M Pocquet

Z Wu

Y Bessiere

I Queinnec

X Lefebvre

M Mauret

M Spérandio

A control system for nitrification / denitrification over nitrite in SBR with simultaneous mitigation of N2O emissions

Keywords: SBR, Control, Nitrification, Denitrification, Nitrous oxide

A control system for limiting N2O emissions in a SBR treating an ammonium rich effluent (~500 mgN.L -1 ) with nitrification / denitrification over nitrite is presented. This system is based on the use of derivatives of DO and ORP signals for on line control of aerobic and anoxic periods. During the 120 days of operation, the effect of different operating conditions on N2O emissions during nitrification has been analyzed. Experimental results have highlighted the effect on N2O emissions of: HNO2 concentration, DO concentration and the presence of COD during nitrification. The control system allows reducing the N2O emission factor (N-N2O per N removed) to less than 1%.

Introduction

Nitrification combined with denitrification (N/DN) over nitrite is an interesting process for biological treatment of ammonium-rich wastewaters. In this system the needs of oxygen and organic matter are lower than those of conventional biological nitrification and denitrification. This alternative allows reducing energy consumption and chemical use for denitrification (example: methanol). However, in system treating N-rich wastewater, ammonium-oxidizing bacteria (AOB) can produce a significant amount of N2O which is a major greenhouse gas. It was demonstrated that a major part of this gas was produced by the mechanism of autotrophic denitrification, i.e. the reduction of nitrite by AOB [START_REF] Kampschreur | Nitrous oxide emission during wastewater treatment[END_REF]. The N2O emissions of systems with N/DN over nitrite can be much higher than those of conventional system with full nitrification (Rodriguez-Caballero et al., 2013). In Sequencing Batch Reactor (SBR) operated for N/DN over nitrite, N2O emission factors can reach more than 5% (mass of N-N2O produced per N removed) which clearly counter balances the benefits of partial nitrification, i.e. reduction of oxygen demand and COD needs (Rodriguez-Caballero et al., 2013). Moreover, alternating aerobic and anoxic periods has also a detrimental effect on N2O emissions (Rodriguez-Caballero and Pijuan, 2013). Therefore the objective of this work is to develop a control system for N/DN over nitrite with simultaneous mitigation of N2O production. Compared to previous systems based on on-line nitrite measurement [START_REF] Lemaire | Advanced control system to treat ammonia-rich effluent and reduce N2O emission by using an SBR operating via nitrite pathway[END_REF], this controller is based on simple sensors, cheap and easy to maintain: dissolved oxygen (DO) and Oxidation-Reduction Potential (ORP).

Material and Methods

Experimental set-up and operating conditions

Experiments were performed in a lab-scale reactor (1.43-1.63 L of liquid) equipped with an intermittent aeration system (fine bubble injection), a mechanical stirrer (Rushton type with constant stirring rate of 465 rpm) and a water jacket for temperature control (maintained at 28 ± 0.5 °C). pH (H8481HD from Schott), dissolved oxygen (DO) (Visiferm DO Arc 120 from Hamilton) and oxidation-reduction potential (ORP) (PL89225Pt from Schott) were monitored during all the study. Air flow rate was imposed with a mass flowmeter. The gas was collected at the top of the reactor with a flow of 0.2 L.min -1 for analysis of N2O (X-STREAM X2GP from Emerson; range: 0-100 ppm). All the data were monitored every 20 seconds. A control panel connected to a computer allowed managing stirring, aeration and all inputs and outputs of the reactor. The system was inoculated with a sludge sample from the activated sludge process of a WWTP with stable nitrification (Graulhet, France, 50 000 PE; sludge retention time of 20 d) receiving domestic and industrial wastewater (70% from tanneries). The initial concentration was closed to 4 gVSS.L -1 . The lab scale reactor was operated in SBR mode with automatic feeding and wastage. A sludge retention time (SRT) of 15 d -1 was maintained. The system was fed with a high strength synthetic wastewater (462 mgN-NH4 + .L -1 ) and a complementary solution was used as a source of organic carbon (18.8 gCOD.L -1 ) for heterotrophic denitrification. During all the study, chemical species were quantified by ionic chromatography (IC25, 2003, DIONEX, USA). Ammonium, VSS, COD were determined using standard methods. The reactor was operated over a 120 days period in sequencing batch mode with a cycle composed of alternating aerobic period (nitrification) and anoxic period (denitrification), followed by settling (20 minutes) and withdraw (8 minutes). Aerobic period was initiated by the feeding of ammonium-rich effluent (20 mL.min -1 , 10 min) and anoxic period started with addition of complementary solution respectively (6.25 mL.min -1 , 2.5 to 7.5 minutes). During the first 35 days of operation (case A), nitrification and denitrification durations were fixed to 100 minutes and 180 minutes respectively. Then automatic control of aerobic and anoxic durations has been activated based on DO and ORP signals with the method described below (cases B, C, D).

Calculations

The N2O gas phase concentration measured is converted into emission rate (ER) with the following equation, considering that N2O behave like an ideal gas. The total amount of N2O emitted during an aerated period is calculated by integration of the N2O-ER. Emissions factor (noted EF) is calculated as the ratio of N2O emission to the amount of ammonium removed.

N2O-ER (gN-N2O.L -1 .h -1 ) = N2O(ppmv) .10 -6 .Qgas(L.h -1 ).Vm -1 (25.0376 L.mol -1 at 28.0 °C and 1 atm).2.MN(14.00674 g.mol -1 ).V -1 (L liquid medium)

The concentration of free nitrous acid (HNO2, noted FNA) is calculated according to [START_REF] Anthonisen | Inhibition of Nitrification by Ammonia and Nitrous Acid[END_REF].

Results and discussion

Example of a typical SBR cycle

A typical cycle of the SBR obtained after 45 days of operation (phase B) is presented in Figure 1 (a). Ammonium is fed during the first 10 minutes (period 1) of the aerobic period with a volumetric exchange ratio of 12.3 %. In that example, it corresponds to a total amount of ammonium to nitrify of 56.7 mgN-NH4 + .L -1 which is partly oxidized into nitrite during this feeding period as nitrification takes place as soon as period 1 begins. Aeration is maintained (period 2) until the depletion of ammonium. All the nitrified ammonium was converted into nitrite as nitrite oxidation was stopped after 40 days due to the operational conditions. Actually, Nitrite Oxidising Bacteria (NOB) were outcompeted due to simultaneous effect of ammonium inhibition, substrate privation (aeration is stopped after ammonia depletion) and high temperature (growth rate of ammonium oxidizing bacteria becoming higher than those of NOB). Aeration is interrupted from periods 3 to 6 maintaining anoxic condition. Period 3 corresponds to the introduction of 18.75 ml of secondary solution (organic carbon source) into the reactor. Then, nitrite was removed by denitrification (periods 3 and 4). After the exhaustion of nitrite, a settling period (period 5) was followed by a last decanting period (period 6). Emission of N2O is observed throughout the aerobic period. The N2O profile in the gas phase shows a maximum (28.9 ppm). After that, the rate of N2O emission decreases according to the decrease of nitritation rate. In this example, 1.84% (emission factor) of the amount of nitrogen oxidized during the aerobic period has been transformed into N2O. During all the study, N2O was mainly emitted during aerobic periods. Emissions during the anoxic phase were quantified punctually by using nitrogen gas injection and were negligible compared to aerobic emissions, which can be explained by a sufficient COD/N ratio during denitrification period. It was also observed that the anoxic N2O production was limited if the aeration was stopped after the end of N2O peak emissions, i.e. a few minutes after the ammonium depletion.

Automatic control system

Evolution of DO, ORP, the first derivative of DO and the second derivative of ORP are presented in Figure 1 (b). DO increases rapidly as soon as the cycle begins, then stabilizes during nitritation period and starts to increase again when ammonia is depleted. The point α is determined by a threshold limit on DO derivative, corresponding to the end of nitritation. In the meanwhile ORP decreases during anoxic period and this rate accelerates as soon as the nitrite is depleted (commonly attributed to the start of sulfate reduction activity). The point ω corresponding to the end of denitrification is detected with a threshold limit on the second derivative of ORP. The automatic control system is based on these signals. When the point α is reached, this signifies that ammonium is depleted and the system automatically moves to the next period of the cycle. Similarly, the end of denitrification is detected when the point ω is reached and the system moves automatically to the settling period. Threshold values used for the control of the cycle presented in Figure 1 (b) are 14 mgO2.L -1 .h -1 and -12000 mV.h -2 for α and ω respectively. The first and second derivatives are calculated with linear regression associated to the 15 last data which correspond to the last 5 minutes. The use of derivatives signals allows the system to be relatively insensitive to possible drift of DO and ORP absolute values due to calibration default.

Effect of the control strategy on performances and N2O emissions

The impact of automatic control, cycle configuration and operating parameters has been analyzed on 4 successive case studies presented in Table 1. Period A corresponds to the reference condition with constant duration of aerobic and anoxic periods. The control system has been activated during case studies B, C and D. In addition, cycle configuration was slightly adapted during phases C and D (configuration 2): an aerated period (15 minutes) before the feeding of ammonium-rich wastewater has been introduced and the aerated period has been extended by 20 to 40 additional minutes after the exhaustion of ammonium (α). A step-feed mode was tested during phase C, and different air flow rates and COD/N ratios were also compared. For each case study, a same operating condition was maintained during at least 10 cycles. removal performance of the process (increasing the median of nitrogen removal rate from 0.29 gN-NH4 + .L -1 .d -1 for case A to 0.85 gN-NH4 + .L -1 .d -1 and 0.60 gN-NH4 + .L -1 .d -1 for subcases B1 and B2 respectively) but dramatically impacts the N2O emission factor, increasing the median of this factor between cases A, B1 and B2 (2.74 % for case A, 8.30 % and 4.28 % for subcases B1 and B2 respectively). A high variability of the N2O emission factor is observed both for the reference period (case A) with time based system (configuration 1) and with the automatic control without modification of the cycle configuration (B1, B2). The transition between configuration 1 and configuration 2 (additional short initial aerobic phase, introduction of a time delay after the end of nitrification) allows reducing N2O emissions while controlling nitrogen removal performances. A mean N2O-EF of 1.0 % is obtained for a mean nitrogen removal rate of 0.60 gN-NH4 + .L -1 .d -1 for the subcase D1. These values are managed by the amount of ammonium removed for periods automatically controlled by the system coupled to the optimized configuration. For subcases D2 and D3, this pair of mean values decreases to [0.27 % -0.30 gN-NH4 + .L -1 .d -1 ] and [0.32 % -0.31 gN-NH4 + .L -1 .d -1 ] respectively. The feeding of ammonium in 5 steps (case C) allows decreasing the mean N2O-EF to 0.24 % with a mean nitrogen removal rate of 0.21 gN-NH4 + .L -1 .d -1 .

Relation between FNA concentration and N2O emission factor

For each cycle the maximal nitrite and free nitrous acid concentrations (reached at the end of aerobic phase) were measured or calculated. On one hand, a large variation of N2O emissions was observed for the same level of nitrite concentration and no clear relationship can be demonstrated between these parameters (not shown here). On the other hand, the relationship between the maximum free nitrous acid concentration and N2O emission factor appears clearly in Figure 3 a. It appears that the N2O emission factor is exponentially correlated to the free nitrous acid concentration with a threshold value around 0.70 µgN-HNO2.L -1 beyond which the N2O emission factor increases significantly with a great dispersion. This observation is close to those presented in [START_REF] Lemaire | Advanced control system to treat ammonia-rich effluent and reduce N2O emission by using an SBR operating via nitrite pathway[END_REF]. Indeed, these authors have found that an FNA concentration lower than 0.80 µgN-HNO2.L -1 leads to an N2O-EF lower than 1.0 %. The range of FNA concentration corresponding to each case is presented in Figure 3 b with a statistic distribution illustrated by box plots. The largest dispersion of N2O emission factors (case B) has been obtained for the smallest dispersion of FNA concentrations. This can be explained by the fact that another parameter (see below DO effect) influences the emissions during this period. On the contrary, in cases C and D, low values of N2O-EF and relatively small dispersion of N2O-EF corresponds to a large dispersion of FNA concentrations whose maximum value is located below the threshold value.

Relation between DO concentration and N2O emission factor

The N2O emission factor versus DO is presented in Figure 3 c (average value of DO throughout the aerobic period). The statistic distribution of oxygen concentrations in cases A, B, C and D is presented Figure 3 d. For cases A and B, emissions of N2O increases when DO decreases. Highest emission factors (3 to 12% of nitrogen removed) were observed at low DO in case B, which is simultaneously the case with highest concentration of FNA. In cases C and D, the effect of DO does not appear clearly; N2O emission factors remain lower than 1.22 % throughout these periods despite an important variability of the dissolved oxygen concentration from 1.42 mgO2.L -1 to 5.10 mgO2.L -1 . In different previous studies higher emission of N2O was detected at low DO concentrations [START_REF] Kampschreur | Nitrous oxide emission during wastewater treatment[END_REF][START_REF] Pijuan | Effect of process parameters and operational mode on nitrous oxide emissions from a nitritation reactor treating reject wastewater[END_REF]. But in this study, this effect is more pronounced for cycles operated with the cycle's configuration 1 (cases A and B) than those operated with the configuration 2 (case C and D).

Correlated effects of operational parameters

A correlation (almost linear) appears between DO and pH for cases A and B. Variation of DO and pH depends mainly on two similar factors: (1) the variation of biodegradable COD present in aerobic phase which controls the CO2 production by heterotrophic activity and influences the DO for a given air flow rate and (2) the variation of air flow rate which controls the stripping rate of CO2 and the DO as well. In this study the presence of biodegradable COD during nitrification in aerobic phase is due to residual organic matter not used during the previous anoxic phase for cycles operated with the configuration 1. This excess leads both to DO and pH decrease, the later leading to an increase of FNA as this concentration is directly correlated to the pH value. For cases C and D, the correlation between DO and pH does not appear. Indeed, the introduction of the aerobic period before ammonium feeding allows removing the residual organic matter and increasing the pH with the stripping of inorganic carbon. The extension of the aerated period after the depletion of ammonium also allows increasing the pH as well. Thus, the optimization of the configuration leads to the decorrelation between DO and pH during nitrification and for a given DO concentration, the pH will be higher with the configuration 2 leading to a lower free nitrous acid concentration. This can partly explain the relation between DO and N2O-EF in Figure 3 c. Below a pH value of 8.2, the FNA concentration crosses the threshold value of 0.70 µgN-HNO2.L -1 for cases A and B corresponding to a DO lower than 2.5 mgO2.L -1 (Figure 3 e). For cases C and D, the pH is higher for this range of DO and FNA concentrations remain lower than 0.70 µgN-HNO2.L - 1 .

Control strategies for the mitigation of N2O emissions in SBR

Results clearly demonstrated the benefit of the control system based on configuration 2 (periods C, D) for minimising N2O emission factor. The N2O emission factor as a function of the nitrogen removal rate is presented in Figure 3 f. Highest emissions of N2O were clearly observed for case B, with a low DO, high FNA concentration and high nitrogen removal rates. For a given nitrogen removal rate and a given DO concentration the N2O-EF is lower with the optimized configuration (C and D). Thus, this optimized control system allows exploring interesting ranges of DO and nitrogen removal rates in term of process performances while, at the same time, limits the accumulation of FNA concentration, thus reducing significantly emissions of N2O. Finally several recommendations can be proposed for N2O emission mitigation in SBR with nitrification / denitrification over nitrite:

(1) Limit the presence of readily biodegradable COD during nitrification which impacts the competition between nitrifiers and heterotrophs on oxygen, decreases the DO and possibly the pH during nitrification. This was done in this study by introducing an aerobic phase before Nrich effluent feeding, this could be done also by an optimal pre-anoxic (DN) phase in SBR systems receiving wastewater containing both COD and ammonium.

(2) Limit the accumulation of FNA below 0.70 µgN-HNO2.L -1 . This can be done by limiting the amount of ammonium to be removed during nitrification (step-feed, reduction of the volumetric exchange ratio) or by increasing the pH. It was done in this study by means of a 5step feed strategy (case C) coupled to a slight increase of the pH due to modification of cycle configuration.

(3) Extend the aerobic period after the exhaustion of ammonium. The system avoids switching in anoxic conditions as the N2O emitted by AOB during nitrification is not finished (extension of the aerobic period after the detection of the DO bending point).

Conclusions

An automatic control system for nitrification / denitrification over nitrite is presented in this study. The DO and ORP bending points allow detecting the end of nitrification and denitrification respectively. By adapting the length of aerobic and anoxic periods, the system allows to outcompete NOB and maintain full conversion of ammonia into nitrite.

Several effects on N2O emissions were highlighted during nitrification. (1) The N2O-EF is exponentially correlated to the free nitrous acid concentration with a threshold value of 0.70 µgN-HNO2.L -1 beyond which the N2O-EF increases significantly.

(2) The effect of DO on N2O EF is exacerbated for cycles with a high FNA accumulation.

(3) The presence of easily biodegradable COD during nitrification increased the N2O-EF.

The optimisation of control system allows reducing the N2O emission factor to less than 1% for a nitrogen removal rate of 0.6 gN-NH4 + .L -1 .d -1 . Moreover, N2O emission factor could be limited to less than 0.2% for a nitrogen removal rate of 0.3 gN-NH4 + .L -1 .d -1 .

Figure 1 .

 1 Figure 1. (a) Evolution of N-NH4 + , N-NO2 -, DO, pH and N2O during nitrification and denitrification. (b) Evolution of DO, ORP, the first derivative of DO and the second derivative of ORP during nitrification and denitrification.

Figure 2 .

 2 Figure 2. Statistic distributions of the evolution of the nitrogen removal rate (a) and N2O emission factor (b) during four different case studies. The median is represented with red thick lines, the average with green crosses, extreme values with green triangles, whiskers with vertical black lines delimited with horizontal black lines on top and bottom.

Figure 3 .

 3 Figure 3. Maximum free nitrous acid concentration (a) and average DO during nitrification (c) versus N2O EF for case studies A, B, C and D. Statistic distributions of the free nitrous acid concentration (b) and DO (d) during case studies. Average pH during nitrification versus DO (e) during case studies. N2O EF versus nitrogen removal rate for case studies B, C and D (constant nitrogen removal rate with a high N2O EF variation during phase A). Lines correspond to trend curves relative to the different data series. Each point corresponds to a SBR cycle. For statistic distributions, the median is represented with red thick lines, the average with green crosses, extreme values with green triangles, whiskers with vertical black lines delimited with horizontal black lines on top and bottom.

Table 1 . Characteristics of four successive case studies (with subcase) of the operating SBR.

 1

	Case study	Automatic control	Cycle configuration	Step-feed	NH4 + removed (mgN.L -1 )	COD/N	Qgas (L.h -1 )
	A	NO	1	1	55.0	4.5	70
	B1 B2	YES	1	1	55.0	4.5 3.4	111.4
	C1 C2	YES	2	5	68.7	4.1 6.5	54.8 26.1
	D1 D2 D3	YES	2	1	51.0 36.0	5.0 8.4	111.4 83.1

Acknowledgements

Mathieu Pocquet was financially supported by the French National Research Agency (ANR) and the project was co-funded by Région Midi-Pyrénées. The authors would like to thank E. Mengelle, M. Bounouba, D. Delagnes for their contribution.


Yann Goueffon

Laurent Arurault
email: arurault@chimie.ups-tlse.fr

Catherine Mabru

Claire Tonon

Pascale Guigue

Black anodic coatings for space applications: study of the process parameters, characteristics and mechanical properties

Keywords: aluminium, black anodizing, thermal control, microcracking, optical treatment

Black inorganic anodized aluminium alloys are used for managing passive thermal control on spacecraft and for avoiding stray light in optical equipment. Spalling of these coatings has sometimes been observed after thermal cycling on 2XXX and 7XXX aluminium alloys. This phenomenon could generate particulate contamination in satellites and may affect mission lifetime. In this work, the influences of the four main steps of the process (pretreatments, sulphuric anodizing, colouring and sealing) on the coating characteristics have been studied for a 7175 T7351 aluminium alloy. The chemical heterogeneity of the coating has been underlined, and its mechanical behaviour observed through crazing. Scratch-testing, used to evaluate coating adhesion to its substrate, revealed the negative impact of thermal cycling.

Introduction

Due to the vacuum of space, thermal regulation of satellites is passively managed by radiative exchange between its external surfaces and the environment. Radiation interactions occur with the Earth, the moon, deep cold space, the sun and the external parts of the satellite. Satellite temperatures in a space environment are often passively controlled by thermo-optical properties of suitable surfaces, ie having convenient solar absorptance (α s ) and emittance (ε). Thus, the utility of black coatings is their α s /ε ratio close to one, which allows passive thermal control of equipment. Black coatings are also useful to avoid stray light in optical instruments. They can consist of painted surfaces [START_REF] Mccroskey | Aeroglaze Z306 black paint for cryogenic telescope use: outgassing and water vapor regain[END_REF], metal surfaces [START_REF] Magdy | Black nickel electrodeposition from a modified Watts bath[END_REF], oxide layers obtained by micro-arc oxidation [START_REF] Shrestha | Microstructural and thermo-optical properties of black Keronite PEO coating on aluminium alloy AA7075 for spacecraft materials applications[END_REF] or anodic films coloured with organic or inorganic dyes [START_REF] Levesque | Black anodize as a thermal control coating for space station freedom[END_REF][START_REF] Shih | Study of the black electrolytic coloring of anodized aluminium in cupric sulfate[END_REF]. Black anodic films (α s > 0.93 ; ε n > 0.90) including inorganic dyes are mainly used because of their low cost, their corrosive and wear resistances during storage, as well as the low risk of contaminating the spacecraft's instruments, especially by outgassing. The anodic film is grown by electrochemical oxidation of the metal surface, without addition of any substances. From this point of view, the anodic film is not a coating, but a conversion interface tightly bound to the metallic substrate.

The adhesion concept is not considered in this case. However in 2005, an alert from the European Space Agency (ESA) (ESA Alert, 2005) mentioned many cases of particle detachment from black anodic films, supported especially on 2XXX and 7XXX type aluminium alloys, after thermal cycling. Such particle pollution is very hazardous for the lifetime of the satellite, potentially inducing the disturbance of any optical or mechanical mechanisms.

The aim of this research work is to study the operational parameters of the black inorganic anodizing process, especially the pre-treatment steps, and their impact on the anodic film characteristics and properties. Adhesion measurements will be particularly discussed.

Experimental procedure

The substrate material was the 7175T7351 aluminium alloy. Its chemical composition in weight percent is: 1.62%Cu, 0.12%Fe, 2.42%Mg, 0.01%Mn, 0.06%Si, 5.75%Zn, 0.041%Ti, 0.21%Cr and Al the remainder. Moreover, all chemical compounds used were analytical grade. Aqueous electrolyte solutions were obtained using deionised water.

Process of elaboration

The preparation of black anodic films including inorganic dyes followed the ESA Standard (ESA ECSS-Q-70-03A, 2006) for spacecraft design. The process involved four main consecutive steps: surface pretreatments , anodizing, colouring and sealing.

The alloy sheet (3x20x40mm or 10x20x160mm) was degreased with ethanol and then etched in Na 2 CO 3 (6.2 g/L) and Na 3 PO 4 (12.5 g/L) aqueous mixed solution (pH = 11) for 5 minutes at 93°C and neutralised (this step is often called desmutting) in aqueous HNO 3 (50 %v/v) for 3 minutes at room temperature. The samples were rinsed in distilled water at the end of each step.

The aluminium sheet was then used as anode and a lead plate (3x40x40mm) as counter-electrode in the electrochemical cell. The anodizing was run for 60 minutes in the galvanostatic mode (J a = 1.2±0.1 A/dm²) using a sulphuric acid solution (150g/L) thermally regulated at 20°C. The samples were then immediately rinsed in distilled water.

The anodized part was coloured by immersion in two successive baths: firstly for 15 min. in a purple solution of cobalt acetate (200g/L) regulated at 43 ±2°C, and then for 10 min in a yellow bath of ammonium hydrosulphide (30g/L) at room temperature. The samples were rinsed in distilled water after each step to remove the excess solution.

The sealing step occurred for 25 minutes in a bath regulated at 98 ±2°C and made up of nickel acetate (NiCH 3 COO, 4H 2 O) and boric acid (5g/L both).

Thermal aging

In space, satellites are directly lighted by the sun and then pass into the shadow of Earth, causing thousands of thermal cycles during their lifetime. To simulate the space environment, the ESA Standard (ESA ECSS-Q-70-04A, 1999) recommends performing 100 cycles between -100 and 100°C under vacuum (10 -5 Pa) with dwells times of 5 minutes minimum and a slope of 10°C per minute. These conditions were defined for the general case and all kinds of materials and equipment (from polymers to electronics). In the current paper, we chose to evaluate the samples, when not otherwise specified, after 10 cycles of -140 to 140°C in a Sun Electronic System EC11 environmental chamber. This could be the maximum range of temperature reached by a satellite orbiting earth. A nitrogen atmosphere was used first to dissociate vacuum and thermal effects. Dwells times of 10 minutes and warming/cooling speeds of 10°C/min were used.

Microscopic and chemical analysis

The meso-and nanostructures of the coatings were observed by field-emission gun scanning electron microscopy (FEG-SEM) with a JEOL JSM6700F device. Energy Dispersive X-ray (EDX) spectroscopy analyses were performed on the cross-sections of the samples to identify the chemical composition inside the film, and especially the location of the dyes.

The microstructure of the aluminum substrate was observed by optical microscopy after metallographic attack for 30 seconds in an acid bath (1%vol. HF, 1.5% HCl, 2.5% HNO 3 and 95% of distilled water).

Characterization and mechanical analysis

Surface roughness was measured before and after pretreatment thanks to a perthometer concept using a stylus method. Although the profile of the surface was scanned along a line of 5.6mm, the arithmetic roughness (Ra) was calculated without the first and the last 0.8mm of the profile to avoid edge effects.

Adhesion measurements were evaluated qualitatively using 180° peel tests with tape strengths of 250 and 500g/cm (ESA ECSS-Q-70-13A, 1999) and quantitatively using a scratch-test device (CSM Revetest instrument) with a diamond stylus (Rockwell, 200µm radius tip). Scratch tests were configured with an increasing normal load from 1 to 30N, a loading speed of 30N/min and an advance speed of 5mm/min. The resulting scratch-print of about 5 mm length was then observed by SEM to identify the mechanisms of degradation and to determine the corresponding normal load.

Four-point bending tests were also performed to generate stresses in the coating and then to observe the possible formation and evolution of cracks. Tests were done on a tensile testing device (Adamel DY 26 of 100kN) with a distance between supports of 40mm for the face under compression and 120mm for the face under tension. Samples (10x20x160mm) were observed with the SEM after loading.

Results and discussion

The physical and chemical characteristics of the material surface, especially the microstructure and the porosity, were studied during the different steps of the preparation process with a view to explaining the final interface properties. Moreover, the aluminium alloy included various irregularly shaped constituent particles, in the range of some microns (Fig. 2a). The EDX analysis revealed that particles contained Al, Fe, Cu, Zn, Mg, in accordance with previous works (Savas and Hearthman, 2008;[START_REF] Gao | An analytical electron microscopy study of constituent particles in commercial 7075-T6 and 2024-T3 alloys[END_REF] that identified for example Al 23 CuFe phases [START_REF] Gao | An analytical electron microscopy study of constituent particles in commercial 7075-T6 and 2024-T3 alloys[END_REF] in similar aluminium alloys. The sample showed a metallic appearance, while the arithmetic roughness (R a ) was (0.9±0.1) µm.

The etching in the Na 2 CO 3 / Na 3 PO 4 aqueous solution blackened the sample. At a mesoscopic scale a tortuous surface appeared (Fig. 2b) with long spots including increasing contents of Zn, Mg and P. The high pH of the etching solution (pH = 11) favours both the chemical dissolution of the "type A" particles [START_REF] Gao | An analytical electron microscopy study of constituent particles in commercial 7075-T6 and 2024-T3 alloys[END_REF][START_REF] Buchheit | A Compilation of Corrosion Potentials Reported for Intermetallic Phases in Aluminum Alloy[END_REF] and the immediate re-precipitation of their corresponding hydroxides and/or phosphates at the metal surface.

The desmutting step in nitric acid solution dissolved these precipitates at the surface in acidic conditions, leaving a clean surface with abundant irregularly shaped pores a few microns across (Fig. 2c). These pores explain that R a then increased to (1.8 ±0.1) µm.

The anodizing step

Fig. 2d shows a typical top view of the anodized sample at the mesoscopic scale, while Fig. 3 reports the porosity at higher magnification. The growth of the porous anodic film is usually considered to be a complex process resulting from antagonistic reactions: the electrochemical oxidation of the metal substrate to the Al 2 O 3 film, and the chemical dissolution of the anodic layer in highly acidic conditions. The characteristics of the resulting porosity depend on various operational parameters, such as the aluminium alloy used [START_REF] Konieczny | The influence of cast method and anodizing parameters on structure and layer thickness of aluminium alloys[END_REF], the experimental anodizing conditions (content, stirring and the bath temperature [START_REF] Aerts | Influence of the anodizing temperature on the porosity and the mechanical properties of the porous anodic film[END_REF]) and the electrical conditions [START_REF] Regone | Al-based anodic films structure observation using field emission gun scanning electron microscopy[END_REF]. Our results show that the anodic film on AA 7175T7351, includes an erratic spongy structure, clearly different from Keller's long-established model [START_REF] Keller | Structural Features of Oxide Coatings on Aluminum[END_REF] made up of regular hexagonal cells, each including a straight central pore perpendicular to the initial substrate surface. This difference is explained by the alloying elements and especially the number and the nature of the constituent microparticles. The usual detailed explanation is based on changes of local current distributions due to potential differences between the metal matrix and the intermetallic particles.

The thickness of anodic films obtained in sulphuric based electrolytes can vary from 2 to 300µm [START_REF] Arurault | Pilling-Bedworth ratio of thick anodic aluminium porous films prepared at high voltages in H 2 SO 4 based electrolyte[END_REF]. In the conditions used here ([H 2 SO 4 ] = 150g/L at 20°C, J a = 1.2 ±0.1 A/dm² for 60 min.), the resulting thickness was about 20µm, i.e. in accordance with the ESA specification (between 10 and 35µm) for spacecraft design (ESA ECSS-Q-70-03A, 2006).

3.1.3. The sealing step EDX analysis (Fig. 4) shows the distribution of cobalt, nickel and sulphur in the cross-section of the anodic film of a coloured and sealed sample. [START_REF] Thompson | Porous anodic alumina: fabrication, characterization and applications[END_REF] previously indicated that the anodic films obtained in sulphuric based electrolyte have a sulphur content around 12-17 weight%. The second possible origin is the introduction of sulphur during the colouring (second)

step; the increase of sulphur comes from the precipitation of cobalt sulphide taking place inside the pores during the second step of colouring [START_REF] Sharma | Anodizing and inorganic black coloring of aluminium alloys for space applications[END_REF] according to:

Co(CH 3 COO) 2 + (NH 4 ) 2 S → CoS + 2 CH 3 COONH 4 (1)
This second origin was confirmed by the presence of cobalt concentrated near the surface (Fig. 4), black CoS being formed in the upper part of the pores due to the immersion colouring process.

Nickel was only detected at less than five microns under the surface. This element comes from the sealing bath but its penetration is limited by the presence of CoS particles in the upper part of the pores.

The resulting surface of a coloured and sealed sample was also studied by FEG-SEM at different scales (Fig. 5a and5b). The plan views reveal no pores on the sample surface but many microcracks at the mesoscopic scale. The mechanism of hydrothermal sealing, usually occurring in pure boiling water, is mainly considered as the hydration of the anodic film, converted into pseudo-boehmite [START_REF] Alwitt | Anodized aluminium coatings for thermal control, Part I: Coating process and stresses[END_REF]:

Al 2 O 3 + (n+1) H 2 O → 2AlOOH + n H 2 O (n>1) (2)
In this case, the pore closing results from the expansion of the molar volumes, as well as the additional precipitation from the nickel acetate. So, the aim of the sealing step is to close the pores, especially their upper part (Fig. 5a and5b), to increase efficiency against corrosion and wear resistance, and protect the dyes against the external atmosphere.

But, the main problem remains understanding the origin of the microcracks and their correlation with the mechanical properties in general and with the potential flaking of the final coating in particular.

Mechanical behaviour

The anodizing process is an electrochemical conversion of the metal surface and does not involve the addition of external material. That is the reason why anodic films are considered sufficiently adherent for the usual applications so the concept of adhesion has not been studied much.

The coloured sealed films were studied from the mechanical point of view, before and after thermal cycling. This cycling simulated the space environment, potentially inducing internal stress, cracking and ultimately flaking of the coatings.

Preparation process, storage and coating stresses

Whatever the process used to prepare the coatings (PVD, CVD...), they usually have residual stresses. For example, Alwitt et al. [START_REF] Alwitt | Anodized aluminium coatings for thermal control, Part I: Coating process and stresses[END_REF]Alwitt et al., 1993) showed that, before sealing, non-coloured anodic films on 5657 aluminium alloy anodized in sulphuric acid have tensile stresses around 30 MPa, depending on the substrate and the process parameters, but not on the thickness. However, the sealing step seems to be much more critical because the thermal expansion coefficient is roughly five times greater in the substrate than in the coating (Alwitt et al., 1992). The sealing bath, regulated at 99°C, caused expansion of the aluminium alloy and subsequent tensile stress in the anodic film. However, once out of the sealing bath, they noted that residual stresses were largely compressive in the sealed films, increasing with sealing time and the thickness of the film.

The influence of humidity on residual stresses of the uncoloured films during long-term exposure to various humidity levels at room temperature were also studied to simulate the effects of space device storage [START_REF] Alwitt | Anodized aluminium coatings for thermal control, Part I: Coating process and stresses[END_REF]. Aluminium oxide is indeed hygroscopic, and the amount of water adsorbed depends on humidity and temperature. At a level of humidity of 0.02 ppm water, coating stresses are greater than 100MPa tensile while they are between 10 and 30MPa compressive at ambient humidity. On the contrary, a decrease of humidity dehydrates the sealed film, contracting the oxide and inducing tensile stresses. [START_REF] Alwitt | Anodized aluminium coatings for thermal control, Part I: Coating process and stresses[END_REF] particularly noticed that the effects produced by low humidity (0.025ppm) on films grown on the 3XXX and 5XXX series aluminium alloys are reversible and comparable to the effects of a vacuum (10 -6 Torr).

Cracking

Due to the difference of thermal expansion coefficient between the anodic film and the substrate, heating induces tensile stresses in the coating that can lead to cracking if the fracture limit is reached.

On the anodized 7175 T7351 aluminium alloy, few cracks were observed after the colouring step performed at 43°C. In all the cases, coloured or not, the samples always cracked just after the 99°C sealing step as shown in Fig. 5a. However, on the 5657 aluminium alloy studied by Alwitt et al [START_REF] Alwitt | Anodized aluminium coatings for thermal control, Part I: Coating process and stresses[END_REF], no crazing was observed after preparation in spite of high tensile stresses detected during the sealing step. [START_REF] Alwitt | Anodized aluminium coatings for thermal control, Part I: Coating process and stresses[END_REF] studied the crazing of anodic sulphuric films, sealed but not coloured, on 3XXX and 5XXX series aluminium alloys during heating. They noted that in most cases the first cracks appeared between 100°C and 200°C. But they also noticed that for specimens stored in a damp atmosphere, this critical temperature decreases to around 70°C (Alwitt et al., 1993). In addition, Alwitt et al. (Alwitt and McClung, 1993) found that the crazing temperature depends on the aluminium alloy used, and for a given alloy, on the heat treatment it experienced. Thus, the lower (lower than 100°C) crazing temperature of anodic films produced in sulphuric acid on the 7175 T7351 alloy compared to 3XXX and 5XXX aluminium alloys can be explained. Indeed, alloying elements, and especially copper, change the growth mechanisms of the anodic film; thus the shape of the pores as well as the chemical composition of the anodic films differ from one alloy to another (Garcia-Veraga et al., 2006;[START_REF] Tsangaraki-Kaplanoglou | Effect of alloy types on the anodizing process of aluminium[END_REF][START_REF] Liu | Grain orientation effects on copper enrichment and oxygen generation during anodizing of al-1at[END_REF][START_REF] Shimizu | The influence of θ' precipitates on the anodizing behaviour of binary Al-Cu alloys[END_REF], changing the mechanical characteristics. For example, the interface between the film and its substrate can be modified by the presence of a layer around 2 nm thick, rich in copper (nearly 40 at. %) [START_REF] Skeldon | Interraction of alloying elements during anodizing of dilute Al-Au-Cu and Al-W-Zn alloys and consequences for film growth[END_REF].

Finally, in the case of coloured films, the dye content of the pores during sealing must also change the thermo-mechanical properties of the anodic film.

Fig. 6 -SEM plan views of a coloured and sealed anodic film (a) before thermal cycling (b) after 10 thermal cycles (-140°C/+140°C). Fig. 6 reveals that thermal cycling (10 cycles between -140°C/+140°C) had no influence on the number of cracks but clearly increased the cracks' width. This can be attributed to a combined effect of dehydration and thermal differential stresses, higher temperatures (compared with the sealing temperature) inducing higher tensile stresses in the anodic film. These results are partially in agreement with similar works (Alwitt et al., 1993) that reported no significant influence of minimal temperature on the crazing while the maximum temperature was found to be highly critical. However, Alwit et al. (Alwitt et al., 1993) showed that the number of cracks always increased as a function of the number of cycles performed on the sample, which is not the case for the present anodized 7175 T7351 aluminium alloy.

The evolution of cracking was observed under mechanical loading, using four-point bending tests to put coatings under both tensile and compressive stresses and to then evaluate adhesion, as done in previous works with other coatings [START_REF] Richard | Four-point bending tests of thermally produced Wc-Co coatings[END_REF][START_REF] Dalmas | Characterization of adherence and cracking within coated materials by an acoustic emission method : application to a WC-Co coating on a steel substrate[END_REF]Ollendorf and Scheinder, 1999]. SEM observations (Fig. 7a and Fig. 7b) of a strictly identical sample area were performed before and after applying tensile stress. At low loadings (about 2kN), new cracks, mainly perpendicular to the tensile direction of stress, appeared. The substrate was then under elastic strain. Higher loadings (about 10kN; plastic strain of the substrate) did not cause new cracks. However, it is interesting to note that, during bending, when plastic strains were reached, the face under compressive stress exhibited film flaking whereas the other (under tensile stresses) did not. In particular, this means that compression stresses induced crack propagation along the interface leading to particle detachment, i.e. flaking.

Flaking

The major risk for space applications is not simply the cracking of the surface layer but flaking, because the particles released could settle on cold parts, typically mirrors, lenses or mechanisms.

Then the satellite instruments would be damaged decreasing the mission's lifetime. In addition, a high level of flaking could change the thermo-optical properties of the surface affecting thermal control.

Fig. 8 -Black anodic coating flaked after pulling off a tape for securing a thermocouple Until now, flaking was always observed on parts that had been thermally cycled. The loss of adhesion was revealed for example on pulling off tapes used to fix thermocouples to the surface during thermal cycles (Fig. 8). However, few cases of self-flaking were observed: particles spontaneously separated from the substrate only under thermal stress. In all cases flaking occurred on crazed surfaces when the islets delineated by cracks came away. Flaking is thus directly linked to crazing and was previously observed on aluminium alloys especially in the 2XXX and 7XXX series (ESA Alert, 2005).

Peel tests highlighted the influence of thermal cycling on flaking. These tests revealed that black anodized parts, thermally cycled five times between -70 and +100°C, showed a "low removal" of the coating with 250 g/cm tape and "moderate removal" at 500 g/cm. After five cycles between -140°C and +140°C, the removal of the oxide increased at 250 g/cm, while there was complete film removal at 500 g/cm and self-flaking was also observed (without any pulling forces applied).

Nevertheless, this test is only qualitative and comparative because the results were strongly dependent on the type of tape used and on the experimental conditions. Thus, other methods to evaluate adhesion quantitatively are necessary.

The scratch test is particularly suited for hard coatings on soft substrates [START_REF] Darque-Ceretti | Adhésion -Adhérence[END_REF][START_REF] Bull | Overview of the potential of quantitative coating adhesion measurement by scratch testing[END_REF] and is used for instance to study the adhesion of various coatings including TiN [START_REF] Holmberg | A mode l for stresses, crack generation and fracture toughness calculation in scratched TiN-coated steel surfaces[END_REF] or CN [START_REF] Liu | Determination of adhesion energy of CNx thin film on silicon from micro-scratch testing[END_REF]. Increasing the normal load of the stylus, a critical load can be measured when failure of the coating occurs.

Numerous failure modes for thin films are known [START_REF] Bull | Failure mode maps in the thin film scratch adhesion test[END_REF] and models have been developed

to calculate adhesion energy thanks to the critical load [START_REF] Laugier | An energy approach to the adhesion of coatings using the scratch test[END_REF][START_REF] Rickerby | A review of the methods for the measurement of coating-substrate adhesion[END_REF]. For the present black anodized 7175 T7351 aluminium alloy, scratch tests showed that the critical load was always lower after thermal cycling than before as shown for instance in Fig. 9 (11 and 8 N, before and after thermal cycling, respectively). This therefore provides quantitative confirmation that thermal cycling induces a decrease in the adhesion of the anodic film to the alloy. al., 2007;[START_REF] Baleix | Oxidation and oxide spallation of heat resistant cast steels for superplastic forming dies[END_REF][START_REF] Lours | Direct observations and analysis of the spallation of alumina scales grown on PM2000 alloy[END_REF].

Conclusion

In the current study, we demonstrated that the chemical composition of black inorganic anodic films is heterogeneous, especially due to the colouring and the sealing steps, which only modify the external part of the coating. Nevertheless, on AA 7175 T7351, the coating was spongy throughout its volume due to alloying elements, which modify the mechanical behaviour of the film compared to other AA series. Thermal cycling enlarges cracks and decreases adhesion of the coating according to the scratch-test measurements. Anodic films may crack because of two different phenomena: the difference in the coefficient of thermal expansion between substrate and coating and the drying of the coating.

Further investigations are currently in progress to evaluate the impact of the anodizing and colouring parameters on the characteristics of the film and especially its adhesion. In addition, the roles of humidity and temperature on film cracking during the process and thermal cycling are currently being studied in order to accurately describe the flaking mechanism.

Figures Captions

Fig. 1 -Optical microscopic 3D views of the AA 7175T7351 substrate after degreasing and metallographic attack.

3. 1

 1 Fig.1shows the initial surface of the AA 7175T7351 just after degreasing and metallographic

Fig. 2 -

 2 Fig. 2 -SEM plan views of AA 7175T7 at different stages of preparation (a) after degreasing (b) after etching (c) after deoxidation (d) after anodizing

Fig. 4 -

 4 Fig. 4 -EDX analysis of coloured and sealed anodic film on the AA 7175T7351.

Fig. 5 -

 5 Fig. 5 -FEG-SEM plan views of a coloured and sealed anodic film on the AA 7175T7351 (a) x 200 ; (b) x 60,000.

Fig. 7 -

 7 Fig. 7 -SEM plan view of a strictly identical sample area of a coloured and sealed anodic film (a) before four-point bending (b) after four-point bending

Fig. 9 -

 9 Fig. 9 -Scratch test of a black anodic film followed by SEM (pictures at the top); Acoustic Emission (Left axis) and friction coefficient between stylus and sample (Right Axis) (a) before thermal cycling (b) after 10 thermal cycles (-140°C/+140°C)

Fig. 2 -

 2 Fig. 2 -SEM plan views of AA 7175T7 at different stages of preparation (a) after degreasing (b) after etching (c) after deoxidation (d) after anodizing

Fig. 5 -

 5 Fig. 5 -FEG-SEM plan views of a coloured and sealed anodic film on the AA 7175T7351 (a) x 200 ; (b) x 60,000.

Fig. 6 -

 6 Fig. 6 -SEM plan views of a coloured and sealed anodic film (a) before thermal cycling (b) after 10 thermal cycles (-140°C/+140°C).

Acknowledgements

The authors thank Peter Winterton for his helpful comments.


J Lemetayer

C Lacour

F Lefebvre

C Gobin

A Cessou

B Lecordier

Simultaneous gas and spray PIV measurements in an optical engine

ABSTRACT

Engine internal aerodynamic reveals complex flow involving multi-scale turbulence, flow structure compression, two-phase interactions and are responsible for the mixing process and the engine performance and efficiency. Recent efforts have been made to describe in details in-cylinder flows either with temporal (HR-PIV) or spatial (Tomo-PIV) resolution. However, for gasoline direct injection engine (GDI), the dynamic coupling between gas and spray droplets should also be undertaken for a correct evaluation of in-cylinder aerodynamics. Simultaneous measurements of instantaneous gas and droplets velocities during intake and compression strokes are proposed by means of two-phase PIV based on fluorescence [START_REF] Driscoll | Simultaneous air/fuel phase PIV measurements in a dense fuel spra[END_REF][START_REF] Kosiwczuk | Simultaneous velocity field measurements in two-phase flows for turbulent mixing of sprays by means of two-phase PIV[END_REF]. The technique is adapted to the constraints of optical engine and associated to specific algorithms development for the liquid phase. The engine test bench consists in a mono-cylinder GDI engine (AVL) which operates up to 3000 rpm in optical configuration with a displacement volume of 450 cm 3 and a compression ratio of 8.5. The optical accesses to the combustion chamber is enabled by a quartz-glass liner. The injection system is composed of a solenoid multi-hole injector Bosch fed up by a pressurized volume to ensure a stable injection pressure up to 100 bar. An injector power control module (EFS IPOD) is used to drive the injector and control injection timings in the engine cycle.

The two-phase PIV technique is based on the use of two different dyes dissolved in the seeding particle and gasoline, producing fluorescent emissions on separated spectral bands for each phase [START_REF] Driscoll | Simultaneous air/fuel phase PIV measurements in a dense fuel spra[END_REF][START_REF] Kosiwczuk | Simultaneous velocity field measurements in two-phase flows for turbulent mixing of sprays by means of two-phase PIV[END_REF]. The phase separation is enabled by a detection system consisting of a dichroic window distributing the collection signal on two synchronized PIV cameras (Hamamatsu 12 bits 2018×2048 pixels) equipped with Nikkor lenses (50mm f/#2) and adapted pass-band filters. Angular controls are mounted at the base of the dichroic sheet and both cameras in order to adjust with precision the common camera field of view. A refined adjustment based on a polynomial approach of 5 th degree is then numerically performed from calibration grid images to ensure a perfect images overlap and to correct image distortion induced by the glass liner. Dyes excitation is performed with two lasers at different wavelengths (532 nm for the gas and 355 nm for the spray) in order to independently adjust the PIV acquisition delays to the high velocity shift between phases in the early stage of injection. The use of two wavelength also improve the spectral separation of fluorescence signal and then the phase discrimination. An original synchronisation of the laser and camera with the engine cycle is ensured by means of a programmable time board to get rid of engine speed fluctuations and guarantee a fixed working frequency for the lasers while limiting injection and fouling to the acquisition triggering.

Velocities of the gas and the liquid phases can thus be acquired simultaneously for engine conditions where the two phases are present, typically early after the start of injection [START_REF] Kosiwczuk | Simultaneous velocity field measurements in two-phase flows for turbulent mixing of sprays by means of two-phase PIV[END_REF][START_REF] Lemétayer | Two-phase PIV to investigate spray/flow interaction in Gasoline Direct Injection engine[END_REF]. Prior to the correlation step, a pre-processing of the fluorescence images is performed to enhance the correlation level. A masking technique, with adaptation of the masking surface at each angular position is also used. The PIV post-processing is then adapted to the present configuration with two different algorithms for each phase. The velocity fields of the gaseous phase are obtained by a multi-pass subpixel shift correlation algorithm based on the correlation of the seeding patterns [START_REF] Lecordier | Etude de l'interaction de la propagation d'une flamme prémélangée avec le champ aérodynamique, par association de la tomographie laser et de la vélocimétrie par images de particules[END_REF]. Interrogation window size of 32×32 pixels (1.85×1.85 mm 2 ) with an overlap of 50% has been used with a vectors filtering based on a minimum value of the Signal to Noise Rate (SNR) and on a median filter which are adapted to each experimental condition. This enables to reject most of non-valid vectors. Gas phase velocity calculation for internal engine flow is validated in our configuration by means of simultaneous Mie-based PIV and fluorescence based PIV. Comparison of mean and instantaneous velocities show less than 5 % differences. The density of the liquid phase is heterogeneous with a very dense part near the injector nozzle and a dispersed part after the breaking of the liquid sheet. In the spray dispersed part, velocities are processed with a particle approach, whereas in the dense part of the spray, a specific algorithm based on pattern correlation is developed. Preferential direction and topology of the spray are taken into account through the window shape and size.

ISPIV2017

The 12 th International Symposium on Particle Image Velocimetry June 19 (Mon) -21 (Wed), 2017 / Grand Hotel, Busan, Korea Operating conditions correspond to homogeneous combustion at partial load. The start of injection is fixed at 300 BTDC. The injection duration is adjusted for the engine rotation speed in order to fulfil the unity equivalence ratio. For the present study, measurements have been performed with engine rotation speed of 2000 rpm and injection pressure of 100 bar and 50 bar. Acquisitions are performed during intake and compression phases from 288 BTDC till 50 BTDC. Figure 1 presents the gas velocity fields coloured by the velocity amplitude (left) without injection for an acquisition at 264 BTDC, during the intake phase. The intake valve jet is developed and forms a large-scale rotating structure, giving rise to the tumble motion. The velocity fields measured with injection is presented in the middle row for the same conditions and the corresponding spray velocity is shown on the right row. The comparison of these velocity fields show that the gas velocities are strongly impacted by the presence of the spray injection. The spray dynamics and interactions with the carrier phase must be taken into account for a correct evaluation of engine internal aerodynamics. The tumble positions and amplitudes are extracted from 1 scalar map and analysed during the engine cycle. The tumble trajectory is modified under influence of the spray development and is limited to a vertical displacement. The spray injection also induces a reduction of the tumble intensity reaching 50% at 50 BTDC, near the compression point. The modification of injection pressure shows that the more the pressure increases the more the tumble level is reduced near the compression point. These results put forward that direct injection strongly impacts the internal aerodynamic in the present configuration.

The flow structure modifications will be described in details based on the superimposed spray and gas velocity field in the cylinder volume. Topology of the spray dynamics and its modification due to engine turbulent flow field will also be analysed by means of a comparison of the spray velocity field measured at rest and the one measured in the engine volume.

The whole results obtained in this study will help understanding the dynamic coupling between the engine in-cylinder flow and the high-pressure spray, it will also provide more realistic internal engine aerodynamic in the presence of injection for engine simulation.

Figure 1

 1 Figure 1 Mean velocity fields at Θ = -264 dv for the gas phase without (a) and with (b) injection, for the spray droplets (c).

ACKNOWLEDGEMENTS

This work was conducted as part of the ESSENCYELE project (moteur ESSENCe injection directe hYbride Electrique abordabLE), funded by the Agence De l'Environnement et de la Maîtrise de l'Energie (ADEME). The material support from the EFS Company is acknowledged.


Julien Lemetayer

Corine Lacour

Franck Lefebvre

Carole Gobin

Armelle Cessou

Bertrand Lecordier
email: bertrand.lecordier@coria.fr

Recent developments in two-phases fluorescence PIV: Application to the

ABSTRACT

The purpose of the present work is to develop and to demonstrate the ability of the two-phase PIV by fluorescence technique for the study of two-phase flows dynamics. In particular, the dynamic of a high-pressure spray and its interaction with surrounding air is investigated in this study. For the single-phase flows, the PIV is a well-established technique to measure the velocity of the continuous phase but not easily applicable on two-phase flows without adaptation and new developments. Ten years ago, fluorescence PIV were developed to investigate dynamic of spray injection [START_REF] Driscoll | Simultaneous air/fuel phase PIV measurements in a dense fuel spray[END_REF] [START_REF] Kosiwczuk | Simultaneous velocity field measurements in two-phase flows for turbulent mixing of sprays by means of two-phase PIV[END_REF]. For phase discrimination, each phase is labelled by two different dyes allowing the acquisition of two separate pairs of images which are then analysed separately and provide the instantaneous velocity fields of the two phases. In the present paper, an extension of our previous work is proposed [START_REF] Kosiwczuk | Simultaneous velocity field measurements in two-phase flows for turbulent mixing of sprays by means of two-phase PIV[END_REF] by associating two PIV lasers at 532 nm and 355 nm in order to make separation of fluorescence signals of the two phases more efficient and then to access to velocity measurements in the denser part of the spray. This arrangement permits also an optimisation of the two PIV delays in order to account for the high velocity difference between the phases in high pressure injections. For the imaging system, the two fluorescence signals are collected simultaneously on two 12 bits Hamamatsu cameras (C9300 -4M pixels). Two selected passband filters are placed in front of the cameras, for the discrimination of the fluorescence in each phase. To ensure a perfect overlap of the velocity measurements, image deformation technique based on 5 th order polynomial functions are applied on particle images to correct any differences of magnifications, orientation and optical deformation of images. In the first part of the paper, the selection of a couple of fluorescent dyes adapted to our experimental conditions is addressed. In the second part, a novel PIV measurement based on pattern correlation is proposed to measure the velocity of the liquid phase in the dense part of the liquid jet near the injector nozzle. In the final part, our fluorescence PIV technique is validated on two high pressure gasoline injectors placed in a closed pressurized chamber (single and multi-holes injectors up to 100). bar). The major criteria of dyes selection are: maximum of absorption spectrum close to the laser wavelengths; fluorescence signal in a spectral range close to maximum quantum efficiency of the PIV cameras; high fluorescence efficiency, especially for the dye labelling the seeding particles; distinct fluorescence emission to permit the optical filtering of the two phases; the solubility of the dyes in the liquid phase and the seeding particles. Considering all these criteria and numerous tests, our dye selection has been stopped on Stilbene 420 for the seeding particles and Pyrromethene 597 for the liquid phase. Figure 1 presents instantaneous images in the case of single hole injector without any processing of the fluorescence signal of each dye. These images show that there is no signal from the seeding particles of the gas remaining on the spray image (Figure 1-A). The gas image (Figure 1-B) shows that the spray signal is correctly filtered in spite of the high intensity of droplets Mie scattering and fluorescence. This is confirmed in Figure 1-C by the superimposition of both phases at the location of the red rectangle plotted in Figure 1-B. Indeed, the spray droplets (in red) are clearly distinguished from the seeding particles of the gas (in black).

ISPIV2017

The 12 th International Symposium on Particle Image Velocimetry June 19 (Mon) -21 (Wed), 2017 / Grand Hotel, Busan, Korea For the PIV calculation on the two phases, different strategies have to be adopted. For the gas phase, the images are very similar to those obtained in PIV based on Mie scattering imaging. The velocity fields are thus evaluated using a conventionnal multi-pass crosscorrelation PIV algorithm (32×32 pixels -1.85×1.85 mm 2 , 50 % overlap, validation > 90%). For the liquid phase, two complementary algorithms have been used depending on the local density of spray. In the lowest density region (Zone B -Figure 1-A), a localised PIV cross-correlation on the droplet positions is adopted [START_REF] Kosiwczuk | Simultaneous velocity field measurements in two-phase flows for turbulent mixing of sprays by means of two-phase PIV[END_REF]. In the denser part of the spray, near the injector nozzle (Zone A -Figure 1-A), this approach cannot be applied (no individual particle) and a standard multi-pass PIV algorithm failed to evaluate the velocity of the liquid. In those regions, an original PIV algorithm based on a specific management of the interrogation window taking into account the spray topology (angle, expansion) is proposed (Figure -pattern correlation method). This approach is derived from previous work on flames [START_REF] Balusamy | Direct measurement of local instantaneous laminar burning velocity by a new PIV algorithm[END_REF]. On the spray images, particle and pattern correlation methods are individually applied providing two velocity fields which are combined in the final step of the treatment. In Figure , a sample of an instantaneous velocity field of the liquid phase for a single-hole injector is plotted. The regions of application of each method are easily observable as a function of droplet density. In the region of intermediate density, the two approaches provide identical velocity measurements. The impact of gasoline injection on entrainment of surrounding gas in the external and internal part of spray is investigated for the two injectors to validate the measurements through a comparison with previous studies. An example of instantaneous measurement of both phases in the case of single hole injector is presented in Figure and demonstrates that the two-phase PIV by fluorescence is a powerful tool to measure instantaneously and simultaneously the velocity of the two phases in a dense spray and then to access to the relative spray/gas velocity in 2D which is a fundamental data to characterize the dynamic interactions between phases.

Figure 1 -

 1 Sample of a couple of fluorescence images of spray (A) and seeding (B) signals and their overlap (C)

Figure 4 -Figure 2 -

 42 Figure 4 -Instantaneous velocity fields of the spray (left) and the gas phase (right) 4 ms after the start of injection for a single-hole injector

Figure 3 -

 3 Figure 3 -Instantaneous velocity fields of spray calculated by the particle approach (blue) and by the pattern correlation method (red) 2 ms after the ASOI

ACKNOWLEDGEMENTS

This work was conducted as part of the ESSENCYELE project (moteur ESSENCe injection directe hYbride Electrique abordabLE), funded by the Agence De l'Environnement et de la Maîtrise de l'Energie (ADEME).


Mokrane Malek

Daniel Allano

Bertrand Lecordier

Frédéric Corbin

Gilles Godard

Denis Lebrun
email: denis.lebrun@coria.fr

Frederic Corbin

Digital in-line holography with a divergent beam for 3D velocity field measurement in small volumes

Introduction

Digital in-line holography (DIH) is commonly used to characterize flows seeded by microparticles. These particles are illuminated with a source of coherent light and the diffraction patterns are directly recorded by CCD camera [START_REF] Meng | Holographic particle image velocimetry: from film to digital recording[END_REF], [START_REF] Sheng | Using digital holography microscopy for simultaneous measurements of 3D near wall velocity ans wall shear stress in a turbulent boundary layer[END_REF], [START_REF] Salah | Application of multiple exposure digital in-line holography to particle tracking in a Bénard-von Kármán vortex flow[END_REF], [START_REF] Malek | Digital in line holography system for threedimensional-three components (3D-3C) Particle Tracking Velocimetry[END_REF], [START_REF] Cao | Hybrid digital holography imaging system for 3D Dense Particles field measurement[END_REF], [START_REF] Garcia-Sucerquia | Digital in-line Holographic microscopy[END_REF]]. The reconstruction of particle images is often achieved by applying the Fresnel integral to the diffraction patterns [START_REF] Xu | Tracking particles in four dimensions with inline holographic microscopy[END_REF], [START_REF] Fournier | Digital in-line holography : influence of the reconstruction function on the axial profile of a reconstructed particle image[END_REF]] or by using the fractional Fourier transformation [Coëtmellec et al. (2003)]. In the case of far-field diffraction, this reconstruction can be seen as a wavelet transform (WT) of the hologram transmission function [START_REF] Onural | Diffraction from a wavelet point of view[END_REF]].

The main advantage of DIH is a simple optical arrangement of the recording system composed only by few optical components. This characteristic allows its implementation in several applications, particularly where optical access is difficult. Finally, for more or less all the applications, low-power laser sources such as modulated laser diodes are suitable.

In this paper, we show that DIH can be used to determine 3D velocity fields in small volumes. A diverging reference wave emerges from a fiber coupled laser diode and the curvature radius of the wave leads to a transverse magnification of 4. This configuration allows us to reconstruct images of droplets with a diameter smaller than 5µm and moving with velocities ranging up to 1m/s. The use of divergent beam necessitates an accurate calibration step of the sample volume.

In section 2, the theoretical background used for digital in-line reconstruction of holograms is recalled. In section 3, a description of the proposed method for calibration step is explained. In the same section, an accuracy assessment of this configuration for the axial and transverse localizations is proposed. In the last section, experimental results of 3D particle displacements in air-flow seeded with oil micro-droplets are presented.

Theoretical background : Hologram recording with a diverging beam

The principle of the recording in the case of in-line holography is shown in Fig. 1. A coherent light source illuminates an opaque objet of diameter d located in the plane ( e X , e Y ) at a distance e Z from the CCD sensor plane (X',Y'). In the far-field configuration, the distance between the CCD camera and the source light, noted s Z , can be considered as the curvature radius of the used spherical wave. ) and by assuming that a spherical wave illuminates the object field, the intensity of the diffraction patterns recorded on the CCD camera at distance e Z is described by the following expression [START_REF] Tyler | Fraunhofer Holography Applied to Particle Size Analysis: A Reassessment[END_REF], Vikram (1987) ] :

( ) ( ) r F Z r F Z r Z Y X I eq eq eq eq eq Z d eq Z d eq eq Z 2 2 2 1 sin 2 1 ) ' , ' ( λ λ λ λ π λ ⎟ ⎟ ⎠ ⎞ ⎜ ⎜ ⎝ ⎛ + ⎟ ⎟ ⎠ ⎞ ⎜ ⎜ ⎝ ⎛ - = (1)
where:

e eq KZ Z = , Kd d eq = , e s s Z Z Z K - = , 2 2 2 ' ' Y X r + = and
( )

eq eq eq eq eq Z d Z r d Z r d J d r F eq λ π λ π π λ ⎟ ⎟ ⎠ ⎞ ⎜ ⎜ ⎝ ⎛ = 1 2 2 .
λ is the wavelength and . The object transmittance is

Z Z e Y e Source Z s X' Y' X e
expressed as follows:

⎪ ⎩ ⎪ ⎨ ⎧ ≤ + = otherwise 0, ; 2 if 1, ) , ( 2 2 d Y X Y X O e e e e
(2) Note that equation ( 1) is established by applying the Huygens-Fresnel principle in the scalar diffraction. The first term corresponds to the reference wave; the second describes the interferences between the reference wave and the object diffracted wave. The third term is similar to the diffraction produced by the same aperture as the object. This last term can be neglected in the case of the Fraunhofer approximation [START_REF] Buraga-Lefebvre | Application of wavelet transform to hologram analysis: three-dimensional location of particles[END_REF]].

The diffraction pattern expressed by the equation ( 1) shows that the recording by the spherical wave can be modeled as a collimated recording configuration (see Fig. 2). The equivalence between both configurations is obtained by applying a magnification term ( By this equivalence, the hologram formed by a particle of diameter d, illuminated by a divergent wave, and located at a distance e Z from the CCD sensor, is similar to the hologram generated by a particle of diameter d K × illuminated by a plane wave and located at a distance e Z K × from the CCD. This interesting result leads to a very simple way for the processing of digital holograms. It means that the diffraction pattern produced by a given object can be processed exactly as if the hologram was recorded with a plane wave and all the conditions of reconstruction (spatial sampling, recording distance, spatial resolution) being expressed in an equivalent magnified space. This principle is illustrated in conventional optical holography representation by Fig. 3. axis systems:

⎪ ⎪ ⎩ ⎪ ⎪ ⎨ ⎧ = = = K Z Z K Y Y K X X r e r e r e
, where conjuguate planes lead to:

s s r Z Z Z K + =
(3)

In our case, the holograms are reconstructed by a wavelet transform (see [START_REF] Buraga-Lefebvre | Application of wavelet transform to hologram analysis: three-dimensional location of particles[END_REF] for more details). A sample volume is reconstructed plane-by-plane by computing, for different scale parameters, the WT of the recorded diffraction pattern. Each scale parameter of the wavelet is directly related to the axial position r Z . In the reconstruction space, the axial position of particle image is obtained by the Maximum of the Wavelet Transform Modulus (MMWT) according optical axis positions (i.e. r Z ).

Calibration method

In practice, the distance between the CCD sensor and the light source (i.e. s Z ) has to be accurately estimated by an appropriate calibration method.

This condition gives

] [i e Z = A× ] [i e Z ∆
+B. The estimated value of Z s is obtained iteratively as described below by a progressive forcing of the parameter A to tend toward 1, and B is an estimated value of Z e0 .

In this work the calibration plate is composed of five opaque disks of diameter 10µm deposed on the surface of an optical glass plate. Four disks are placed at the corner of a rectangle of 1mm by 0.7mm, the fifth disk being centered to the others. An example of disk images at reconstructed Z r = 464.75mm is shown in Fig. 3.a. These images are reconstructed by using the hologram recorded at the reference position. The intensity profiles, according X r and Z r coordinates, of the disk noted P5 are respectively shown in Fig. 3

[i e Z ∆ , compute ] [i e Z ) by ] [ 0 ] [ 0 ] [ i rj s i rj s i e Z Z Z Z Z + = )
) )

; 3) Compute the coefficients A and B of the linear regression

B Z A Z i e i e + ∆ × = ] [ ] [ ) 4) Increment the value of s s s Z Z Z ∆ + ← 0 0 ) ) ; 5) Repeat the steps 2, 3 and 4, until ε > -1 A .
In this calibration method, ∆Z s =10 -3 mm and ε=0.001 have been imposed.

The estimated value of s Z is obtained after satisfaction of the stop condition (i.e. step 5 of the above algorithm) and a reference axial position (i.e. 0 e Z ) is deduced from the parameter B of the linear regression.

The Fig. 4 shows the histogram of the measured depth-error δZ e and x-y error (i.e.δx and δy) obtained by comparing the measured locations with the imposed x-y-z positions. For this experiment, 21 holograms have been recorded and 10 reconstructed images have been processed for each hologram. From these results, the mean square errors are estimated to 7.62µm in depth and to 1.60µm in the x-y direction.

Experimental results

Experimental setup

In the present work, the experimental setup presented in Fig. 5 has been used to validate our DIH system in real experimental conditions. The divergent source light comes from an optical fiber coupled with a modulated laser diode (CUBE 640-25FP) of wavelength λ= 640nm. In continuous mode, this laser delivers a maximum power of 25mW and a TTL signal is used to drive it in pulsed mode. The diffraction patterns are recorded on the interline transfer CCD chip of a Hamamatsu camera (C9300-024). Here, a single frame with a double exposure is used to record the hologram. The size of recorded images is 2048×2048 pixels in 8 bits and the pixel size is equal to 7.4µm. The studied flow is a laminar air jet of 8 mm in diameter (D). This jet, seeded with fine oil droplets in a diameter range between 1 to 5µm, is produced at the outlet port of a porous media. The flow is protected from external environment by a stream of filtered air of 100 mm in diameter in order to avoid any unwanted particles or dust during recording (see Fig. 5). After the reconstruction and 3D localization of particle images, the couples of particle are identified and the displacement of each one is measured.

The first step of the experiment has been to calibrate our recording setup by using the method presented previously in the section 3. The estimated value of the distance between the light source and the CCD sensor is 125 = s Z mm with an optical magnification ranging between 3 and 5. In that configuration, the sample volume of particle field is about 4×4×10 mm 3 .

3D velocity measurement

In order to validate the potential of our approach for the velocity measurement of small oil droplets, several ten of holograms with different velocity ranges and particle concentration have been recorded. In this experience, the electronic shutter of the CCD camera is set to 110µs to reduce the effect of the background light. The time delay between two laser pulses and pulse duration are respectively equal to 10µs and 1µs. The Figure 6 shows an example of reconstructed image. On a magnification of this image (Fig. 6.b), two couples of particle images are clearly observed and two velocity vectors are then measured.

Conclusion

Digital in line holography using a divergent light emerging from a fiber coupled laser diode is well suited for recording holograms of micrometric seeding particles. This capability is obtained by the magnification ratio produced by the curvature radius of the reference wave. The digital reconstruction of particle images by wavelet transform is used to obtain 3D instantaneous spatial distribution of the particle field. This method has been validated in a laminar air jet seeded with oil droplets in the diameter range of 1-5µm and with velocities about 1 m.s -1 . It is a significant advance in applications of digital in-line holography in fluid mechanics, where the size of seeding particles is often few pixels in diameter to obtain diffraction patterns with significant visibility. The experimental results obtained in the case of a laminar jet are in good agreement with our expectations and shows the possibilities of this method for instantaneous evaluation of the velocity fields in a sample volume of about 1mm 3 .

Fig. 1 :

 1 Fig. 1: Optical configuration of recording in-line holograms with a divergent beam.

Fig. 2 :

 2 Fig. 2: Hologram recording under Gabor configuration (collimated case): equivalence with the spherical wave configuration showed in Fig. 1.

Fig. 3 :

 3 Fig. 3: Digital in-line holography with a spherical recording wave. Equivalent coordinate system under plane wave configuration

For

  the calibration, a set of diffraction patterns is recorded from known objects successively localized at different axial positions. Each object position ] [i e Z is adjusted by a micrometer translation stage to a well-defined distance ] [i e Z ∆ centered on the reference plane ( 0 e Z ).The objective of this calibration step is to estimate the value of s Z by assuming a linear

  Fig. 3: a) Example of reconstructed disk images at Z r =464.75mm. d = 10 µm, b) X r profile and c) the Z r profile of disk P5.

]

Fig. 4 :

 4 Fig. 4: Estimation of location accuracy by processing 21 holograms of calibrated disks. a) Histograms of deptherror

Fig. 5 :

 5 Fig. 5: Recording system with divergent light.

Fig. 6 :

 6 Fig. 6: Example of reconstructed images. Here, particles are located at a distance mm Z e 98 = from the

Fig. 7 :

 7 Fig. 7: Evolution of the normalized Y-velocity (

15th Int. Symp on Appl. Laser Techniques to Fluid Mechanics, Lisbon, Portugal, July 05 -08, 2010

Acknowledgement:

This work has been supported by the French National Agency (ANR) in the frame of the program : VIVE3D (ANR-07-1188-532) "Vélocimétrie Instantanée Volumique pour les Ecoulements Tridimensionnels''.