Ti/TiC MMC synthesis have also been considered. Typical industrial heat treatment of a MMC billet, 1h at 900°C, was modeled and the results showing an increase of the total amount of carbide in the composite from 16 mass% to 22 mass% are in rather good agreement with the experimental value (21 mass%). This illustrates the potential benefits of thermodynamic and kinetic modeling, combined to in-situ X-ray diffraction, in understanding and optimizing industrial processes for MMC synthesis.

Fig. 1

 1 Fig. 1 Example of diffraction pattern collected from the sample at room temperature with peak identification

Figure 2 .

 2 a. During the sequential To cite this paper : J. Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

9 3. 2

 92 Isothermal heating at 1073K (800°C) for 2h3.2.1. Raw dataThe time-dependent change in the TiC(200) peak during isothermal heating at 1073K (800°C) is shown on Figure3. The colorscale is related to the peak intensity. Note that the darker area around 4500s was due to synchrotron beam refill. It can be clearly seen from Figure3that a new diffraction peak appears at higher 2θ values for the TiC(200) peak, corresponding to the formation of a TiC population with smaller cell parameters, i.e. a sub-stoichiometric composition TiC y . The peak position of the TiC y composition remains constant as a function of time at 1073K (800°C), meaning that the lattice parameter and therefore the stoichiometry of the phase, remain constant. Finally, the intensity of the TiC y diffraction peak increases whereas that of the starting TiC 0.96 population decreases. Note that the same results were observed with the TiC(111) peak.

Fig. 3

 3 Fig. 3 Time-dependent change in the TiC(200) during isothermal treatment at 1073K (800°C). The colorscale indicates the peak intensity (on-line version).

Fig. 4

 4 Fig. 4 Time-dependent change in cell parameter during isothermal treatment at 800°C. t=0 corresponds to the beginning of the heating step.

Fig. 5

 5 Fig. 5 Bright field STEM view of core-shell microstructure of TiC particles in Ti-TiC MMC, obtained after 1min at 900°C.

Figure 6

 6 Figure6presents the time-dependent variation in the amount of the three identified populations of TiC particles (see Section 3.1) during isothermal annealing at 1073K (800°C). As already observed in Figure3, the quantity of the TiC y population increases whereas that of the initial TiC 0.96 composition decreases. More interestingly, the population of the initial smallest crystallites (TiC 0.96 _SC) is consumed after only 3 min of heat treatment. In addition, 50% conversion of TiC 0.96 _BC into TiC y phase is reached after only 6min of heat treatment. Finally, after 1h30 at 1073K (800°C), a complete reaction is not observed as ~25 % of TiC 0.96 _BC remains in the sample.

Fig. 6

 6 Fig.[START_REF] Quinn | Solid-State Reaction Between Titanium Carbide and Titanium Metal[END_REF] Changes in the quantity of different populations of TiC calculated using Rietveld refinement during isothermal treatment at 800°C. t=0 corresponds to the beginning of the heating step.

Fig. 7

 7 Fig. 7 Calculated time-dependent change in the mass fraction of TiC particles in a Ti matrix during isothermal heat treatment at 1073K (800°C). The solid line corresponds to one class of particles with a unique TiC diameter of 60nm. The dashed line shows the results obtained with two classes of TiC particles with different diameters: 6000 particles of 60nm, and one particle of 1µm. The dotted line corresponds to three classes of particles: 6000 particles of 60nm, 200 particles of 160nm and two particles of 1µm. The experimental values and the final asymptotic amount of TiC are also reported (filled circles and dasheddotted horizontal line respectively).

Fig. 8

 8 Fig. 8 Schematic model of the initial configuration of the system used in the DICTRA simulations

17

 17

4. 2

 2 Fig. 10 (a) Time-dependent change in the C content at the external Ti matrix interface during isothermal treatment at 1073K (800°C). (b) Time-dependent change in the calculated mass fraction of TiC particles in the Ti matrix at the very beginning of the isothermal treatment, during the dissolution step.

Figures 10 .

 10 Figures 10.a and b indicate that the dissolution of the smallest particles in order to allow carbon saturation of the Ti matrix is expected to be achieved after a typical isothermal time at 1073K

Table 1

 1 Chemical analysis of the Titanium powder used as starting material (SCA-CNRS Solaize)

	). Small cylinders of powder compacts (3 mm

Table 2

 2 Chemical analysis of the TiC0.96 powder used as starting material (SCA-CNRS, Solaize)

				Element concentration (ppm in mass)
		Fe	Cr	V	Ni	Ca	Cu, Cl, K, Zn, S, Er, As
	TiC 0.96 (starting powder)	1200	118	170	116	77	traces

Table 3 X

 3

	2.2 STEM characterization
	Sample for STEM characterization was placed in a liquid tight graphite crucible and then
	immersed for 1 min in a liquid aluminium bath hold at 900°C. After 1min, the crucible was
	quenched in water. STEM characterization was performed on a JEOL-JEM 2100F microscope
	in Bright field mode, under an accelerating voltage of 200 kV and with a magnification of
	x40000.

-ray fluorescence analysis of the TiC0.96 powder used as starting material (SCA-CNRS, Solaize) To cite this paper : J. Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

	3.3.2. Rietveld refinement results
	10

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.a simple linear combination. The relative amount of the two classes of particles is defined in order to get a 50% volume fraction of the smallest class, a value that should allow the initial fast reaction rate to be captured. Calculation results are reported in Figure

  Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

  , B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

	21

Acknowledgements

This work was undertaken in the framework of the COMETTi project funded by the French national research agency (ANR) [grant number ANR-09-MAPR-0021]. O.D. is very grateful to Dr. S. Fries and Pr. I. Steinbach from ICAMS institute at Bochum University (Germany) for allowing DICTRA calculations on their informatics cluster. J.A. thanks ID15B beamline staff for their help during beam time and ESRF for the provision of beamtime through in-house research during his post-doctoral position. The authors thank Gilles RENOU for the STEM observation performed at the "Consortium des Moyens Technologiques Communs" (CMTC, http://cmtc.grenoble-inp.fr). Composite powders were provided by the Mecachrome company (www.mecachrome.fr).

To cite this paper : J. Andrieux, B. Gardiola, O. Dezellus, Synthesis of Ti matrix composites reinforced with TiC particles: in-situ synchrotron X ray diffraction and modeling, Journal of Materials Science 2018, Accepted for publication, doi 10.1007/s10853-018-2258-8.

Conflicts of interest statement

The authors declare that they have no conflict of interest.


Quentin Scanvion

Sandrine Morell-Dubois

Cécile M Yelnik

Johana Bene

Sophie Gautier²

Marc Lambert
email: marc.lambert@chru-lille.fr

Pr Marc Lambert

FAILURE OF RIVAROXABAN TO PREVENT THROMBOSIS IN FOUR PATIENTS WITH ANTI-PHOSPHOLIPID SYNDROME

Comment on: Failure of rivaroxaban to prevent thrombosis in four patients with anti-phospholipid syndrome

Quentin Scanvion, Sandrine Morell-Dubois, Cécile M Yelnik, Johana Bene,

Sophie Gautier, Marc Lambert

To cite this version:

Letter to the Editor Sir, we read with a great interest the letter published recently by Dufrost et al. [START_REF] Dufrost | Failure of rivaroxaban to prevent thrombosis in four patients with anti-phospholipid syndrome[END_REF] on failure of rivaroxaban to prevent thrombosis in patients with APS, as well as the interview between Dr.

Hannah Cohen and Prof. Bernard Lauwerys [START_REF]Rivaroxaban versus warfarin to treat patients with thrombotic antiphospholipid syndrome. Dr. Hannah Cohen about the results of the RAPS trial[END_REF]. We think that this subject is important because we have observed growing use of direct oral anticoagulants (DOAC) in APS patients.

Rivaroxaban in Anti-Phospholipid Syndrome (RAPS) study is the only randomized controlled trial studying the use of rivaroxaban versus warfarin to prevent thrombosis recurrence [START_REF] Cohen | Rivaroxaban versus warfarin to treat patients with thrombotic antiphospholipid syndrome, with or without systemic lupus erythematosus (RAPS): a randomised, controlled, open-label, phase 2/3, non-inferiority trial[END_REF].

The primary outcome was the percentage change in endogenous thrombin potential, selected because of the low frequency of clinical outcomes and the ability of this particular test to measure biological activity of both drugs according to Cohen's interview. The authors conclude that rivaroxaban could be an effective alternative to vitamin K antagonist (VKA) therapy in APS patients since no thrombotic event was observed. Although we appreciate the randomized controlled nature of the study, we think that author's conclusion is an overstatement because of the short duration of the study follow-up (210 days), and the nature of the study which was not designed to demonstrate clinical non-inferiority. We also think that predicting a drug's efficacy in APS only on the basis of anticoagulation does not reflect the complexity of APS pathophysiology, which involves platelets, endothelial cells, monocytes, and immune and inflammatory mechanisms as well [START_REF] Giannakopoulos | The pathogenesis of the antiphospholipid syndrome[END_REF].

To illustrate, we report two new cases of APS patients; they both relapsed under DOAC, after having been stable under VKA therapy.

The first patient is a 66 year-old SLE female diagnosed in 1967 (oral ulcers, polyarthritis, ANA, immune thrombocytopenic purpura). warfarin. An aortic valve replacement by mechanical prosthesis was required.

We notice that DOACs are increasingly prescribed to treat APS patients, whatever the expression of their disease. Among the DOACs, rivaroxaban is primarily prescribed for the treatment of APS according to initial RAPS results [START_REF] Cohen | Rivaroxaban versus warfarin to treat patients with thrombotic antiphospholipid syndrome, with or without systemic lupus erythematosus (RAPS): a randomised, controlled, open-label, phase 2/3, non-inferiority trial[END_REF].

However evidence-based medicine inspires caution. The Task Force on APS Treatment

Trends of the 15 th International Congress on Antiphospholipid Antibodies stated that there was insufficient evidence to make recommendations at this time regarding the use of these DOACs in the APS, which can only be considered when there is known VKA allergy/intolerance, in patients with only venous APS. Thus VKA remains the mainstay of anticoagulation in thrombotic APS and the non-adherence is not a reason to switch [START_REF] Erkan | 14th International Congress on Antiphospholipid Antibodies Task Force Report on Antiphospholipid Syndrome Treatment Trends[END_REF].

In APS-ACTION registry, an international multicenter prospective database, of 428 thrombosis APS patients, 19 were under DOAC, of whom 6 had relapse during the 2-years follow-up (15.8% annual thrombosis risk), compared to 1.5% risk in VKA-receiving patients [START_REF] Unlu | Antiphospholipid syndrome alliance for clinical,trials,and international networking (APS action) clinical database and repository analysis: direct oral anticoagulant use among among antiphospholipid syndrome patient[END_REF], suggesting that DOAC and VKA do not have similar efficacy. Furthermore Cohen's study included only patients at low risk (neither prior arterial thrombotic event nor previous relapse with INR between 2.0 and 3.0). Because it is difficult to evaluate if "standard intensity" [START_REF]Rivaroxaban versus warfarin to treat patients with thrombotic antiphospholipid syndrome. Dr. Hannah Cohen about the results of the RAPS trial[END_REF] anticoagulation by VKA will be safe before starting, DOAC should be considered only as a second line treatment. Moreover an inaugural venous event does not exclude later arterial thromboses. Our first patient had no prior arterial history before the catastrophic APS. Thus we endorse the conclusion of Dufrost et al. [START_REF] Dufrost | Failure of rivaroxaban to prevent thrombosis in four patients with anti-phospholipid syndrome[END_REF] and do not agree that rivaroxaban offers an effective, safe and convenient alternative to warfarin in APS patients, as Cohen suggested [START_REF]Rivaroxaban versus warfarin to treat patients with thrombotic antiphospholipid syndrome. Dr. Hannah Cohen about the results of the RAPS trial[END_REF].

The manifestations of APS span a heterogeneous clinical spectrum. Awaiting randomized controlled trials with clinical outcome [START_REF] Pengo | Efficacy and safety of rivaroxaban vs warfarin in high-risk patients with antiphospholipid syndrome: Rationale and design of the Trial on Rivaroxaban in AntiPhospholipid Syndrome (TRAPS) trial[END_REF][START_REF] Woller | Apixaban for the Secondary Prevention of Thrombosis Among Patients With Antiphospholipid Syndrome: Study Rationale and Design (ASTRO-APS)[END_REF] and prolonged follow-up to clarify whether DOACs are efficient alternatives to VKAs, cautions and stringent clinical review are especially necessary in known high risk patients (table 1).

  SLE associated-APS was diagnosed in 2007 after a deep vein thrombosis of the right arm in the context of LA and anti-beta2gp1 positivity. She remained relapse-free under VKA therapy during 8 years and then switched to rivaroxaban (20mg) in 2015. Because of refractory ITP (corticosteroids, rituximab and splenectomy) eltrombopag was started. Six months later (one year after rivaroxaban switching), she experienced chest pain associated with elevated troponin, leading to the diagnosis of myocardial micro-thrombosis on MRI of the heart. Concomitantly, multiple cerebral ischemia and cutaneous arterial micro-thrombosis (confirmed by biopsy) occurred within a week, which led to the diagnosis of catastrophic APS. She improved after corticosteroids, intravenous immunoglobulin, curative heparin therapies, plasmapheresis and rivaroxaban/eltrombopag withdrawal. The second patient is a 44 year-old female diagnosed with primary triple-positive APS in 2007 (stroke with Libman-Sacks endocarditis and obstetrical morbidity) and treated by warfarin. Despite subtherapeutic INR, no relapse was observed on yearly-repeated echocardiography and cerebral MRI during the 7 years follow-up. Rivaroxaban (20mg) was initiated in 2014. The compliance was suboptimal. In May 2017 Libman-Sacks endocarditis relapsed with new ischemic strokes. Remission improved after switching by heparin, then

Table 1 : Clinical phenotypes of APS require caution with the use of direct oral anticoagulants. Lack of evidence-based medicine No evidence-based medicine

 1

	Venous thrombotic APS without previous	Arterial and small vessel thrombotic APS
	relapse with INR between 2.0 and 3.0
		Libman-Sacks endocarditis
		Association with pro-thrombotic treatment
		Triple-positive aPL profile
		Poor compliance (no biological monitoring
		and short half time)
		Relapsing patient despite anticoagulant.

Conflicts of interest:

Prof. LAMBERT receives fees from BAYER, BMS-PFIZER and DAICHY-SANKIO. QS, SMD, CMY, JB and SG declare no conflicts of interest.

Funding source: No specific funding was received from any bodies in the public, commercial or not-for -profit sectors to carry out the work described in this manuscript.


Wolfram Kloppmann

Neus Otero

Daren C Gooddy

Dan Lapworth

Ben Surridge

Emmanuelle Petelet

Christine Flehoc

Nicole Baran

Limitations of the isotopic composition of nitrates as a tracer of their origin

Nitrogen and oxygen isotopes are traditionally considered and frequently used as tracers of nitrate sources in watersheds used for drinking water production. The enrichment of synthetic nitrate-containing fertilizers in 18O due to the contribution of atmospheric oxygen in the production process confers a specific isotopic fingerprint to mineral fertilizers. In spite of the still widespread use on nitrate-containing synthetic fertilizers, their characteristic N and O isotope signatures are rarely unambiguously observed in nitrate-contaminated groundwater. We postulate, in line with Mengis et al. (2001), that fertilizer-derived nitrate is not directly and rapidly transferred to groundwater but rather retained in the soil-plant system as organic N and then mineralized and re-oxidized (termed the mineralization-immobilization turnover, MIT) thereby re-setting the oxygen isotope composition of nitrate and also changing its N isotope ratios. We show examples from watersheds on diverse alluvial/clastic and carbonate aquifers in eastern and northern France where, in spite of the use of mineral fertilizers, evidenced also through other isotopic tracers (boron isotopes), both N and O-isotope ratios are very homogeneous and compatible with nitrification of ammonium where 2/3 of oxygen is derived from soil water and 1/3 from atmospheric O 2 . These field data are corroborated by lysimeter data from Canada. Even if in areas where ammonium is derived from chemical fertilizers, N values still tend to be lower than in areas where ammonium is derived from manure/sewage, this is clearly a limitation to the dual isotope method (N, O) for nitrate source identification, but has important implications for the nitrogen mobility and residence time in soils amended with synthetic fertilizers (Sebilo et al., 2013).


Guangbo Hao
email: g.hao@ucc.ie

Haiyang Li

Abhilash Nayak

Stéphane Caro

Stephane Caro

Design of a Compliant Gripper With Multimode Jaws

Introduction

A recent and significant part of the paradigm shift brought forth by the industrial revolution is the miniaturization of electromechanical devices. Along with the reduction of size, miniaturization reduces the cost, energy and material consumption. On the other hand, the fabrication, manipulation and assembly of miniaturized components is difficult and challenging. Grippers are grasping tools for various objects, which have been extensively used in different fields such as material handling, manufacturing and medical devices [START_REF] Verotti | A Comprehensive Survey on Microgrippers Design: Mechanical Structure[END_REF]. Traditional grippers are usually composed of rigid-body kinematic joints, which have issues associated with friction, wear and clearance/backlash [START_REF] Nordin | Controlling mechanical systems with backlash-a survey[END_REF]. Those issues lead to poor resolution and repeatability of grasping motion, which makes the high-precision manipulation of miniaturized components challenging. In addition to being extremely difficult to grip sub-micrometre objects such as optical fibres and micro lens, traditional grippers are also very hard to grip brittle objects such as powder granular. This is because the minimal incremental motion (i.e. resolution) of the jaw in the traditional gripper is usually larger than the radius of the micro-object or already causes the breaking of the brittle object. Figure 1 shows a parallel-jaw gripper as an example. Although advanced control can be used to improve the gripper's resolution, its effort is trivial compared to the resulting high complexity and increased cost [START_REF] Nordin | Controlling mechanical systems with backlash-a survey[END_REF].

Figure 1: Comparison of traditional parallel-jaw gripper's resolution and size/deformation of objects

Although mechanisms are often composed of rigid bodies connected by joints, compliant mechanisms that include flexible elements as kinematic joints can be utilised to transmit a load and/or motion. The advances in compliant mechanisms have provided a new direction to address the above rigid-body problems easily [START_REF] Howell | Compliant Mechanisms[END_REF]. The direct result of eliminating rigid-body kinematic joints removes friction, wear and backlash, enabling very high precision motions. In addition, it can be free of assembly when using compliant mechanisms so that miniaturization and monolithic fabrication are easily obtained. There are mainly two approaches to design compliant grippers. The first one is the optimisation method [START_REF] Zhu | Topology optimization of hinge-free compliant mechanisms with multiple outputs using level set method[END_REF] and the second is the kinematic substitution method [START_REF] Hao | Conceptual designs of multi-degree of freedom compliant parallel manipulators composed of wire-beam based compliant mechanisms[END_REF]. The former optimizes the materials' distribution to meet specified motion requirements, which includes topology optimization, geometry optimization and size optimization. However, the optimization result often leads to sensitive manufacturing, and therefore minor fabrication error can largely change the output motion [START_REF] Wang | Design of Multimaterial Compliant Mechanisms Using Level-Set Methods[END_REF]. Also, the kinematics of the resulting gripper by optimization is not intuitive to engineers. The latter design method is very interesting since it takes advantage of the large number of existing rigid-body mechanisms and their existing knowledge. It renders a very clear kinematic meaning, which is easily used for shortening the design process. Parallel or closed-loop rigid-body architectures gain an upper hand here as their intrinsic properties favour the characteristics of compliant mechanisms like compactness, symmetry to reduce parasitic motions, low stiffness along the desired degrees of freedom (DOF) and high stiffness in other directions.

Moreover, compliant mechanisms usually work around a given (mostly singular) position for small range of motions (instantaneous motions). Therefore, parallel singular configurations existing in parallel manipulators may be advantageously exploited [START_REF] Amine | Classification of 3T1R parallel manipulators based on their wrench graph[END_REF][START_REF] Maraje | Operation modes comparison of a reconfigurable 3-PRS parallel manipulator based on kinematic performance[END_REF][START_REF] Nurahmi | Dimensionally homogeneous extended jacobian and condition number[END_REF][START_REF] Rubbert | Using singularities of parallel manipulators for enhancing the rigid-body replacement design method of compliant mechanisms[END_REF][START_REF] Rubbert | Design of a compensation mechanism for an active cardiac stabilizer based on an assembly of planar compliant mechanisms[END_REF][START_REF] Zlatanov | Constraint Singularities as C-Space Singularities[END_REF]. Parallel singularity can be an actuation singularity, constraint singularity or a compound singularity as explained in [START_REF] Amine | Classification of 3T1R parallel manipulators based on their wrench graph[END_REF][START_REF] Maraje | Operation modes comparison of a reconfigurable 3-PRS parallel manipulator based on kinematic performance[END_REF][START_REF] Nurahmi | Dimensionally homogeneous extended jacobian and condition number[END_REF]. Rubbert et al used an actuation singularity to typesynthesize a compliant medical device [START_REF] Rubbert | Using singularities of parallel manipulators for enhancing the rigid-body replacement design method of compliant mechanisms[END_REF][START_REF] Rubbert | Design of a compensation mechanism for an active cardiac stabilizer based on an assembly of planar compliant mechanisms[END_REF]. Another interesting kind of parallel singularity for a parallel manipulator that does not depend on the choice of actuation is a constraint singularity [START_REF] Zlatanov | Constraint Singularities as C-Space Singularities[END_REF]. Constraint singularities may divide the workspace of a parallel manipulator into different operation modes resulting in a reconfigurable mechanism. Algebraic geometry tools have proved to be efficient in performing global analysis of parallel manipulators and recognizing their operation modes leading to mobility-reconfiguration [START_REF] Husty | Algebraic methods in mechanism analysis and synthesis[END_REF][START_REF] Nurahmi | Reconfiguration analysis of a 4-RUU parallel manipulator[END_REF][START_REF] He | Design and Analysis of a New 7R Single-Loop Mechanism with 4R, 6R and 7R Operation Modes[END_REF]. The resulting mobility-reconfiguration can enable different modes of grasping in grippers. Thus, the reconfigurable compliant gripper unveils an ability to grasp a plethora of shapes or adapt to specific requirements unlike other compliant grippers in literature that exhibit only one (mostly parallel mode) of these grasping modes [START_REF] Beroz | Compliant microgripper with parallel straight-line jaw trajectory for nanostructure manipulation[END_REF][START_REF] Hao | Design and static testing of a compact distributed-compliance gripper based on flexure motion[END_REF].

Though there are abundant reconfigurable rigid-body mechanisms in the literature, the study of reconfigurable compliant mechanisms is limited. Hao studied the mobility and structure reconfiguration of compliant mechanisms [START_REF] Hao | Mobility and Structure Re-configurability of Compliant Mechanisms[END_REF] while Hao and Li introduced a position-space based structure reconfiguration approach to the reconfiguration of compliant mechanisms and to minimize parasitic motions [START_REF] Hao | Position space-based compliant mechanism reconfiguration approach and its application in the reduction of parasitic motion[END_REF]. Note that in rigid-body mechanisms, using the underactuated/adaptive grasping method [START_REF] Birglen | Self-adaptive mechanical finger and method[END_REF][START_REF] Laliberté | Underactuation in robotic grasping hands[END_REF], a versatile gripper for adapting to different shapes can be achieved. In this paper, one of the simplest yet ubiquitous parallel mechanisms, a fourbar linkage is considered at a constraint singularity configuration to design a reconfigurable compliant fourbar mechanism and then to construct a reconfigurable compliant gripper. From our best understanding, this is the first piece of work that considers a constraint singularity to design a reconfigurable compliant mechanism with multiple operation modes, also called motion modes. This remaining of this paper is organised as follows. Section 2 describes the design of a multi-mode compliant four-bar mechanism and conducts the associated kinematic analysis. The multi-mode compliant gripper is proposed in Section 3 based on the work presented in Section 2, followed by the analytical kinetostatic modelling. A case study is discussed in Section 4, which shows the analysis of the gripper under different actuation schemes. Section 5 draws the conclusions.

Design of a multi-mode compliant four-bar mechanism 2.1 Compliant four-bar mechanism at its singularity position

A comprehensive singularity and operation mode analysis of a parallelogram mechanism is reported in [START_REF] Nayak | A Reconfigurable Compliant Four-Bar Mechanism with Multiple Operation Modes[END_REF], using algebraic geometry tools. As a specific case of the parallelogram mechanism, a four-bar linkage with equilateral links as shown in Fig. 2 is used in this paper, where the link length is l. Link AD is fixed, AB and CD are the cranks and BC is the coupler. Origin of the fixed frame, O0 coincides with the centre of link AD while that of the moving frame O1 with the centre of BC. The bar/link BC is designated as the output link with AB/CD as the input link. The location and orientation of the coupler with respect to the fixed frame can be denoted by (a, b, ϕ), where a and b are the Cartesian coordinates of point O1 attached to the coupler, and ϕ is the orientation of the latter about z0-axis, i.e., angle between x0 and x1 axes. The two constraint singularity positions of the equilateral four-bar linkage are identified in Fig. 3 [START_REF] Nayak | A Reconfigurable Compliant Four-Bar Mechanism with Multiple Operation Modes[END_REF]. At a constraint singularity, the mechanism may switch from one operation mode to another. Therefore, in case of the four-bar linkage with equal link lengths, the DOF at a constraint singularity is equal to 2. In this configuration, points A, B, C and D are collinear and the corresponding motion type is a translational motion along the normal to the line ABCD passing through the four points A, B, C and D, combined with a rotation about an axis directed along z0 and passing through the line ABCD. Eventually, it is noteworthy that two actuators are required in order to control the end-effector in those constraint singularities in order to manage the operation mode changing. Based on the constraint singularity configuration of the four-bar rigid-body mechanism represented in Fig. 3, a compliant four-bar mechanism can be designed through kinematically replacing the rigid rotational joints with compliant rotational joints [START_REF] Howell | Compliant Mechanisms[END_REF][START_REF] Hao | Conceptual designs of multi-degree of freedom compliant parallel manipulators composed of wire-beam based compliant mechanisms[END_REF] in the singularity position. Note that, the singularity position (Fig. 3(a)) is the undeformed (home) configuration of the compliant mechanism. Each of the compliant rotational joints can be any type compliant rotational joint such as cross-spring rotational joint, notch rotational joint and cartwheel rotational joint [START_REF] Howell | Compliant Mechanisms[END_REF].

In this section, the cross-spring rotational joint as the rotational/revolute joint (RJ) is employed to synthesize the design of a reconfigurable compliant four-bar mechanism based on the identified singularity (Fig. 4). In Fig. 4, the RJ-0 and RJ-2 are traditional cross-spring rotational joints, while both the RJ-1 and the RJ-3 are double cross-spring joints. Each of the two joints, RJ-1 and RJ-3, consists of two traditional cross-spring rotational joints in series with a common rotational axis and a secondary stage (encircled in Fig. 4). This serial arrangement creates symmetry and allows for greater motion and less stress in the mechanism. It should be mentioned that using these joints can allow large-amplitude motions compared to notch joints, which thus serves for illustrating the present concept easily. Note that these joints are not as compact and simple (with manufacture and precise issues) as circular notch joints. In addition, the parasitic rotational shift of these joints will be minimized if the beams intersect at an appropriate position of their length [START_REF] Henein | The art of flexure mechanism design[END_REF].

We specify that the Bar-0 is fixed to the ground and the Bar-2 is the output motion stage, also named coupler. Bar-1, Bar-2 and Bar-3 correspond to links, CD, BC, and AB, respectively, in Figure 3. The link length can be expressed as l=LB+LR (1) where LB and LR are the lengths of each bar block and each compliant rotational joint, respectively, as indicated in Fig. 4.

Like the rigid-body four-bar mechanism shown in Fig. 3(a), the output motion stage (Bar-2) of the compliant four-bar mechanism has multiple operation modes under two rotational actuations (controlled by two input angles α and β for Bar-1 and Bar-3, respectively), as shown in Fig. 4. However, the compliant four-bar mechanism has more operation modes than its rigid counterpart as discussed below. A moving coordinate system (o-xyz) is defined in Fig. 4, which is located on Bar-2 (link BC), which coincides with the fixed frame (O-XYZ) in the singularity position. Based on this assumption, Bar-2's operation modes of the compliant fourbar mechanism are listed below: a) Operation mode I: Rotation in the XY-plane about the Axis-L, when α ≠ 0 and β = 0. b) Operation mode II: Rotation in the XY-plane about the Axis-R when α = 0 and β≠0. In this mode, the cross-spring joints can tolerate this constrained configuration (close enough to the singularity) due to the induced small elastic extension of joints, but do not work as ideal revolute joints anymore. c) Operation mode III: General rotation in the XY-plane about other axes except the Axis-L and Axis-R, when α≠β. Similar to the constraint in operation mode II, the cross-spring joints in this mode are no longer working as ideal revolute joints. d) Operation mode IV: Pure translational motions in the XY-plane along the X-and Y-axes (mainly along the Y-axis), when α=β.

It is noted that the rotational axes associated with α and β are both fixed axes (as indicated by solid lines in Fig. 4); while Axis-L and Axis-R (as indicated by dashed lines in Fig. 4) are both mobile axes. The rotational axis of α is the rotational axis of joint RJ-0, and the rotational axis of β is the rotational axis of joint RJ-3. Axis-L is the rotational axis of joint RJ-2, which moves as Bar-3 rotates. Axis-R is the rotational axis of joint RJ-1, which moves as Bar-1 rotates. As shown in Fig. 3, in the initial singular configuration, Axis-L overlaps with the axis of α; and Axis-R lies in the plane spanned by Axis-L and the Axis of β.

It should be also pointed out that it is possible for operation modes II and III having large-range motion thanks to the particular use of rotational joints which may not be true for circular notch joints anymore.

These operation modes are also highlighted in Fig. 5 with verification through the printed prototype. In order to simplify the analysis, let α and β be non-negative in Fig. 5. The primary motions of output motion stage (Bar-2) are the rotation in the XY plane and the translations along the X-and Y-axes; while the rotations in the XZ and YZ planes and translational motion along the Z-axis are the parasitic motions that are not the interest of this paper.

Kinematic models

The kinematics of the compliant four-bar mechanism is discussed as follows, under the assumption of small angles (i.e., close to the constraint singularity). According to the definition of the location and orientation of the Bar-2 (link BC) with respect to the fixed frame, we can have the primary displacement of the Bar-2 as:

1) Displacement of the centre of Bar-2 along the Y-axis: b-0=l(sin α+sin β)/2 ≈ l(α+β)/2 if small input angles (2) 2) Rotation of Bar-2 about the Z-axis:

ϕ-0= α-β (3)
Using the assumption of small angles, the displacement of the centre of Bar-2 along the X-axis is normally in the second order of magnitude of the rotational angles, which is trivial and can be neglected in this paper. Note that this trivial displacement is also affected by the centre drift of the compliant rotational joints [START_REF] Zhao | A novel compliant linear-motion mechanism based on parasitic motion compensation[END_REF].

Design of a multi-mode compliant gripper 3.1 Compliant gripper with multiple modes

In this section, a multi-mode compliant gripper using the compliant four-bar mechanism presented in Fig. 4 as a gripper jaw mechanism is proposed (shown in Fig. 7). Instead of the cross-spring joints in the compliant four-bar mechanism, the commonly-used rectangular short beams as the rotational joints (with rotation axis approximately in the centre) are adopted for the final gripper designed in this paper, as shown in Fig. 7(a). The reason for using the rectangular short beams is mainly twofold. Firstly, comparing with the cross-spring joints, the rectangular joints are compact, simple enough, and easy to fabricate. Secondly, the rectangular joints have larger motion range than that of the circular notch joints (as used in Appendix A). In addition, using the rectangular short-beams allow not to work as pure rotational joints as discussed in Section 2.1.

In order the make the whole mechanism more compact, the compliant gripper is a two-layer structure with two linear actuators to control the two rotational displacements (α and β) in each jaw. The top layer actuator is for determining β, and the bottom layer is for determining α. The design of the compliant gripper is further explained in Fig. 8, with all dominant geometrical parameters labelled except the identical out-of-plane thickness (u) of each layer, and the gap (g) between the two layers.

As can be seen in Figs.

Kinetostatic modelling

Under the assumption of small rotations, the relationship between the linear actuation and the rotational actuation in the slider-crank mechanism (the left jaw is taken for studying) can be modelled below:

bb or / a r a r

    ( 8
)
tt or / a r a r     (9) where at and ab represent the input displacements of the top and bottom actuators along the X-axis, respectively. A minus sign means that the positive linear actuation causes a negative rotational actuation (based on the coordinate system illustrated in Fig. 8). Here, r is the lever arm as shown in Fig. 8.

The rotational displacement of RJ-4 in the added slider-crank mechanism can be approximately obtained as follows. The rotational displacement of RJ-5 in each layer can be ignored due to the specific configuration of the added slider-crank mechanism as shown in Fig. 8, where the crank parallel to the Y-axis is perpendicular to the coupler so that the coupler is approximately straight over motion under the condition of small rotations. 2)-( 9), the input-output kinematic equations of the compliant gripper can be obtained:

tb () 2 l b a a r   ( 12
) tb aa r    (13)
As indicated in Eqs. ( 12) and ( 13), the amplification ratio is a function of design parameter r denoted in Fig. 8. Using the above kinematic equations, the kinetostatic models of the compliant gripper can be derived from the principle of virtual work [START_REF] Howell | Compliant Mechanisms[END_REF], with at and ab being the generalised coordinates.

t t b b t b tb d d d d UU F a F a a a aa      (14)
where Ft and Fb represent the actuation forces of the top and bottom linear actuators along the X-axis, corresponding to at and ab, respectively. U is the total elastic potential energy of the compliant gripper, which is calculated as below:

2 2 2 2 2 2 2 0 0 1 1 2 2 3 3 4 4t 4 4b p t p b 2 2 2 2 2 2 b t b t t t b 0 1 2 3 4 4 22 p t p b 1 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2 1 1 1 1 1 1 ( ) ( ) ( 2 ) ( ) ( ) ( ) 2 2 2 2 2 2 11 22
U k k k k k k k a k a a a a a a a a k k k k k k r r r r r r r k a k a                       (15)
where k0, k1, k2, k3 correspond to the rotational stiffnesses of RJ-0, RJ-1, RJ-2, RJ-3 in the compliant four-bar mechanism, respectively. k4 is the rotational stiffness of the RJ-4 in each layer. kp is the translational stiffness of the prismatic joint in each layer.

Note that the reaction forces from gripping objects [START_REF] Hao | Design and static testing of a compact distributed-compliance gripper based on flexure motion[END_REF] can be included in Eq. ( 14), which, however, is not considered in this paper.

Combining the results of Eqs. ( 14) and ( 15) with the substitution of Eqs. ( 4)-( 11), we have

t b t t t 1 2 3 4 p t 1 2 3 4 2 t p b 2 2 2 2 t t 2 t 1 2 1 1 ( ) ( 2 )( ) ( ) ( ) 2 8 2 2 4 ( 2 ) 2 /2 (
) / a a a a a k k k k k a r r r r r r r r r kk U F a k k k a k a r r r F rr                  (16)
b b t b 0 2 4 p b 0 2 4 2 b p t 2 2 2 2 b b b 1 1 1 ( ) ( 2 ) ( ) 2 2 2 4 ( 2 ) ( /
) 2 /2 a a a a k k k k a r r r r r r r k k k k a k a r r r r U F a F             (17)
Equations ( 16) and ( 17) can determine the required forces for given input displacements, which can be rearranged in a matrix form:

t 11 12 b 21 2 b 2 t a kk a kk F F                 (18)
where the stiffness coefficients of the system associated with the input forces and input displacements are

k k k k kk r r r r k kk r k k k kk r r r           
Therefore, the input displacements can be represented with regard to the input forces as:

FF c c k k FF c c k a a k                                 (19)
We can further obtain the following stiffness equations for all compliant joints used in this paper [START_REF] Hao | Designing a monolithic tip-tilt-piston flexure manipulator[END_REF], which can be substituted into Eqs. ( 18) and ( 19) to solve the load-displacement equations.

EI Eut k k k ll EI Eut k k k ll EI Eut k ll          ( 20
)
where E is the Young's modulus of the material and I is the second moment of inertia of the cross-section areas.

With the help of Eqs. ( 12), ( 13) and ( 19), we can have the required output displacements for given input displacements/forces:

                                                  (21)
Note that in the above kinetostatic modelling, a linear assumption is made. However, in order to capture accurate load-dependent kinetostatic characteristics for larger range of motion, nonlinear analytical models should be used [START_REF] Ma | Modeling large planar deflections of flexible beams in compliant mechanisms using chained beam-constraint-Model[END_REF].

Case study

In this section, a case study with assigned parameters as shown in Finite element analysis (FEA) simulation was carried out to show the four operation modes of the compliant gripper, in comparison to the 3D printed prototype (Fig. 9). Here, Solidworks 2017, with a meshing size of 1.0 mm and other settings, in default is used for FEA. Figure 10 illustrates the comprehensive kinetostatic analysis results of the proposed compliant gripper including the comparison between the analytical modelling and FEA. It can be observed that linear relations of all figures have been revealed where lines in either model are parallel each other. The FEA results have the same changing trends as the analytical models, but deviate from the analytical models in certain degrees. The discrepancy between the two theoretical models may be due to the assumptions used in the analytical modelling such as neglecting centre drift of rotational joints. It should be pointed out that although stress analysis is not the interest of this paper, the maximal stress (29 MPa) was checked in FEA, which is much less than the yield strength of the material during all simulations for the case study.

In Fig. 10(a), with the increase of ab, the difference between the two models (analytical and FEA) goes up if at=0, while the difference between the two models decreases if at=0.50 mm and 1.0 mm. This is because of different line slops of two models. Generally speaking, the larger at, the larger the deviation of two models, where the maximal difference in Fig. 10(a) is about 20%. In Fig. 10(b), the line slops of two models are almost same, meaning that the increase of increase of ab has no influence of the discrepancy of two models for any value of at. Also, the larger at, the larger the deviation of two theoretical models. A real prototype made of polycarbonate was fabricated using CNC milling machining, which is shown in Fig. 11(a). Each layer of the compliant gripper (Fig. 7) was made at first and then two layers were assembled together. The gripper prototype was tested in a customised-built testing rig as shown in Fig. 11(b). The singleaxis force loading for both the top layer actuation and bottom layer actuation was implemented. The micrometer loads displacement on the force sensor that is directly contacted with the top layer input or bottom layer input. The force sensor reads the force exerted on one input of the gripper, and the two displacement sensors indicate the displacements of two inputs. Testing results are illustrated in Fig. 12, which are compared with analytical models and FEA results. It is shown that the analytical displacement result is slightly larger than the experimental model, but is slightly lower than the FEA result. The difference among three models is also reasonable, given that FEA always take all parts as elastic bodies (i.e., less rigid system) and testing is imposed on a prototype with fabricated fillets (i.e., more rigid system).

Conclusions

In this paper, we present the first piece of work that employs a constraint singularity of a planar equilateral four-bar linkage to design a reconfigurable compliant gripper with multiple operation modes. The analytical kinetostatic mode of the multi-mode compliant gripper has been derived, which is verified in the case study. It shows that the FEA results comply with the analytical models with acceptable discrepancy. The proposed gripper is expected to be applied in extensive applications such as grasping a variety of shapes or adapting to specific requirements.

The design introduced in this paper uses a two-layer configuration for desirable compactness, which adversely results in non-desired small out-of-plane rotations. However, the out-of-plane rotations can be reduced by optimising currently-used compliant joints or employing different compliant joints with higher out-of-plane stiffness. Note that the compliant gripper can be designed in a single layer for monolithic fabrication and elimination of out-of-plane motion, at the cost of a larger footprint. This can be done by using the remote rotation centre, as shown in Appendix A. A single layer gripper can be more easily designed at micro-scale on a silicon layer for MEMS devices.

Despite the work mentioned above, there are other aspects to be considered in the future, including but not limited to the following: (i) An analytical nonlinear model for a more accurate large-range kinetostatic modelling of the compliant gripper; (ii) Design optimisation the compliant gripper based on a specific application; (iii) Output testing with comparison to the analytical model for a specific application; and (iv) Developing a control system to robotise the compliant gripper.

  (a) Resolution of a jaw: ∆Jaw (b) Diameter of micro object: Dmicro (c) Diameter of brittle object: Dbrittle, with a small breaking deformation ∆b ∆Jaw> Dmicro/2 ∆Jaw> ∆b/2 (but Dbrittle can be larger than 2∆Jaw)

Figure 2 :

 2 Figure 2: A planar equilateral four-bar linkage

Figure 3 :

 3 Figure 3: Constraint singular configuration of the planar equilateral four-bar linkage

Figure 4 :Figure 5 :

 45 Figure 4: A compliant four-bar mechanism at its constraint singular position (as fabricated)

Figure 6 :

 6 Figure 6: The generic kinematic configuration (close to constraint singularity) of the four-bar linkage

Figure 7 :

 7 Figure 7: The synthesized multi-mode compliant gripper

Figure 8 :

 8 Figure 8: Design details of the multi-mode compliant gripper

  (a) Top layer with two slider-crank mechanisms (indicated by dashed square) (b) Bottom layer with two slider-crank mechanisms (indicated by dashed square) bottom layer

Figure 10 (

 10 c) reveals the same conclusion as that of Fig. 10(b).Figure 10(d) shows the similar finding to that of Fig. 10(a), except the larger at, the smaller the deviation of two models. It is clearly shown that in Figs. 10(b), 10(c) and 10(d) the general discrepancy of the two models is much lower than that in Fig. 10(a).

Figure 9 :Figure 12 .

 912 Figure 9: Gripper operation modes under input displacement control

  (a) Monolithic design of the reconfigurable compliant gripper (b) Actuating the linear actuator 1 only in bi-directions

Table 1 : Geometrical parameters

 1 Table 1 is presented to verify the analytical models in Section 3.2. The overall nominal dimension of the compliant gripper is 130 mm × 70 mm. The Young's of compliant gripper is given by E=2.4G Pa, which corresponds to the material of Polycarbonate with Yield Strength of σs > 60 MPa, and Poisson Ratio of v= 0.38.

	l	t	r	h	l1	l2	w1	w2	u	g
	25 mm	1 mm	18 mm	25 mm	5 mm	15 mm	24 mm	19 mm	10 mm	3 mm

Acknowledgment

The author would like to thank Mr. Tim Power and Mr. Mike O'Shea from University College Cork for the excellent prototype fabrication work as well as their kind assistance in the experimental testing. The work in this paper was funded by IRC Ulysses 2015/2016 grant.

In order to reach the constraint singularity, i.e. allowing the overlapping of two revolute joints, the design uses isosceles trapezoidal flexure joints with remote center of rotation. However, this one layer mechanism has a quite large footprint, and requires more extra space for incorporating the two actuators. Moreover, due to the use of circular notch joints, operation modes II and III in this gripper may not produce large-range motion. The modelling and analysis of the present monolithic design in this appendix is left for future study.

(e) Fabricated prototype


Aymen Sendi
email: aymen.sendi@laas.fr

Grérory Besnard
email: gbesnard@laas.fr

Philippe Menini
email: menini@laas.fr

Chaabane Talhi
email: talhi@laas.fr

Frédéric Blanc
email: blanc@laas.fr

Bernard Franc
email: bfranc@laas.fr

Myrtil Kahn
email: myrtil.kahn@lcc-toulouse.fr

Katia Fajerwerg
email: katia.fajerwerg@lcc-toulouse.fr

Pierre Fau
email: pierre.fau@lcc-toulouse.fr

Sub-ppm Nitrogen Dioxide (NO 2 ) Sensor Based on Inkjet Printed CuO on Microhotplate with a Pulsed Temperature Modulation

Keywords: NO 2, CuO nanoparticles, temperature modulation, gas sensor, selectivity. I

Nitrogen dioxide (NO 2 ), a toxic oxidizing gas, is considered among the main pollutants found in atmosphere and indoor air as well. Since long-term or short-term exposure to this gas is deleterious for human health, its detection is an urgent need that requires the development of efficient and cost effective methods and techniques. In this context, copper oxide (CuO) is a good candidate that is sensitive and selective for NO 2 at sub-ppm concentrations. In this work, CuO nanoparticles have been deposited by inkjet printing technology on a micro hotplate that can be operated up to 500°C at low power consumption (55 mW). The optimum detection capacity is obtained thanks to a temperature modulation (two -consecutive temperature steps from 100°C to 500°C), where the sensing resistance is measured. Thanks to this operating mode, we report in this study a very simple method for data processing and exploitation in order to obtain a good selectivity for the nitrogen dioxide over few interferent gases. Only four parameters from the sensor response allow us to make an efficient discrimination between individual or mixed gases in humid atmosphere.

INTRODUCTION

Humans spend more than 90% of their time in closed environments, even though this indoor environment offers a wide variety of pollutants [START_REF] Lévesque | Indoor air quality[END_REF] [START_REF] Namiesnik | Pollutants, their sources, and concentration levels[END_REF].

Indoor air pollution is a real health threat, so measuring indoor air quality is important for protecting the health from chemical and gaseous contaminants. Nitrogen dioxide (NO 2 ) is a dangerous pulmonary irritant [START_REF] Lévesque | Indoor air quality[END_REF]. NO 2 is generated by multiple sources of combustion in indoor air, such as smoking and heaters, but it also comes from outside air (industrial sources, road traffic) [START_REF] Cadiergues | Indoor air quality[END_REF]. NO 2 may have adverse effects of shortness of breath, asthma attacks and bronchial obstructions [START_REF] Koistinen | The INDEX project: executive summary of a European Union project on indoor air pollutants[END_REF]. It is also classified as toxic by the "International Agency for Research on Cancer (IARC)" [START_REF] Loomis | The International Agency for Research on Cancer (IARC) evaluation of the carcinogenicity of outdoor air pollution: focus on China[END_REF], hence the necessity for sensor development for accurate NO 2 detection is an acute need. Among sensors techniques, the metal oxide gas (MOX) sensors are promising candidates because of their high performance in terms of sensitivity on top of their low production cost. The copper oxide (CuO) material is highly studied because of its high sensitivity and its ability to detect oxidant gaseous compounds, but also for other indoor air pollutants, such as acetaldehyde (C 2 H 4 O), formaldehyde (CH 2 O), NO 2 , CO, etc. However, CuO suffers from a major disadvantage which is the lack of selectivity with respect to targeted gas.

In this study, our main objective is to develop an innovative and simple pulsed-temperature operating mode associated with an efficient data processing technique, which enables good selectivity toward NO 2 in gas mixtures. This technique is based on few parameters extracted from the dynamic response of sensor versus temperature changes in a gaseous environment. These parameters are: the normalized sensing resistance, the values of the slope at the origin, the intermediate slope and the final slope of the response of NO 2 against different reference gases, such as C 2 H 4 O, CH 2 O and moist air. The selectivity of NO 2 was examined in relation to air moisture with 30% humidity, C 2 H 4 O at 0.5-ppm, CH 2 O at a concentration of 0.5-ppm and the binary mixture of these gases with 0.3-ppm of each.

In Section II of the paper, we describe the materials and methods used in our work. Section III presents our results and the discussion. We conclude this work in Section IV.

II. MATERIALS AND METHODS

The sensitive layer made of CuO nanoparticles is deposited by inkjet printing on a silicon microhotplate [START_REF] Dufour | Technological improvements of a metal oxide gas multi-sensor based on a micro-hotplate structure and inkjet deposition for an automotive air quality sensor application[END_REF]. The ink is prepared with 5% CuO weight, which was dispersed in ethylene glycol by an ultrasonic bath for about one hour. The dispersions obtained were allowed to settle for 24 hours. The final ink was collected and then used for printing using Altadrop equipment control, where the numbers of the deposited drops of ink were controlled [START_REF] Dufour | Technological improvements of a metal oxide gas multi-sensor based on a micro-hotplate structure and inkjet deposition for an automotive air quality sensor application[END_REF]. This technique is simple and allowed us to obtain reproducible layers thicknesses of a few micrometers depending on the number of deposited drops. In addition, this technique permits to have a precisely localized deposit without need of additional complex photolithographic steps [START_REF] Morozova | Thin TiO2 films prepared by inkjet printing of the reverse micelles sol-gel composition[END_REF]. The CuO layer is finally annealed in ambient air from room temperature to 500°C (rate 1°C/min) followed by a plateau at 500°C for 1 hour before cooling to room temperature (1°C/min). This initial temperature treatment is necessary because CuO requires operating temperatures between 100°C<T<500°C. The thermal pretreatment is necessary to generate ionized oxygen species in atomic or molecular form at the oxide surface and therefore to improve the reactivity between the reacting gas and the sensor surface [START_REF] Dufour | Technological improvements of a metal oxide gas multi-sensor based on a micro-hotplate structure and inkjet deposition for an automotive air quality sensor application[END_REF].

In this study, we have used a pulsed temperature profile, presented in a previously published work, which showed that optimized sensitivity can be achieved with the use of two different temperature stages at 100°C and 500°C respectively. This dual temperature protocol also reduces the total power consumption of the device (see Figure 1). The CuO sensor was placed in a 250 ml test chamber and the test conditions were as follow: -A flow rate of 200 ml/min, controlled by a digital flowmeter.

-A relative humidity (RH) level of 30% is obtained by bubbling synthetic air flow controlled by a mass flow controller.

-The measuring chamber is at ambient temperature, controlled by a temperature sensor placed inside the vessel.

-A bias current is applied to the sensitive layer, controlled by a Source Measure Unit (SMU).

We started a test with single gas injections after a phase of two-hours stabilization in humid air, then with injections of binary mixtures for 16-min. During 32-min, moist air is injected between two successive gas injections. This time is enough to clean the chamber and stabilize the sensor to its baseline.

The gas injections concentrations are summarized in Table 1. A schematic representation of these injections is presented in Figure 2. During all the experience (stabilization phase, gases injections and stage between two successive gases) , the sensor is powered by a square signal voltage applied on the heater in order to obtain two temperature steps as shown in Figure 1. To ensure a constant overall flow, we adapted the gas injection sequences duration in correlation with the heating signal period (see Figure 3).

The resistance variation of CuO is measured under a fixed supply current of 100 nA in order to obtain voltage measurements in the range of few volts, far from compliance limit of the measuring device (20V). We also verified that this procedure (temperature cycling) doesn't affect the sensor reproducibility in terms of baseline or the sensor sensitivity.

Under such test conditions, we achieved a continuous 6.5 hours testing period without observing any drift on the raw sensor response. The sampling period is 500 ms, which gives us 60 points on a 30-second response step, this acquisition rate being enough for accurate data processing. Finally, we analyzed the sensor responses at each steps, according to the different gas injections, using a simple method of data processing in order to have the better selectivity of NO 2 .

III. RESULTS AND DISCUSSION

A.

Method of analysis During each gas injection, 16 periods of temperature modulation are applied. After verifying the reproducibility of sensor responses along these cycles, we only present here the responses of the last cycle, which is stabilized and reproducible from one cycle to another. As mentioned previously, we used new simple data treatment methods to obtain the better selectivity toward NO 2 with respect to several interferent gases. Among the multiple possible criteria, we chose representative variables that take into account the dynamic sensor behavior during a change of gaseous conditions and during a pulsed temperature; these criteria are obtained from the sensor resistance slopes during the gas response on each 30-second-steps. The data acquisition relies on the decomposition of the response into three distinct domains (see Figure 4):

-Starting Slope: from the 1st point to the 10th point (in yellow), -Intermediate slope: from the 10th point to the 30th point (in red), -Final slope: from the 30th point to the 60th point (in black).

The normalized resistance is measured from the last point on each step: this is the absolute difference between the resistance of the sensor under a reference gas (like moist air) and the resistance of the sensor under targeted gas(es), at the final cycle of each injection: R n = ((R gas -R air ) / R air ) * 100

(1) By treating these four parameters, a good selectivity of NO 2 can be obtained with respect to moist air, C 2 H 4 O and CH 2 O.

B.

Slope at the origin The slope at the origin is calculated on the first 10 points of each temperature step of the 8th cycle of each injection. Regarding the reference gas (moist air); we took the response of the 8th cycle of the last sequence under humid air before the injection of gases. The values of these slopes (in Ohms/ms) are shown in Figure 5; each bar represents the value of the slope at the origin of each gas at 500 and 100°C. Figure 5 clearly shows that the calculations of this parameter enable us to differentiate NO 2 from the other reference gases by measures on the plateau at 100°C. Regarding the other step at 500°C we note that the response is almost zero, because the transition from a cold to a hot state decreases the detection sensitivity and therefore reduces the sensors resistance variations under gas. We also note that with this criterion we evidence a significant difference between the value of the starting slope under NO 2 compare to others under other gases and mixtures without NO 2 .

C.

Intermediate slope The intermediate slope is calculated from the 10th point to the 30th point of each temperature step of the last cycle during each gas injection and compared with the value under air. The calculated values are shown in Figure 6. According to Figure 6, the value of the intermediate slope of individual NO 2 injection or of the gas mixtures in which there is NO 2 , is not prominent compared with the other reference gases even on the plateau at 100°C. This parameter is less effective than the previous one to detect NO 2 in gas mixtures.

D.

Final slope The final slope is the slope of the second half of the gas response; it is calculated between the 30th point and the 60th point. The response to the different temperature steps of the last cycle during NO 2 injection and the reference gases is presented in Figure 7. It is worth noting that the value of the final slopes of NO 2 or gaseous mixtures, which contain the NO 2 , are very different compared to the other reference gases on the two stages at 100°C only. This parameter allows us to select NO 2 with respect to other interfering gases.

E.

Normalized resistance As previously presented, the normalized resistance is calculated with respect to humid air from each last cycle level. The reference resistance used is the resistance of each stage of the last wet air sequence before the gas injection. The results obtained from these calculations are shown in Figure 8. Figure 8 shows a slight variation in the values of the normalized resistance between two similar and successive temperatures for the same gases. This slight variation can be explained by data dispersion which is +/-2%, due to the fact that the normalized resistance is calculated from the raw values during the gas injection and the raw values of resistances during the injections of humid air that may be slightly different between two similar and successive temperature stages.

The CuO sensor response to sub-ppm NO 2 levels when injected individually or in combination with another gas, is specific when measurements are taken on the low temperature plateau at 100°C.

IV. CONCLUSION

The selectivity and sensitivity of our CuO sensor has been studied by different operating modes and simple methods of analysis. Specific temperature modulation was applied to the metal oxide with the use of temperature steps at 500 and 100°C. The response of CuO sensitive layer toward gases representatives of indoor air pollution (C 2 H 4 O, CH 2 O, NO 2 , humid air) has been studied. These responses were analysed with several parameters, such as the study of the slope of resistance variation at the origin, the intermediate slope, the final slope and the normalized resistance measured at each temperature steps. The study of these different parameters shows that the CuO material is able to detect sub-ppm levels of NO 2 with a good selectivity compared to different interfering gases. To still improve the selectivity of gas sensor device to a larger variety of polluting gases, we plan to integrate these CuO sensors in a multichip system, which will allow us to use in parallel new metal oxide layers with specific temperature profiles and data analysis criteria.

Figure 1 .

 1 Figure 1. CuO temperature profile.

TABLE 1 .Figure 2 .

 12 Figure 2. Synoptic representative of a sequence of gas injections.

Figure 3 .

 3 Figure 3. Diagram of a gas sequence.

Figure 4 .

 4 Figure 4. Representative diagram of a response of a gas sensor during an injection cycle and showing the 3 domain slopes.

Figure 5 .

 5 Figure 5. Representation of the slopes at the origin under different gases of the 8th cycle in "Ohms/ms" according to the temperature of the sensors.

Figure 6 .

 6 Figure 6. Representation of the intermediate slopes under different gases at the 8th cycle in "Ohms/ms" according to the temperature of the sensors.

Figure 7 .

 7 Figure 7. Representation of the final slopes under different gases at the 8th cycle in "Ohms/ms" according to the temperature of the sensors.

Figure 8 .

 8 Figure 8. The normalized resistance of the different gas injections at the 8th cycle according to the temperature.

Copyright (c) IARIA, 2018. ISBN: 978-1-61208-621-7 ALLSENSORS 2018 : The Third International Conference on Advances in Sensors, Actuators, Metering and Sensing

ALLSENSORS 2018 : The Third International Conference on Advances in Sensors, Actuators, Metering and Sensing

Copyright (c) IARIA, 2018. ISBN: 978-1-61208-621-7

ACKNOWLEDGMENT

The authors express their gratitude to neOCampus, the university project of the Paul Sabatier University of Toulouse, for the financial support and the Chemical Coordination Laboratory of Toulouse for the preparation of the CuO nanoparticle powder. This work was also partly supported by the French RENATECH network.