Monica Vegliò
email: monica.veglio2@ensam.eu

Antoine Dazin

Olivier Roussette

Gérard Bois

FP17-2014-veglio PRESSURE MEASUREMENTS IN AN AXIAL COMPRESSOR: FROM DESIGN OPERATING CONDITIONS TO ROTATING STALL INCEPTION

The reduction of the environmental impact is nowadays the crucial challenge for the aeronautic industry. The following of a lower consumption of the vehicles has led to more compact and high loaded engines, increasing the internal flow unsteadiness and the occurrence of unstable phenomena, especially for compression stages. In this work the casing pressure characteristics in a single stage axial compressor both for normal-to-stall transient regime and for stalled conditions will be discussed. A special unsteady measurement window was designed, covering an entire rotor pitch, in the circumferential and in the axial extent, and allowing to solve pressure fields with high spatial resolution for several relative rotor/stator positions. A phase locked average technic was applied to obtain the pressure fields during normal and stalled operating conditions. The emergence of a spike-type pressure perturbation was captured during transient; the gradual alignment of the tip leakage vortex trajectory with the inlet rotor section was observed before the precursor appeared; this corresponds, following [1], at one of the criteria for spike stall inception machines. An original way to observe the rotating stall phenomenon is proposed; the phase-locked technic used allowed reconstructing the average pressure field in the perturbation reference frame. The salient rotating stall pattern features will be described.

INTRODUCTION

The reduction of the environmental impact is nowadays the crucial challenge for the aeronautic industry. The following of a lower consumption of the vehicles has led to more compact and high loaded engines, increasing the internal flow unsteadiness and the occurrence of unstable phenomena, especially for compression stages. A better understanding of the engines' behavior during transition from stable to unstable operating regimes should allow reducing the surge safety margin and finalizing control technics, delaying the instabilities appearance. A great effort has been done for 20 years in trying to unravel the details of stall onset processes. The multitude of stall inception patterns observed was classified by [START_REF] Camp | A Study of Spike and Modal Stall Phenomena in a Low-Speed Axial Compressor[END_REF] in two main groupings, spike-type and modal waves. Being the most common in real aeronautical compressor engines, spike-type precursor has attracted the interest of the scientific community. It is about a rapid perturbation of the flow near the casing, extending on few rotor pitches and rotating at 70-80% of rotor speed. Many studies have been carried out in order to understand which parameters mostly affect the emergence of this kind of precursor; Camp & Day [START_REF] Camp | A Study of Spike and Modal Stall Phenomena in a Low-Speed Axial Compressor[END_REF] have, for instance, underlined the importance of the incidence angle, whereas [START_REF] Dobrzynski | Influence of the rotor tip stagger angle on axial compressor inception[END_REF] focused on the effects of the stagger angle. Other works focused on the flow mechanisms leading to the emergence of this kind of perturbation. The experimental investigations of [START_REF] Deppe | Spike stall inception in axial flow compressors[END_REF] and the numerical simulations of [START_REF] Vo | Criteria for spike initiated rotating stall[END_REF] are in perfect agreement, identifying two main criteria for spike initiated rotating stall: the tip leakage vortex trajectory tends to align with the rotor inlet plane, whereas at the trailing edge a backward flow is observed. Following [START_REF] Deppe | Spike stall inception in axial flow compressors[END_REF], while the vortex trajectory progressively aligns with the leading edge plane the incoming flow spills to the next blade-to-blade channel. Yamada et al. [START_REF] Yamada | An Explanation for Flow Features of Spike-Type Stall Inception in an Axial Compressor Rotor[END_REF] studied, instead, spike flow pattern under particular conditions, "mild stall conditions"; they observed the emergence of a tornado-like vortex, connecting the blade suction side with the casing during spike evolution process. The more common set-up used in order to capture this kind of phenomena is to analyze pressure and/or velocity evolution for some discrete circumferential positions at the rotor inlet and outlet sections. More sophisticated measurement set-ups have also been used ( [START_REF] Yamada | An Explanation for Flow Features of Spike-Type Stall Inception in an Axial Compressor Rotor[END_REF], [START_REF] März | An Experimental and Numerical investigations Into th Mechanism of Rotating Instability[END_REF], [START_REF] Weichert | Detailed Measurements of Spike Formation in an Axial Compressor[END_REF]) allowing capturing the instantaneous pressure field pattern on the whole blade-to-blade channel with high time and space resolution. The aim of this work is to provide further information about the inception and the pattern of rotating instabilities. Unsteady pressure investigations were carried out in a low speed single stage axial compressor, using a specially designed measurement window. Combining frequency domain analysis and phase-locked averaging data treatment, the evolution of the pressure behavior from design conditions to stalled regime will be discussed. While the classical way to observe the instabilities involves instantaneous representations, the originality of this work consists in extracting the instability pattern, by depolluting it from all the other concurrent phenomena.

THE CME2 COMPRESSOR

The CME2, showed in Fig. 1, is a low speed single stage axial compressor, especially designed for instabilities investigations.

Figure 1 Compressor test rig

The operating point during stalled conditions remains effectively stable; although an unsteady phenomenon is installed, the average values of the flow rate and of the pressure rise are constant. The main features of the test rig are listed in Tab.1. The experiments reported in this paper were realized at a rotational speed of 3200 rpm. In a general way, the desired operating point was reached by throttling the vane. Furthermore, the method used to stall the compressor involved non-stop slowly throttling beyond the stability limit until rotating stall was reached.

INSTRUMENTATION

All the results presented in this work concern unsteady pressure investigations on the casing. A pressure field measurement window was specially designed, giving high spatial resolution and allowing solving phase-locked average for several rotor/stator relative positions. 98 measuring holes are regularly dispatched on the window, covering an entire rotor pitch for an axial extension of 78 mm (see details in Fig. 2). The two possible mounting positions of the device allow doubling the amount of measuring locations (in Fig. 2 green dots are consistent with one window position and the red ones with the same window turned of 180°).

Figure 2 Measurement grid

Four miniature high response pressure transducers (ENDEVCO 8507C-1/2) were used and moved to cover the green dotted grid. The transducers were mounted behind a small cavity; its resonance frequency is 7 times higher than the blade passing frequency (BPF). It was verified that the cavity did not produce any perturbation at low frequencies on the phenomena of interest. The sampling rate was varied following the test purpose: 35000Hz (21 samples per blade passage) for normal regime acquisitions and 50000Hz (30 samples per blade passage) for unsteady regime tests. A low-pass filter at 10000Hz was applied by the pressure transducer conditioner to avoid aliasing. The data were acquired by a National Instrument PXI 6123 module via a Labview™ code.

RESULTATS

Global Performance Curve.

Fig. 3 presents the performance characteristics of the compressor, for the whole stage (dots) and the single rotor (stars). The pressure rises were obtained as difference between the calming chamber (see Fig. 1) and the outlet sections, of the stator and the rotor respectively.

Figure 3 Performance characteristics of the compressor at 3200 rpm

A colour code distinguishes the loading and the unloading paths, highlighting the presence of a hysteresis cycle. Both performance curves obtained by throttling and opening the valve are divided in two portions:

-normal regime, positive slope region, for q throttling > 3.85kg/s and q opening >4.2kg/s; -stalled regime, negative slope region, for q throttling <3.2kg/s and q opening <3.85kg/s;

The transient between the two operating conditions, following either the loading or the unloading path, is really abrupt. Once the rotating phenomenon is installed, starting from about 3.4kg/s, the regime becomes stable again.

No intermediary pseudo-stable regions are observed. So, accordingly with [START_REF] Yamada | An Explanation for Flow Features of Spike-Type Stall Inception in an Axial Compressor Rotor[END_REF], mild stall conditions, with multiple rotating stall cells, should not take place. All the investigations on which is based this work were carried out by throttling the valve (black curves in Fig. 2)

The four operating flow rates surveyed are color marked in Fig. 3; starting from the right side: the design operating point q DP =5.06 kg/s, three partial flow rates, the first one close to the stability limit q PP =4.15 kg/s, the second during normal-to-stalled transition q TP =3.81 kg/s and the last one for completely developed stall q RSP =3.35 kg/s.

Phase-Locked Average Casing Pressure Field for Normal Regime

The data issued from all the locations marked in green in Fig. 2, for the two highest flow rates, were treated by means of a phase-locked average technic to reconstruct the casing pressure field. As already mentioned, one of the advantages of the measurement window in use is the possibility to calculate the pressure field for several rotor/stator relative positions. The results selection in Fig. 4 lets emerge the presence of a low pressure region near the blade suction side evolving in magnitude and position with the flow rate; it moves progressively backward from the axial position 20mm to the leading edge plane, while the flow rate is reducing from the design value toward the stability limit. This low pressure region is the track of the tip leakage vortex bursting zone.

According with [START_REF] Vo | Criteria for spike initiated rotating stall[END_REF], the alignment of the tip leakage vortex trajectory with the leading edge plane is a criterion for spike initiated rotating stall. As it will be presented in the next paragraph, a sudden spikelike perturbation actually appears during the normal to stall transient. The chance that the low pressure region observed is linked to the same phenomenon as for [START_REF] Yamada | An Explanation for Flow Features of Spike-Type Stall Inception in an Axial Compressor Rotor[END_REF], namely the casing leg of a tornado-like vortex, can be rejected. This kind of structure was observed [START_REF] Yamada | An Explanation for Flow Features of Spike-Type Stall Inception in an Axial Compressor Rotor[END_REF] during mild stall regime that is a weak multicells rotating stall pattern, which develops progressively between the normal and completely stalled conditions. As it can be noticed from Fig. 3, no intermediary pseudo-stable regions are observed in the performance curve of the CME2.

Spike-Type Rotating Stall Inception

The common way to investigate the flow during rotating stall onset is to perform simultaneous acquisitions at several locations dispatched on the inlet section circumference.

A different point of view was adopted in this work, trying to rationalize the placement of the measurement means on the available grid. The four pressure transducers were placed in pair, at the inlet and outlet rotor sections respectively, as marked in Fig. 2 (Circumferential locations 0°violet rings -and 12°-orange rings -for the axia l coordinates 0mm and 66mm).

The throttling valve was slowly closed, starting from the normal regime, passing through the stability limit to reach the perturbed operating conditions. The Fig. 5 focuses on the transient from normal to stalled regime, the time period observed corresponding to the operating point marked in dark blue in Fig. 3 (q TP =3.81kg/s). The pressure signal was smoothed by means of a 3 rd order Butterworth filter with a cut frequency of 250 Hz. While the spectrum analysis could not provide any information about the transient regime, the time pressure evolution (top box) clearly reveals the emergence of a perturbation before the stall develops, both at the inlet and the outlet sections (bottom boxes). The phenomenon appears approximately 5 rotor revolutions before the rotating stall is completely established (point A).

Figure 5 Spike stall inception

By convoluting the signals at the same axial position, it was calculated that the perturbation rotates in the same circumferential direction as the machine but with a speed ratio of 0.7 approximately. Using its angular velocity and its time duration, it was evaluated that the circumferential perturbation extent is about 6 rotor pitches. Its shape and all of the characteristics given above bring to consider the phenomenon as a spike-type stall precursor [START_REF] Camp | A Study of Spike and Modal Stall Phenomena in a Low-Speed Axial Compressor[END_REF]. As already mentioned, a pressure signal perturbation is also seen at the outlet section. It is slightly delayed and it presents an inverse pattern with respect to the leading edge side perturbation. Future works, including finer casing pressure investigations, coupled with hot-wire anemometry at the inlet and outlet sections, should give a better understanding of the flow pattern during the transient.

Stalled Operating Conditions.

Rotating stall pattern for the test rig under survey will be presented in this paragraph. The results discussed concern the operating point marked in light blue on the performance curve in Fig. 3 (q RSP =3.35kg/s). It is remembered that during stalled regime the compressor operating point is stable. The same transducers arrangement as for transient investigations (previous paragraph) was used to get the main features of the unsteady phenomenon. The Fast Fourier Transform was calculated for the four locations, by 50000-sample Hanning windows corresponding to 1 second acquisition. Fig. 6 presents the results for the grid location 0mm-0°.

Figure 6 Frequency domain analysis, q RSP =3.35kg/s, transducer location 0mm-0°

Firstly, it can be observed that the analysis correctly captures the blade passage for a frequency of 1610Hz. A detail of the low frequency range is also given in Fig. 6. Two peaks are emerging, the lowest one (f=21Hz) due to the rotating stall phenomenon (RSF=rotating stall frequency), the second being its first harmonic (1 st HRSF). The cross-spectrum calculated between the two circumferential locations 0° and 12°, both for the inlet and the outlet sections (axial positions 0mm and 66mm), allows concluding that the rotating stall observed is characterized by a single cell pattern rotating at a velocity which is 40% of the rotor speed. st HRSF location used, average pressure field could not be solved with respect to the relative rotor/stator position. The field is expressed in a reference system linked to the rotating stall cell, the compressor rotational direction being represented by the arrow in the bottom side. Three main regions can be distinguished in Fig. 7, characterized by different circumferential extent and pressure gradient.

The completely stalled behavior is recognizable in region β, where no pressure gradient exists between the outlet and inlet sections. The pressure field is not completely homogenous on the whole of this no-load region which extends on 6 rotor pitches; a phase of lower pressure stands out (dark blue colored in region β), presenting two cores placed upstream and downstream of the rotor wheel.

The stalled flow region is preceded by a long nonperturbed phase (region γ in Fig. 7), extending on 17 rotor pitches. The pressure gradient corresponds to the average rotor pressure rise for the operating flow rate investigated (q RSP ).

A very high loaded sector (α in Fig. 7) follows the stalling region. It extends on almost 7 rotor pitches and is characterized by a pressure gradient higher than 2 times the average rotor pressure rise expected for the regime. While during normal operating regime the whole circumference is on average submitted to the same pressure conditions, an important redistribution of the load takes place during stall. It seems to be reasonable that a circumferential redistribution of the flow rate takes place in the same time as the load. A stronger flow rate passing across region α could explain the improved operating conditions: the better flow incidence would be responsible of losses reduction.

CONCLUSIONS AND PERSPECTIVE

The need of a better understanding of the flow behaviour during rotating stall regime and its onset phase motivated this work. A special measurement window was designed in order to perform fine casing pressure investigations on an entire blade to blade channel, both in the circumferential and axial directions. Phase-locked average data treatment was performed for normal and stalled regimes. The pressure fields for normal operating conditions let emerge the evolution of the tip leakage vortex trajectory while the flow rate was reducing. This phenomenon was identified as one of the criteria proposed by [START_REF] Vo | Criteria for spike initiated rotating stall[END_REF] for spike-stall inception compressors. The emergence of a spike-stall precursor during normal to stall transition was actually confirmed by the time analysis of the pressure signals at inlet rotor section. A localized perturbation, extending on 6 rotor pitches, was rotating in circumferential direction at 70% of the rotational shaft speed. Simultaneous acquisitions at the outlet axial section showed the presence of a concurrent perturbation, with similar behaviour in speed and extent but with an inverse pattern. The limited circumferential extent (due to the window dimension) of the present measurement region does not allow observing the phenomenon all along its evolution. A new set-up with several circumferentially dispersed measurement locations will be realized. Furthermore, instantaneous and phase-locked average pressure field will be also calculated. Coupled with hot-wire anemometry explorations they will provide another point of view on the flow pattern during precursor emergence. The original approach used to characterize the rotating stall allowed representing the average pressure field throughout a whole stall cycle. A single stalled flow cell, rotating with an angular speed equal to 40% of the shaft speed, was extending on 6 rotor pitches. was extending on 17 rotor pitches and its average load was in agreement with the average pressure rise for the stalled operating conditions investigated. -The second one was more localized, occupying only 7 rotor pitches, and presented a very high load conditions. It is supposed that this particular configuration of the pressure load along the circumference is connected with a redistribution of the flow rate. Further investigation, involving hot wire anemometry measurements at the rotor inlet and outlet sections, will be carried out to verify this statement.

They will provide a fine characterization of the circumferential and radial distribution of the flow during one stall period.

Figure 4

 4 Figure 4 Phase-averaged pressure fields for normal operating conditions

Figure 7

 7 Figure 7 Average casing pressure field during rotating stall More details about the flow structure near the casing can be obtained by observing Fig.7. The color map was calculated from the data for one circumferential position in the measurement grid: the transducers were moved to cover all the axial locations for the line 12° in Fig.2. A phase-locked average technic was applied, similarly to what it had been done for normal operating conditions. Due to single circumferential

ACKNOWLEDGEMENTS

The authors would like to thank the groups Snecma and Turbomeca for their financial support as part of the C.I.R.T. convention (Consortium Industrie Recherche en Turbomachines).


Myriam Gharbi

Patrice Peyre
email: patrice.peyre@ensam.eu

Cyril Gorny

Muriel Carin

Simon Morville

Philippe Le Masson

Denis Carron

Rémy Fabbro

Influence of a pulsed laser regime on surface finish induced by the direct metal deposition process on a Ti64 alloy

Keywords:

The direct metal deposition (DMD) laser technique is a free-form metal deposition process, which allows generating a prototype or small series of near net-shape structures. Despite numerous advantages, one of the most critical issues of the technique is that produced pieces have a deleterious surface finish which requires post machining steps. Following recent investigations where the use of laser pulses instead of a continuous regime was successful to obtain smoother DMD structures, this paper relates investigations on the influence of a pulsed laser regime on the surface finish induced by DMD on a widely used titanium alloy (Ti64). Findings confirm that using high mean powers improves surface finish but also indicate a specific effect of the laser operating mode: using a quasi-continuous pulsed mode instead of fully-cw laser heating is an efficient way for surface finish improvement. For similar average powers, the use of a pulsed mode with large duty cycles is clearly shown to provide smoothening effects. The formation of larger and stable melt pools having less pronounced lateral curvatures, and the reduction of thermal gradients and Marangoni flow in the external side of the fusion zone were assumed to be the main reasons for surface finish improvement. Additional results indicate that combining the benefits from a pulsed regime and a uniform laser irradiation does not provide further reduction of surface roughness.

Introduction

The direct metal deposition (DMD) or laser assisted direct fabrication technique is an attractive manufacturing technique that has been investigated since more than a decade, and allows obtaining sound and near net shape metallic parts from a three dimensional CAD model as indicated by [START_REF] Kobryn | The laser additive manufacture of Ti-6Al-4V[END_REF], or more recently by [START_REF] Pinkerton | Laser direct metal deposition: theory and applications in manufacturing and maintenance[END_REF]. In the DMD process, a high power focused laser beam (usually Nd:YAG, fiber or diode laser) is used as a heat source to selectively generate a molten pool on a substrate, where a coaxial powder feeding increases the material volume and contributes to the formation of a solid layer. The substrate is then scanned on (x,y,z) directions below the (laser + powder nozzle) system and the layer-by-layer manufacturing of complex three dimensional metal parts is finally obtained.

Recently, other layer additive techniques like shape metal deposition, combining a wire feeding and a laser or a TIG welding torch have also been shown to manufacture sound and resistant Ti64 samples [START_REF] Baufeld | Wire based additive layer manufacturing: comparison of microstructure and mechanical properties of Ti-6Al-4V components fabricated by laser-beam deposition and shaped metal deposition[END_REF], with higher deposition rates than powder-based DMD, and a 100% efficiency. Compared with such techniques, the benefits of DMD are expected to come first from the multi-directionality of the powder feeding allowing the formation of more complex shapes, and second to the possibility to manufacture graded materials and to obtain better surface finish.

Many factors influence the DMD process efficiency, stability, or final metallurgical and mechanical properties, because of modifications of the global thermal cycle T = f(t) experienced by each point of the fabricated structure. A large number of parameters are involved in the DMD process. The most frequently investigated are: laser power P 0 (W), scan speed V 0 (m/min), laser diameter d 0 (mm), and average powder flow rate D 0 m (g/min). Other experimental parameters (defocusing distance, gas shielding nature and gas flow rate, time delay between passes, spatial distributions of laser and powder etc.) also can have significant impact on the process. All these parameters, which are highly interactive in nature, influence the thermal history of the part, and contribute to the melt pool shapes, and the resulting layer growth [START_REF] Peyre | Analytical and numerical modelling of the direct metal deposition laser process[END_REF][START_REF] Kumar | Effect of three-dimensional melt pool convection on process characteristics during laser cladding[END_REF]. The manufactured parts are then defined by layer height " h" and width "e", which can vary during the process, especially on complex shapes. However, limitations of the DMD process still remain such as:

(1) the need of a robust and reliable process control to maintain stationary local thermal conditions, and a stable and constant layer growth; (2) the occurrence of a deleterious surface finish usually microscopic roughness R a > 15 m as shown by [START_REF] Maisonneuve | Profil Project: direct manufacturing of aerospace components by laser cladding and laser sintering[END_REF] on Ti64 or [START_REF] Pinkerton | The effect of laser pulse width on multiple-layer 316L steel clad microstructure and surface finish[END_REF] on AISI 316L, that requires intensive post-machining steps. Such surface finish restraints the applications of DMD to the drafts of final parts, contrary to selective laser melting (SLM) that allows obtaining better surface finishes, mostly because of the use of smaller powder grains and the formation of thinner additive layers.

The optimization and understanding of surface finish has never really been addressed in itself in previous work except in [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF]. A common thought is that optimum surface finishes are mostly expected when using thin and stable deposited layers, when limiting melt pool disturbance, but without more precisions concerning process parameters, and melt pool shapes and dynamics. Moreover, as reminded by [START_REF] Pinkerton | Laser direct metal deposition: theory and applications in manufacturing and maintenance[END_REF], one can assume that melt pool stability depends on various internal forces (Marangoni flow, gravitational forces) or external forces (gas pressure, recoil forces near vaporization point, dynamic forces applied by the projected powder grains) applied on the molten metal.

Among the factors susceptible to smoothen surfaces, [START_REF] Zhu | The influence of laser and powder defocusing characteristics on the surface quality in laser direct metal deposition[END_REF] have shown that positioning powder focus below the melt pool, and laser focus above was an efficient method, mostly due to a melt pool enlargement. Optimized process control, and limitations of melt pool fluctuations were also shown by [START_REF] Alimardani | Surface finish in laser solid freeform fabrication of an AISI 303L stainless steel thin wall[END_REF] to be possible contributors to surface finish improvements.

Pulsed laser regimes have already been used to manufacture parts using the DMD technique. [START_REF] Pinkerton | The effect of laser pulse width on multiple-layer 316L steel clad microstructure and surface finish[END_REF] have pointed out a beneficial yet-limited effect of using pulsed laser irradiations (10-40 ms durations) on the surface roughness obtained on a 316L stainless steel. In another publication by [START_REF] Xue | Direct manufacturing of net-shape functional components/test pieces for aerospace, automotive and other applications[END_REF], near 2 m mean roughness R a were obtained by DMD with the use of a pulsed mode on a Ni-based alloy, even if no precise indications on process parameters are given.

In a recent publication [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF], the authors have shown that the generation of deep melt pools (of maximum height = H) combined with thin deposited layers h was a keyfactor for ensuring good surface finish on Ti64 alloy, by favoring a reduction of lateral menisci, and a nearly total suppression of agglomerated particles. The resulting centrifugal Marangoni flow (d /dT (N/m K) < 0) was assumed to act as a shaping contribution to the melt pool and to modify the resulting surface finish during layer additive processes. A simple analytical model was also proposed to calculate the meniscus mean amplitude W p versus melt pool dimensions (height H, width e) and additive layer thickness h. For similar melt pool geometry (assumed to be elliptic), it was shown that smaller menisci (i.e. better surface finish) could be achieved with high H/ h ratio, and large melt pool width values. Another important result from [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF], was the beneficial influence of a top-hat laser distribution which was shown to act as a shaping effect on melt pool stabilized shape.

However, in additive layer manufacturing processes, questions still remained concerning the dependence between temperature distributions in the fusion zone, melt pool shapes, and final surface finish. Additionally, the precise influence of thermo-physical materials' properties was not addressed as we only considered Ti64 for our investigations.

The present paper is a logical continuity of this latter one, considering more specifically the simultaneous influence of pulsed regimes and laser spatial distribution on the surface roughness of Ti64 titanium alloy.

Experimental conditions

Direct metal deposition conditions

DMD tests (Fig. 1) have been carried out in pulsed regime using a HL 10002 continuous wave (cw) disk Yb:YAG laser operating at 1.03 m, with 320-800 W peak laser powers, 2-10 ms pulse durations, and 40-100 Hz pulse frequency (Table 1). Two laser beam spatial distributions were used:

(1) A defocused near Gaussian laser spot ( (W m -2 ) = 5P 0 /( r 2 laser ) exp(-5x 2 /r 2 laser )) provided by a 0.4 mm optical fiber delivery, and respectively 200 mm collimation and focusing lens in the laser head. The laser beam was used with a +5 mm defocusing condition, resulting in a 1.2 mm diameter beam on the substrate (Fig. 2b).

(2) A top-hat laser spot ( (W m -2 ) = P 0 /( r 2 laser )) with 1.7 mm diameter obtained with a 600 m optical fiber, a 100 mm collimating lens and a 280 mm focusing lens.

Argon was used as a driving and shielding gas, in order to ensure powder transportation, and to limit oxidation phenomena. All the tests were carried out with a local Ar shielding (through the nozzle, and around), resulting in approximately 300 ppm O 2 in the interaction zone, as confirmed by local O 2 measurements. Even if the O 2 content in Ti alloys has a well-known detrimental effect on mechanical properties, the influence of oxidation on surface finish was recently shown to be negligible in the [20-500 ppm] range [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF].

For the powder distribution, a coaxial powder delivery nozzle system was used, with the powder flow rate directly proportional to the rotational speed of a delivery hopper. With this system, the Titanium powder (45-75 m, TLS Technik) is delivered coaxially with the laser beam, resulting in a R p ≈ 1.7 mm powder stream radius near the focus point (Fig. 2a), and with average mass feed rates D m in-between 1 g/min and 3 g/min. The mass feed profile was experimentally determined [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF], and was shown to be near-Gaussian (D (g/s/m 2 ) = 5D m /( • r 2 p ) • exp(-5x 2 /r 2 p )). The factor 5 included in the analytical formulation was shown to provide the best fitting with the local mass feed rate distribution measured experimentally.

The relative positioning of laser beam versus powder stream is presented in Fig. 1 (d i = interaction distance between powder stream and laser beam ≈ 1 mm). All the DMD tests have been carried out with 40 mm-length/6-15 mm-height walls, starting from 2 mm-thick titanium sheets (Fig. 1), and with scan speeds in the 0.1-0.6 m/min range. A 10 s time pause was used between each layer to limit heat accumulation on wall edges, and two opposite directions were considered for the deposition (forward and backward). Experimental conditions are summarized in Table 1.

Fast camera analysis

A simultaneous coaxial and lateral recording of melt pool sizes and dynamics was carried out, using two synchronized fast cameras with C-Mos sensors (Photron), at frequency rates up to 4000 Hz, using KG3 filters to cut the laser wavelength (1.053 m), and halogen lights to improve the contrast of recorded images. This allowed us to analyze melt pool dimensions (length L, height H, area S) and flow (fluid velocity v f ) in order to try to correlate them with surface finish data.

Temperature measurements in fusion zones

The determination of temperature fields in melt pools was carried out using a new spectrometer-based method developed by [START_REF] Muller | Temperature measurement of laser heated metals in highly oxidizing environment using 2D single-band and spectral pyrometry[END_REF]. This allowed us to determinate in two steps:

(1) the emissivity and local temperature at the surface of the melt pool on a 500 m diameter spot using a spectrometer, (2) the 2D temperature distribution by correlating gray levels from a C-Mos fast camera (obtained with luminance temperatures). The emissivity was calibrated using the melting point of pure elements such as Niobium and Tungsten. A ε = 0.37 ± 0.02 emissivity was found in the 0.5-0.7 m spectral window (interferential filter) used for melt pool analysis. This allowed us to quantify the thermal field by correlating gray levels of pixels to local temperatures. More data concerning the method are available in [START_REF] Muller | Temperature measurement of laser heated metals in highly oxidizing environment using 2D single-band and spectral pyrometry[END_REF]. The overall objective of thermal analysis was to assess temperature maximum values and gradients, with the objective of predicting thermo-convective effects.

Surface analysis

DMD surfaces exhibit a periodic structure composed of an accumulation of lateral menisci directly due to the equilibrium shapes of the melt pool. Manufactured walls were analyzed using: (1) low and medium magnification optical microscopy to visualize the global surface aspect and the grain microstructure (after Kroll's etching), (2) scanning electron microscopy (SEM-FEG Hitachi 4802 II) to provide a local description of DMD surfaces, (3) 2D and 3D profilometry (Veeco Dektak 150) to quantify surface topography. An example of scanned pictures of DMD walls is shown in Fig. 3.

Experimental characterization of the DMD melt pools

Melt pool dynamics and melt pool dimensions versus process parameters

Different regimes of melt pool behavior could be achieved, depending on the duty cycles dc (=laser pulse duration divided by the laser period T = 1/f), corresponding to the time ratio of laser "on" to total time. Examples of more or less pronounced variations of the melt pool size during the DMD process are shown in Figs. 4 and5. The "S" surface was measured considering the projected image of melt-pools view on CCD sensors, followed by a contour determination using ImageJ software. In Figs. 4 and5b, for a dc = 0.6, the lateral melt pool surface varies of 25% during two laser periods whereas in Fig. 5a, for a smaller dc value (0.3) a 50% fluctuation of the melt pool is shown (Fig. 5a). Using such procedures, we could evidence different kind of behavior for the melt pool: (1) below a given dc value (dc th1 ≈ 0.1-0.25 dependent on the peak power P peak ), a complete solidification is achieved between two pulses, (2) above a high dc value (dc th2 ≈ 0.85-0.95), the melt pool does not fluctuate and can be considered as geometrically stable, and (3) in-between those two dc thresholds, fluctuations of the melt pool gradually increase (Fig. 5a andb).

If we now consider melt pool dynamics (Marangoni convection flow), by analyzing particle velocities at the surface of the melt pool, a centrifugal flow is systematically observed, with average fluid velocities close to 0.4 m/s, consistent with results obtained in cw regime. For low dc value (below dc th1 ), the hydrodynamic flow seems to be reduced, and the fusion zone (FZ) experiences mostly a quasi-hydrostatic behavior.

Analysis of melt pool temperatures

The determination of temperature fields in titanium melt pool is important because it allows us quantifying temperature gradients, and resulting surface tensions gradients and fluid velocities, due to thermo-capillary convection. Another important aspect is the determination of maximum temperatures susceptible to favor vaporization of alloying elements such as aluminum (T v = 2500 • C). Among the main issue for thermal measurements in small melt pools: the determination of the emissivity ε of the melt titanium alloy which requires a pre-calibration, and the need of a good spatial resolution. A specific method was developed for this purpose using a pyrometer system coupled with a CCD fast camera analysis of the fusion zones.

The influence of duty cycles on temperature fields was more specifically investigated.

When comparing two duty cycles (dc = 0.53 and dc = 0.9), the following observations can be done:

(1) Increasing the dc value and the resulting mean power logically tends to increase the maximum temperature in the FZ. This temperature is nearly 3080 K for dc = 0.53 and 3320 K for dc = 0.9. (2) Thermal fluctuations T in the melt-pool during one single laser period decrease with dc (Figs. 5 and6). end of the laser pulse), and t = 10 ms (just before a new pulse), whereas a 0.9 dc value limits T value to (3320 -2850) = 470 K. (4) The maximum temperature achieved for dc = 0.9 corresponds to the boiling point of Ti64 (3320 K), and its time-dependence seems to exhibit a plateau (Fig. 6). This result indicates that a vapor pressure could have been applied on the FZ, resulting in possible shape modifications.

It was also clearly shown that the pulsed regime generates much higher maximum temperatures than the cw regime at similar mean powers. For instance, if we compare the maximum temperature evidenced in Fig. 6c (for dc = 0.53 and P peak = 600 W, P mean = 318 W) with the corresponding test in cw regime (320 W, cw), the cw regime generates a T max = 2600 K versus 3080 K in pulsed regime.

Evolution of surface finishes with process parameters for a

Gaussian beam distribution

General approach

With the use of SEM analysis (Fig. 7a), two kinds of surface roughness were distinguished, as already shown in [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF]:

(1) a microscopic contribution coming from (a) particle agglomerations mostly located in inter-layers areas, (b) solidification lines, (c) possible oxide layers. This micro contribution is favored by small melt pools, and rather low temperature distributions around the melt pool;

(2) the formation of periodic menisci, directly correlated to the melt pool stability (Fig. 7a).

For differentiating macroscopic waviness contribution (W a , W t ) from microscopic roughness (R a , R t ) parameters, a cut-off filter (threshold d th = 80 m) was used (Fig. 7b), just above the maximum diameter of particles (75 m) that usually provoke micro-roughening when agglomerating on the wall edges. The R a and W a values correspond to arithmetic average values (Eq. ( 1)) whereas R p and W p correspond to peak-to-average line values (Eq. ( 2)). W p data can be directly compared to the meniscus height. All measurements were carried out on a d total = 4 mm -length distance, using 3-4 tests for each DMD condition to provide statistically reliable data. [START_REF] Raphet | Etats de surface -Caractérisation[END_REF].

R a , W a = 1 d • d th 0 |Y (z)| • dz (1) R p , W p = 1 d . d th 0 |Y nf (z)| • dz (

Influence of a pulsed regime -case of a Gaussian laser irradiation

Most of the results presented on Ti64 in a recent publication [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF] were obtained for a 1.2 mm Gaussian laser irradiation. In turn, the specific influence of a pulsed regime, considering the same spatial distribution, was important. First of all, wall and melt pool dimensions (width "e", apparent height "H", height par layer " h") induced either in cw or in pulsed regime were compared at similar mean power (P mean = P peak × dc), revealing a very low dimensional change (Fig. 8).

In a second step, the variation of surface parameters versus duty cycles (Figs. 9-11) was analyzed and provided us the following information:

(1) In pulsed regime, the increase in duty cycles limits the amount of agglomerated particles and globally improves surface finish (Fig. 9 for a Gaussian laser spot). This occurs whatever the laser frequency (60 Hz or 100 Hz, Fig. 10). (2) The continuous decrease of both waviness and roughness parameters with dc value reveals a beneficial influence of the mean laser power (P mean = P peak × dc) with increasing duty cycles. High P mean induce thicker melt pools that promote high H/ h and e/ h values [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF] which tend: (1) to reduce particle agglomerations (the local mass feed rate that modifies particle concentration decreases with wider meltpool), ( 2) to reduce meniscus height (the melt-pool lateral curvature radius decreases with deeper and wider melt-pools).

(3) At a given frequency f, the best surface finish is obtained for a quasi-continuous pulsed regime (dc ≈ 0.9). A significant benefit is obtained compared with the cw regime: a factor 2 increase of R, W parameters is obtained between dc = 0.9 and dc = 1 (cw) (Fig. 11). ( 4) At similar dc values, if we compare surfaces manufactured at 60 Hz -P peak = 500 W and 100 Hz -P peak = 600 W, the best surface finish is clearly obtained for the higher peak power (Fig. 10).

Additional tests, carried out at 60 Hz -600 W, have confirmed that the laser frequency in itself was not a key-point. To emphasize the specific influence of a pulsed versus cw regime, surface parameters induced by the two regimes were compared, on a large range of mean powers (Fig. 12). Even at similar mean power (P mean = P peak × dc), corresponding to a similar global heat input, the quasi-continuous pulsed regime induced better Fig. 12. Influence of interaction regime (cw or pulsed) and mean power on surface finish (pulsed regime considered for P peak = 600 W and various dc), Gaussian laser irradiation 1.2 mm. surface finish that the continuous regime. Consequently, a real benefit from the pulsed regime at high duty cycle could be assumed, especially for high mean powers.

Influence of the beam distribution on surface finish obtained in pulsed regime

Previous results presented in [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF] have shown the benefit from using a uniform 1.7 mm "top-hat" laser distribution instead of a 1.2 mm Gaussian irradiation in cw regime. Following this, comparative tests were carried out in pulsed regime to address a possible additional effect of spatial distribution (1.7 mm top-hat instead of 1.2 mm Gaussian) on surface finish. Before addressing surface finish, one has to remind two aspects:

(1) the temperature distribution in the melt pools is more homogeneous for the top-hat distribution, than for the Gaussian irradiation, resulting in reduced thermal gradients and expected Marangoni flow; (2) despite the reduction of thermal gradients, expected to inhibit centrifugal flow that enlarge melt pools, melt pool dimensions are nearly kept constant by the 1.7 mm top-hat irradiation, due to a better spatial distribution of heat input (higher power density values near the edges of the laser irradiation zone).

The analysis of waviness parameter W p (i.e. the height of the lateral menisci) was shown to indicate a moderate surface finish improvement with Gaussian irradiation at similar mean powers (Fig. 13). Moreover, with a 1.7 mm top-hat irradiation, the best Fig. 13. Influence of the laser beam spatial distribution on the surface finish induced in pulsed regime for one peak power P peak = 600 W, two frequencies (60 Hz and 100 Hz), and various duty cycles dc. For similar duty cycles the Gaussian irradiation mostly generates better surface finish.

Fig. 14.

Influence of the duty cycle on melt pool shapes (cross-sections) for a Gaussian irradiation (P peak = 600 W, f = 100 Hz), (a) dc = 0.45, (b) dc = 0.75, (c) dc = 0.9. The concavity of the lower melt pool limit (Â2 angle), and the melt pool apparent height "H" and width "e" increase with dc, but the additive layer h remains constant. The best surface finish is obtained for the higher dc values and the higher  values.

surface finish was obtained for a dc = 1, i.e. for a cw irradiation, whereas for the Gaussian distribution, minimum W values were evidenced for dc ≈ 0.9. Consequently, it seems that the real benefit from a quasi-continuous regime is mostly evidenced for a Gaussian laser irradiation, where thermal gradients in the fusion zone are higher than for the top-hat distribution. It was also assumed that, for mean powers above 600 W (top-hat 1.7 mm), vapor pressure effects acting as a normal force on the melt-pool, could have increased meniscus height by expanding laterally the melt-pool, thus provoking an increase of W p parameters.

Discussion: influence of a pulsed regime on melt pool shapes and surface finish obtained with the DMD technique

For a given spatial distribution of the laser beam (Gaussian), experimental results clearly evidence optimum surface finish with the use of a pulsed DMD regime, especially when using high duty cycles corresponding to quasi-continuous laser irradiations. The benefit of using a near continuous pulsed regime is assumed to be attributed to a combination of physical factors:

(1) The best surface finishes are obtained for high mean powers P mean , corresponding to high H/ h factors, as previously demonstrated in [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF]. The improvement is shown not only on meniscus height (W) but also on microroughness data (R). Globally, thicker and wider melt pools limit solid particle sticking by limiting the local mass feed rate coincident to the lateral edges of the walls, favor particles re-melting, and in turn tend to limit the amplitude of lateral menisci, and the density of agglomerated particles.

(2) Provoking a discontinuity in laser irradiation (≈using a quasi-continuous pulsed regime) allows reducing temperature gradients, interrupting Marangoni lateral flows, and enhancing lateral menisci curvature. The resulting effect is a decrease of the mean velocity in the melt pool, which could not be fully demonstrated experimentally by CCD cameras, but which was confirmed by thermo-hydraulic numerical simulations [START_REF] Morville | 2D finite element modeling of heat transfer and fluid flow during multi layered DMD laser process[END_REF], [START_REF] Morville | Modélisation thermohydraulique du procédé de fabrication Directe par Projection Laser en vue d'améliorer l'état de surface final[END_REF]. This decrease is assumed to be the main reason why W factors are improved by pulsed regimes when carried out with similar mean powers than with a cw regime.

On metallurgical cross-sections of the samples (using Kroll's reagent etching), the use of higher duty cycles logically promotes melt pool enlargements, and more pronounced concavities (increased  angles) due to more severe Marangoni flow (∂ /∂T < 0 on Ti64), together with deeper melt pool (Fig. 14). Large dc values also improve surface finish as shown in Fig. 11. This result is consistent with what has been previously demonstrated in cw regime [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF].

However, when comparing melt pool shapes induced by nearly similar mean powers in cw or pulsed regime (for instance: a 500 W cw irradiation and a (P peak = 600 W -dc = 0.9) pulsed configuration corresponding to a P mean = 540 W mean power), a significant difference is shown on the fusion zone (FZ) height, and on the concavity angles  1 and  2 (Fig. 15). The interesting point is that not only the dimensions, but also the FZ shape is modified by the use of the pulsed regime. This seems to indicate a modification of the local fluid flow, which could not be evidenced with fast cameras, but which has a real effect on melt pool shapes and resulting surface finish.

The resulting conclusion is that not only the apparent (and maximum) melt pool H height and the additive layer h, but also the local  angles at the lateral side of melt pools are the dominant factors that influence the meniscus height, and in turn the W roughness parameter: the higher the  values, the lower the W parameters (Fig. 16). In a previous publication [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF], elliptic melt pool shapes were considered, including H and h in an analytical prediction of W values, but without really addressing the local lateral shape of the FZ in terms of  1 and  2 angles. One of the most unexpected result is that  1 and  2 values are modified by DMD parameters, but do not systematically follow the larger thermal gradients and the higher Marangoni flow. For instance, the comparison between the pulsed mode and the cw mode shows that higher  angles and better surface finish are obtained with the pulsed mode, which is shown to reduce ∂T/∂y gradients in the FZ. A first order estimation of the W value can be obtained using  1 and  2 angles (Eq. ( 3)). However, for the calculation, very local  1 and  2 angles were considered, near the edge of the FZ.

W = h tan(Â 1 ) -tan(Â 2 ) (3)
For nearly identical FZ dimensions, the influence of a pulsed regime on roughness values (=particles agglomerated) is not so obvious. Classically, wider walls tend to reduce local mass feed rates D * m (g/s/m 2 ) in front of the lateral wall edges (Fig. 16), and contribute to a roughness decrease. However, wall dimensions obtained in cw or pulsed regime are nearly identical for similar mean powers (Fig. 8). Consequently, it was assumed that it was the lateral FZ shape modification (increase of  1 and  2 angles ≈ more vertical curvature of the FZ) which limited particle sticking, instead of the reduction of local mass feed rate D * m . If we now consider a uniform beam distribution (top-hat 1.7 mm) already shown to be beneficial for surface improvements [START_REF] Gharbi | Influence of various process conditions on surface finishes induced by the direct metal deposition laser technique on a Ti64 alloy[END_REF], combined with a pulsed regime, no further improvement was shown compared to the Gaussian 1.2 mm irradiation (Fig. 13). Consequently, starting from a cw Gaussian DMD process, two distinct options are possible for improving surface finish: either (1) the use of a quasi-continuous pulsed regime or (2) the use of a cw top-hat irradiation. The combination of these two factors (a pulsed top-hat DMD) did not provide additional improvement of surface finish.

With such optimized conditions, attractive surface finish could be obtained (Table 2), with R a and R p as measured (non-filtered) values of respectively 3-5 m and 7-18 m, and mass efficiency ratios (of approximately 50-75% better with a top-hat 1.7 mm irradiation due to larger melt pools). Corresponding experimental parameters allow combining optimum surface finish with acceptable efficiency rates. However, the manufactured volume per hour remains low (10 cm 3 /h of Ti64), due to low mass feed rates (1 g/min) which are a dominant factor for achieving smooth surfaces. Such roughness data indicate a decrease of the main (non-filtered) roughness parameters by a factor of more than 5, compared with previously published data on Ti64 alloy. With this experimental procedure, post-machining steps are expected to be strongly reduced, and the DMD technique already allows providing similar or even-better surface finish than the SLM technique (Fig. 17).

In a final step, SEM-EDS investigations were carried out in order to check possible compositional changes on DMD samples, especially for those manufactured with high peak powers, where very high local temperatures were measured (Fig. 6). Those temperatures, locally superior to 3200 K, are expected to promote intensive vaporization of Al (boiling temperature T b = 2782 K) or Ti64 (T b = 3330 K). However, EDS analysis did not point out changes Al and V compositions, even for high P mean /V ratio. Moreover, no real hardness difference was shown between samples

Table 2

Best surface finish (non-filtered values) obtained with quasi-continuous pulsed laser regimes ( = experimentally-determined powder efficiency ratio) and a constant Dm = 1 g/min (Rt = total roughness amplitude = peak valuevalley value). Measurements have been carried out along the -z direction.

DMD conditions

Ra (m) Rp (m) Rt (m) Gaussian 1.2 mm -P peak = 600 W 100 Hz ( = 9 ms, dc = 0.9) 2.5 6.6 13.4 0.51 Gaussian 1.2 mm -P peak = 500 W 100 Hz ( = 9 ms, dc = 0.9) 5. manufactured in pulsed regime or in cw regime: all the samples were shown to exhibit HV 0.1 = 420 ± 20 whatever the process conditions. Consequently, and even if further experiments (electron microprobe analysis etc.) should be done to provide more precise data than EDS on the chemical content of DMD walls, it could be assumed that improved surface finish could be obtained together with satisfactory metallurgical behavior, even when carried out with rather high peak powers. However, a more detailed metallurgical investigation of pulsed DMD structures will have to be carried out to investigate the specificity of a pulsed laser regime versus grain morphology, microstructures and resulting mechanical properties.

More widely, it should also be interesting to address in the near-future the influence of the thermo-physical properties of the projected material on the surface finish, which is still difficult to estimate as a whole.

Conclusions

The following conclusions can be drawn from this study:

-This work demonstrated experimentally that the use of quasicontinuous instead of fully continuous laser irradiations can represent a new way to improve the surface finish of Ti64 parts produced by DMD. This was attributed to reduced thermal gradients and Marangoni lateral flow in the melt pool. -With such laser irradiations, combined with low mass feed rates (D m = 1 g/min) resulting in thin additive layers ( h < 0.2 mm), average roughness values R a of near 3 m and peak-to-average values R p near 10 m could be obtained, which result appears to be a huge improvement, compared with previous data reported in the literature. -The combination of a pulsed regime with a uniform "top-hat" irradiation was not shown to add further improvement to surface finish.

These results tend to indicate that using a pulsed laser operating mode, the DMD process can be envisaged for manufacturing near net-shape parts, with limited post-machining steps.

Fig. 1 .

 1 Fig. 1. Experimental DMD set-up and associated diagnostics (laser-powder-melt pool interaction zone), h = layer height, e = wall thickness.

Fig. 2 .

 2 Fig. 2. Experimentally determined (a) powder distribution (for Dm = 2 g/min) in relative units (using CCD camera analysis of a diode laser-illuminated powder stream) and (b) laser beam distribution (for P peak = 450 W, and a d laser = 1.2 mm defocused condition). Both distributions could be approximated by near-Gaussian formulations (f(x) = A • exp(-5x 2 /r 2 )).

Fig. 3 .

 3 Fig. 3. Scanned images of DMD walls (40 mm length) obtained at P = 600 W, V = 0.4 m/min, Dm = 1 g/min, f = 100 Hz in pulsed regime (dc = duty cycle).

Fig. 4 .

 4 Fig. 4. Fast camera analysis of melt pool fluctuations during a DMD in pulsed regime (Dm = 1 g/min, P peak = 600 W, V = 0.4 m/min, f = 100 Hz, laser = 6 ms, dc = 0.6) -Series of pictures captured at 1330 frames/s (t0 = beginning of the laser pulse), obtained with a 1.2 mm Gaussian laser spot. The fusion zone (FZ) surface varies by 40% during the laser period.

Fig. 5 .

 5 Fig. 5. Fast camera analysis of melt pool fluctuations (projected surface S) with time for 2 different duty cycles (P peak = 600 W -100 Hz (10 ms period)): (a) = 3 ms → dc = 0.3 and (b) = 6 ms, dc = 0.6. A 50% fluctuation of the melt pool is obtained for dc = 0.3, and a 25% fluctuation of S is obtained for dc = 0.6.

Fig. 6 .

 6 Fig. 6. Temperature distribution at the surface of the melt pool (100 Hz, P peak = 600 W, V = 0.4 m/min, Gaussian irradiation) using fast camera + pyrometer system -(a) example of side view thermal image of the melt pool captured during a time period, (b) dc = 0.53, (c) dc = 0.9, (d) variation of temperature with time.

Fig. 7 .

 7 Fig. 7. (a) SEM picture of a DMD surface obtained in cw regime (550 m-height menisci, agglomerated particles accumulated in the inter-layers zones and solidification lines) 400 W -2 g/min -0.4 m/min, (b) 2D profile -distinction between micro-roughness R and macro-waviness W using a cut-off filter of 80 m from a starting value Y(z).

  2) With d = d th for R values and d = d total for W values, Y(z) = non filtered line profile with Y(z) = 0 as average line value, and Y nf (z) = non filtered line profile with Y nf (z) = Y(z) + Y max,peak

Fig. 8 .

 8 Fig. 8. Comparison of melt pool dimensions for walls manufactured either in continuous or pulsed regimes at similar mean power (Pmean = P peak × dc): no clear difference is shown between the pulsed (P peak = 600 W, 60 Hz) and cw regime (e = wall and meltpool width, H = apparent melt-pool height = maximum melt-pool height measurable by lateral camera analysis, h = additive layer height).

Fig. 9 .

 9 Fig. 9. SEM pictures of DMD surfaces obtained in pulsed regime (P peak = 600 W -f = 60 Hz -Dm = 1 g/min) -Influence of the duty cycle (a) dc = 0.36, (b) dc = 0.48, (c) dc = 0.6, (d) dc = 0.84.

Fig. 10 .

 10 Fig. 10. 3D profiles of surface finish induced in pulsed regime (P peak = 500 W -f = 60 Hz, Dm = 1 g/min) (a) dc = 0.3 (non-filtered average roughness = R a,nf = 29 m), (b) dc = 0.5 (R a,nf = 17 m), (c) dc = 0.7 (R a,nf = 9 m), (b) dc = 0.9 (R a,nf = 5 m).

Fig. 11 .

 11 Fig. 11. Influence of the duty cycle on surface finishes (Wp, Rp) induced for a Gaussian laser irradiation for two different pulsed regimes: (a) P peak = 600 W -f = 100 Hz, (b) P peak = 500 W -f = 60 Hz.

  (5) Non-filtered roughness values (cut-off = 0), indicate optimum roughness values near R a = 2 m and R p = 5 m. Such values are much lower than those from previously published data on Ti64[START_REF] Maisonneuve | Profil Project: direct manufacturing of aerospace components by laser cladding and laser sintering[END_REF].

Fig. 15 .

 15 Fig. 15. Comparison of cross sections of walls obtained at nearly the same mean power in cw regime or in pulsed regime (a) cw with Pmean = 500 W, (b) pulsed regime with P peak = 600 W and dc = 0.9 (corresponding to Pmean = 540 W). Almost similar wall dimensions, but the pulsed regime induces more concave FZ (higher Â1 and Â2 angles).

Fig. 16 .

 16 Fig.16. Influence of Â1 and Â2 lateral melt pool angles on meniscus amplitude for similar h values. An increase of Â1 and Â2 angles is beneficial for surface finish.

Fig. 17 .

 17 Fig. 17. Improvement of surface finish using an increase of duty cycles on DMD-manufactured walls (P peak = 500 W, f = 100 Hz, V = 0.4 m/min), (a) dc = 0.45, (b) dc = 0.6, (c) dc = 0.75, (d) dc = 0.9.

Table 1

 1 Experimental DMD conditions (with d laser = 2 r laser ). With such conditions, deposition rates are between 0.05 kg/h and 0.1 kg/h.

	Regime	Spatial distribution + d laser (mm)	Laser peak power P peak (W)	Scan speed V (m/min)	Pulse durations (ms)	Frequency f (Hz)	Gas shielding
	Pulsed	Gaussian 1.2 mm	500, 600, 800	0.4	2-10 ms	40, 60, 100	Argon
	Pulsed	Top-hat 1.7 mm	500, 600, 800	0.4	2-10 ms	40, 60, 100	Argon
	Continuous	Gaussian 1.2 mm	320, 400, 500	0.4	-	-	Argon

Acknowledgments

The authors acknowledge the financial support of French National Agency for Research (ANR) under the frame of the ASPECT Project (ANR-09-BLAN-0014). The SEM investigations have been performed on the FEG-HITACHI 4800 of PIMM which has been funded with the help of the SESAME program of Île-de-France Region, Arts et Métiers ParisTech and CNRS.