M Belhaj

P Guérin

M E Zaïm

L Abdeljalil

Influence of the frequential identification tests on the induction machine modelling

Keywords:

An electric network is constituted by generators providing power to various loads. The induction machine is the most common motor in the industrial applications because of its simplicity of construction, its reliability and its low cost. It allows the constant speed applications when it is directly supplied by the network and variable speed ones with the use of power converters. These speed drives induce important harmonic perturbations in the power system and in the motor side. In this context, we are interested by the frequential modelling of the induction machine in order to simulate the line current with its harmonic components. In order to use the induction machine models, it is necessary to identify their parameters and therefore realizing specific and sufficiently informative tests. Several tests of identification are presented in the literature; their differences and their relevance have not been discussed. This study deals with the realization and the comparison of several types of frequential tests (constant voltage, constant V/f, injection of DC current) for induction machine models identification. The comparison of the simulation results with the experimental measurements allows electing the most representative of the induction machine behaviour.

Introduction

With the widespread use of power electronics in the electrical equipments, the absorbed currents are less and less sinusoidal. The number and the unit power of the non-linear loads lead to network voltage disturbance, especially due to the injection of the harmonic perturbations. In the context of a decentralized production, with the possibility of isolated working, these phenomena are easily increased and can then seriously compromise the quality and the availability of the distributed energy. From these different facts, there is an interest of analyzing accurately the interactions between the non-linear loads and the power system.

The induction machine is the most common motor in the industrial applications. It allows the constant speed applications when it is directly supplied by the network and variable speed ones with the use of power converters. These speed drives induce important harmonic perturbations in the power system and in the motor side.

Consequently, the analysis and the simulation of low frequency disturbances [50Hz -1kHz] require an adapted modelling of the machine and its supply. This paper proposes a circuit type frequential model of the induction machine in view of the electrical network simulation. This model allows the reproduction of the main low frequency components of the line current when the machine is supplied by non-sinusoidal voltages. Some of the models studied in the literature use integer orders based on resistances and inductances circuits in parallel or in ladder [START_REF] Vermaelen | Contribution à la modélisation et à la réduction des perturbations conduits dans les systèmes d'entraînement à vitesse variable[END_REF], [START_REF] Kabbaj | Skin effect characterization in a squirrel cage induction machine[END_REF]. Others use non integer orders based on the integration of fractional order impedance in the rotor circuit [START_REF] Riu | Modélisation des courants induits dans les machines électriques par des systèmes d'ordre[END_REF], [START_REF] Canat | Fractional order: frequential parametric identification of the skin effect in the rotor bar of squirrel cage induction machine[END_REF]. In [START_REF] Belhaj | Modélisation fréquentielle de la machine asynchrone en vue de l'analyse des perturbations conduites basses fréquences[END_REF] various induction machine models are compared and their performances and their validity domains are discussed.

In order to use any model, its parameters have to be easily determined. In the literature [START_REF] Vermaelen | Contribution à la modélisation et à la réduction des perturbations conduits dans les systèmes d'entraînement à vitesse variable[END_REF], [START_REF] Canat | Fractional order: frequential parametric identification of the skin effect in the rotor bar of squirrel cage induction machine[END_REF], [6] and [START_REF] Dedene | Estimation of a Global Synchronous Machine Model Using a Multiple-Input Multiple-Output Estimator[END_REF], several procedures of identification tests are proposed. They are realized with blocked rotor (harmonic excitation). Among these measurement methods, it is necessary to choose the most representative, thus, the one that allows the parameters determining in the nearest conditions to the machine real working.

The aim of this work is identifying the best method, thus each one is studied and its representativity is tested using an experimental benchmark. Which is constituted by an industrial 5.5kW induction machine and its variable speed drive. This benchmark allows observing the real phenomena appearing in the machine supplied by non sinusoidal voltage. It is also useful as reference in all the identification phases and for the models validation.

The first part of this paper presents briefly the methods of the induction machine frequential modelling presented in the literature. The most representative model will be used for the identification methods comparison [START_REF] Belhaj | Modélisation fréquentielle de la machine asynchrone en vue de l'analyse des perturbations conduites basses fréquences[END_REF].

In the second part, the identification measurements are treated. We begin by the description of classical and frequential tests. The measurement methods are then compared and the machine representativity is discussed.

The effect of the induction machine non-linearity on the frequential behaviour and the importance of the magnetic flux level during the tests are highlighted.

In the last part, a genetic algorithm is used to determine the parameters of the machine model. The line current computed by the induction machine model is compared to the measurements for each identification method and the efficiency of the various identification measurements is discussed.

Frequential modelling of the induction machine

Concerning the induction machine modelling, several studies have showed that the generalized equivalent circuits are necessary for harmonic analysis [START_REF] Kabbaj | Skin effect characterization in a squirrel cage induction machine[END_REF], [START_REF] Viarouge | Modélisation par circuits équivalents généralises et identification des machines électriques sur un domaine de fréquence étendu[END_REF], [START_REF] Sudhoff | An induction machine model for predicting inverter machine interaction[END_REF]. In this context, these models take into account the frequency effect and two families can be identified :

-Integer order models [START_REF] Alwash | Generalised approach to the analysis of asymmetrical three-phase induction motors[END_REF], [START_REF] Smith | A transient induction motor model including saturation and deep bar effect[END_REF]: They are based on the fundamental model of the induction machine with the rotor conductors represented by RL cells representing the massive conductor [START_REF] Kabbaj | Skin effect characterization in a squirrel cage induction machine[END_REF], [START_REF] Klingshirn | Simulation of polyphase induction machine with deep rotor bars[END_REF].

-Non integer order Models [START_REF] Kabbaj | Skin effect characterization in a squirrel cage induction machine[END_REF], [START_REF] Canat | Fractional order: frequential parametric identification of the skin effect in the rotor bar of squirrel cage induction machine[END_REF], [START_REF] Kamwa | On equivalent circuit structures for empirical modeling of turbine-generator[END_REF] : The application of non integer orders in machine modelling has been used in several studies [START_REF] Lin | Modélisation et identification d'ordre non entier. Applications à une machine asynchrone et à un système thermique[END_REF], [START_REF] Benchellal | Identification of a non-integer model of induction machines, Fractional Differentiation And its Applications[END_REF]. The frequency effect on the single bar impedance is studied in [START_REF] Canat | Contribution à la modélisation dynamique d'ordre non entier de la machine asynchrone à cage[END_REF] with a non integer order modelling. In [START_REF] Khaorapapong | Modélisation d'ordre non entier des effets de fréquence dans les barres rotoriques d'une machine asynchrone[END_REF], this approach is extended to the complete model of the induction machine by including a non integer order representation of the rotor.

Otherwise, a model based on the superposition principle and taking into account the influence of the voltage supply has been studied in [START_REF] Belhaj | Modélisation fréquentielle de la machine asynchrone en vue de l'analyse des perturbations conduites basses fréquences[END_REF]. One part of the model is used to describe the fundamental behaviour of the induction machine and another one represents the harmonic contributions (Fig 1). The summation of the resulting currents gives the machine line current. The fundamental component is represented by the classic single-phase model of the induction machine (Fig 1_a). The harmonic components are obtained with the help of RL circuits representing the massive conductor behaviour [START_REF] Kabbaj | Skin effect characterization in a squirrel cage induction machine[END_REF][12] (Fig 1_b). The superposition principle induces that the fundamental and harmonics phenomena are independent. The advantage of this model is to eliminate the harmonic influence if the induction machine is supplied by pure sinusoidal voltages. Moreover, the experimental identification tests of the two parts of the model, fundamental and harmonics, are completely dissociated. The fundamental model parameters can be identified by usual measurements at no load and at blocked rotor for the nominal frequency of the machine. The harmonic model must be determined with another specific experiments that must be enough informative. According to the study in [START_REF] Belhaj | Modélisation fréquentielle de la machine asynchrone en vue de l'analyse des perturbations conduites basses fréquences[END_REF], this model represents correctly the induction machine frequential behaviour around its operating point thus, it will be used in the following for comparing the identification methods.

Measurement methods in view of the induction machine frequential models parametrical identification

It is clear that the frequential behaviour of the machine is better described using a frequential approach than a temporal approach [START_REF] Dandeno | Recent Trends and Progress in Synchronous Machine Modelling in the Electric Utility Industry[END_REF]. Indeed, the frequential approach is generally on the one hand easier to apply and it represents less risk for the machine compared to some temporal responses like a step response. On the other hand, the measurements are less sensitive to the noise because the sinusoidal signals are defined at a given frequency and they can be easily decorrelated from the noise.

In the literature [START_REF] Vermaelen | Contribution à la modélisation et à la réduction des perturbations conduits dans les systèmes d'entraînement à vitesse variable[END_REF], [START_REF] Canat | Fractional order: frequential parametric identification of the skin effect in the rotor bar of squirrel cage induction machine[END_REF], [6], [START_REF] Dedene | Estimation of a Global Synchronous Machine Model Using a Multiple-Input Multiple-Output Estimator[END_REF], several tests are used to characterize the frequential behaviour of the induction machine. They are called harmonic excitation tests and are frequently realized with blocked rotor. They lead to the identification of the machine single phase impedance and they describe its evolution according to the frequency. The first one consists in imposing a sinusoidal voltage with constant amplitude and variable frequency. The second test is the SSFR method (StandStill Frequency Response) which imposes constant voltage-frequency ratio. Finally, the third test is based on the injection of a sinusoidal current with a DC component for increasing the magnetic flux in the machine.

Constant voltage and variable frequency

This experiment is done with a blocked rotor and the stator is supplied by a sinusoidal reduced voltage with constant magnitude and variable frequency [START_REF] Vermaelen | Contribution à la modélisation et à la réduction des perturbations conduits dans les systèmes d'entraînement à vitesse variable[END_REF]. Before realizing the measurement, it has been verified that the rotor position has no impact on equivalent induction machine impedance.

Usually, these low voltage tests ( < 1V) are realized with an impedance analyzer [START_REF] Vermaelen | Contribution à la modélisation et à la réduction des perturbations conduits dans les systèmes d'entraînement à vitesse variable[END_REF], [START_REF] Revol | Modélisation et optimisation des performances CEM d'une association variateur de vitessemachine asynchrone[END_REF]. Thus, the impedance variation of our induction machine has been then measured with an impedance analyzer (HP4192A) for a supply voltage of 0.5V. In order to realise tests with higher voltages, a three phase controlled voltage source has also been employed. For V3=0.5V, the impedance evolution is similar whatever the measurement device is: impedance analyzer or controlled voltage source. The impedance norm increases in quasi linear manner. Nevertheless, for the tests at V1=32V and V2=16V, the norm increases in a non linear way with a changing slope. This non linearity has been mentioned in [START_REF] Vermaelen | Contribution à la modélisation et à la réduction des perturbations conduits dans les systèmes d'entraînement à vitesse variable[END_REF], but it has not been discussed because the studied tests were only made at low voltages with an impedance analyzer.

Before explaining the phenomenon, it is important to note that the observed non linearity is difficult to analyze when the magnetic hysteresis phenomenon is taken into account [6]. The measured impedance value would also depend on the machine initial state, so it would depend on its last magnetization point. In order to minimize the influence of this phenomenon, we have supplied the machine before each test by a sine voltage at 50Hz before decreasing progressively the current until the retentive effect becomes very low. Thus the hysteresis will be neglected in the following analysis.

Considering the evolution of the magnetic permeability of various steels used in the electrical machines (Fig. 3), we observe that the permeability increases first in a non linear way, reaches a maximum localised in an area usually considered linear and then decreases asymptoticly as when the machine saturates.

The no load operating point of the induction machine is localized at the saturation elbow level (around the point M1 corresponding to H=300A/m and µr=3400) (Fig. 3) which corresponds to a voltage/frequency ratio of 230/50 (V/Hz).

The blocked rotor tests for all the voltages are done with a lower (V/f) ratio than 230/50. Indeed the maximal V/ f values are equal to 32/40, 16/40 and 0.5/40. Thus, the corresponding magnetic flux is low. The measured points at 40 Hz are localized in an area of low field between points M3 and Mn (Fig. 3).When the frequency increases the magnetic state moves towards Mn. Thus, the non linearity which appears in The non linear evolution of the permeability influences necessarily the machine equivalent impedance. We have seen that the realized tests correspond to the first part of permeability curve (M3  M4  Mn). For high frequencies, the permeability becomes very low (point Mn). During the measurements, the machine impedance increases with the frequency and consequently the current and the magnetic field decrease.

In the area Hn<H<H3, the permeability decreases, the magnetic leakage increases and :

-The dispersion coefficient

        - = σ 2 1 2 L L M 1
increases. M is the stator rotor mutual inductance and Li, is the cyclic inductance (i=1 for stator and i=2 for rotor).

-The leakage inductance During the test with low voltage V3 = 0.5V, the magnetic field is very low for all the frequency values and consequently the permeability remains practically constant. Thus, the field has a weak influence on the impedance variation according the frequency. However, this low voltage corresponds to a magnetic state far from the nominal operating point of the induction machine. Concerning the tests at V1 = 32V and V2 = 16V, the permeability varies and consequently the parameter identification can not be representative of a single state of the induction machine.

Standstill Frequency Response test (V/f = cst)

From the analysis of the previous measurements, it seems necessary to keep the same magnetic state in all the frequency bandwidth in order to identify the machine frequential behaviour without introducing the supplementary variations relied to the material characteristics.

The SSFR method (Standstill Frequency Response) [6] is used to determine the frequential characteristics of the blocked rotor machine. The stator is supplied by a sinusoidal voltage source with variable amplitude and frequency (Fig 6). The ratio V/f is kept constant when the frequency varies. Thus the current amplitude is nearly steady. The method is normalized for the synchronous machine identification and adapted for the induction machine after some simplifications [START_REF] Willis | Derivation of induction motor models from standstill frequency response tests[END_REF][21] [START_REF] Retière | Etude des défauts dans les associations onduleur -machine asynchrone. Exemple d'une chaîne de traction[END_REF]. Indeed, the induction machine rotor is symmetrical, so no rotor positioning is required and it is not necessary to separate the measurement of operational direct and quadrature impedances. Moreover, since the induction machine does not have excitation winding, there are no measurements in the rotor.

The three-phase source used previously is well adapted to the SSFR because it can deliver low voltages at low frequencies and relatively high voltages at high frequencies. This controlled voltage source can then imposes a constant V/f ratio. If the stator resistance is neglected, the magnetic flux is considered as steady in the machine.

The magnetic permeability is then independent on the frequency (Fig 7).

Since the choice of the ratio (V/f) can have an influence on the machine impedance, tests have been realized for various values comprised between 50/1000 and 100/500 [V/Hz]. Compared to the nominal point value 230/50 these ratios correspond to operating points localized between the points M3 and Mn (Fig. 3). The evolution of the norm and the argument of the phase-neutral impedance depending on the frequency are presented in corresponds to a f 1 behaviour, according to the skin effect in a massive conductor [START_REF] Belhaj | Modélisation fréquentielle de la machine asynchrone en vue de l'analyse des perturbations conduites basses fréquences[END_REF].

The usual V/f ratios employed in the SSFR method tests are low. This induces a weak flux compared to the nominal one. In order to increase the flux, the V/f ratio must be increased. But nominal ratio is not allowed when the machine has a blocked rotor. Indeed, at V/f = 30/40 (nominal ratio 230/50) the machine already absorbs its nominal current of 10A. When the frequency varies from 50Hz to 1 kHz, the voltage source must supply a higher voltage varying between 30V and 750V.

Knowing that the V/f variations have an important impact on the operational impedance variation, this kind of measurement remains insufficient to represent the nominal behaviour of the machine.

SSFR test with DC current

In order to approach the machine nominal flux, the magnetic level must be increased. In this aim, a current is constituted of a sinusoidal component iac superposed to a continuous component Idc [START_REF] Canat | Fractional order: frequential parametric identification of the skin effect in the rotor bar of squirrel cage induction machine[END_REF], [START_REF] Dedene | Estimation of a Global Synchronous Machine Model Using a Multiple-Input Multiple-Output Estimator[END_REF], [START_REF] Sudhoff | Experimental characterization procedure for use with an advanced induction machine model[END_REF] can be injected between two of the stator phases, according to the diagram of Fig 6 . A voltage source with regulated current is required, in order to maintain constant the supply current and consequently the magnetic state of the machine, when the frequency increases.

The experiment has been realized for three continuous components Idc = 1A, Idc = 3A and Idc = 5A with the same alternative current component of 350mA. These three tests correspond to three magnetic states of the machine.

The tests have been achieved for three DC current values in order to show the influence of the magnetic state on the evolution of impedance. The maximal value of 5A is because of the supply which can't provide more. It would be interesting to make the test with 10A DC current which is the machine nominal current with an adapted supply. with the frequency whatever the magnetic level is. In addition, there is an important variation of the amplitude depending on the magnetic state. Indeed, for a given frequency, the impedance decreases when the magnetic level increases. This kind of test allows increasing the level of the magnetic flux and at the same time it avoids the problems relied to the constant V/f ratio in SSFR test.

In Fig 9, the machine equivalent inductance decreases according to the frequency. This variation agrees with the SSFR test and shows that for fixed Idc the tests are free from the permeability effect.

The induction machine models frequential identification

The analysis of the various identification measurements showed the effect of the machine magnetic state which has an important influence on the operational impedance value and consequently on the identified model parameters. In what follows, the simulation performances of the harmonic model are compared to the experimental measurements.

The harmonic model is composed of two distinct models. The first one is classic and represents the induction machine behaviour for the fundamental frequency. Its parameters are identified using to the usual tests (TABLE I). The second model represents the frequency influence of the induction machine (Fig 1). The parameters of TABLE II are obtained with SSFR test with DC current for three magnetization currents.

The magnetization current increasing from 1A to 5A induces the resistances R1' et R2' diminution (respectively 31% and 59%) and also a diminution of inductance L2' (53%). The strong diminution of inductance L1' (79%)

and the strong increasing of resistance R3' (150%) lead to the equivalent impedance decreasing. It can be noticed that at Idc=5A, the parameters R1' and L1' of the harmonic model are near RS et Nr' of the usual model identified at 50Hz. shows that the model reproduces correctly the frequential behaviour of the machine with blocked rotor for the three magnetic states identified by the three DC currents. However regarding the impedance phase the test with the highest magnetic state (5A) leads to parameters allowing a better representation of the machine.

Simulation

The modelling of the induction machine associated to its identification procedure is often realised in the aim of numerical simulation. The circuit model of the induction machine has been developed in a simulation software (PSIM) in order to evaluate the influence of the frequential identification on the simulated line currents. The real line currents measured on the benchmark are used as reference for the comparison with the simulated results.

Firstly, the measured line current drawn by the induction machine and the corresponding voltage delivered by its PWM speed drive are recorded. In this way, the real recorded voltages supply the induction machine model during the simulations. This operation allows to simulate the real operating conditions of the machine and to release some eventual differences owed to the variable speed drive. Consequently, the simulation can not take into account the possible interactions between the machine and its voltage inverter.

In what follow, the time evolutions and the frequential analysis of the simulated line currents are compared to the measurements. The speed drive operating point is Veff = 230V, f = 50Hz and the line current I = 95%In of the nominal current. In this figure, when the magnetization level increases, the harmonics amplitudes also increase, approaching the real levels. But, the higher magnetization level used (5A) remains insufficient to approach the real harmonics currents.

The differences between simulation and measurement can also be explained by the space harmonic generated by the induction machine (also machine design, winding, slots,…), when it rotates. These space harmonic are not taken into account in the identification because the tests are realized with a blocked rotor.

In Fig 13, line currents simulated by the harmonic model with different method of identification is presented. If the comparison is limited to the fundamental, the line current waveform obtained thanks to all the identification tests agree with measurements. Concerning the harmonics, among all the identification measurements, the SSFR with Idc=5A leads to the best simulation result.

Conclusion

In this article, three identification tests have been compared for the induction machine frequential modelling in order to reproduce the harmonic perturbations in the line currents. After presenting some induction machine models coming from the literature, we used a model called "harmonic" for comparing various methods of identification tests. This model is based on the superposition of two representations: the first one for the fundamental and the second one for all the other frequential components. In order to identify the model parameters, we were various kinds of identification tests have been realized on the experimental benchmark and then analyzed and discussed. The study showed the importance of keeping a constant flux during all the identification test in order to avoid the non linearity due to the variation of the magnetic permeability. Besides, we highlighted the influence of the magnetic flux level on the frequential response of the machine during tests identification.

The comparison of the measurements put in evidence the test which characterizes the best the machine. It requires the injection of sinusoidal current superposed to a continuous component in order to increase the magnetic state of the machine and to approach the nominal operating conditions.

Nevertheless, some differences exist between simulation and measurements and the accuracy of the harmonic model could be improved by taking into account the space harmonics of the induction machine. Another improving way would be building no linear model allowing the parameters evolution according to the magnetic state.

List of figures

  Finally, three tests have been realized for three voltages (V1=32V, V2=16V, V3=0.5V) and for a frequency range of [40Hz -1kHz]. The variations of the neutral to phase impedance are presented in Fig 2.

  Fig 2 can be explained by the first material magnetization characteristic.

  and consequently the imaginary part of the machine equivalent impedance increases according to the frequency until a value of the magnetic excitation (H=Hn) for which the permeability variation becomes small. This phenomenon can explain the non linearity observed between 100 and 400Hz on the neutralphase impedance norm (Fig2). In addition, when the frequency increases, the skin effect in the magnetic steel leads to the machine equivalent inductance decreasing. This phenomenon can be seen in Fig4where the equivalent inductance evolution shows a diminution from 400 Hz for the test at V1=32V and from 200Hz for V2=16V.

  Fig 5 for four V/f values. It appears that the non linearity observed during the identification tests with constant voltage (Fig 2) is not present in the SSFR measurements. The machine equivalent inductance depending on the frequency (Fig 7)

Fig 8

 8 Fig 8 represents the variations of the norm and the argument of the neutral-phase impedance depending on the frequency for these three magnetic states. It can be noticed that the measured equivalent impedance increases

Fig 10 represents

 10 Fig 10 represents respectively the comparison between the measured and calculated impedances. This figure

Fig 11

 11 Fig 11 represents the line currents simulated by the harmonic model with the parameters identified for the three magnetization levels. The time waveforms of the line currents coming from harmonic model are almost the same as the measured ones for all the DC current. Nevertheless, the current ripples are best described with the machine model determined with higher DC current. The proposed model structure (superposition principle) separates harmonics and fundamental representation. The fundamental current is obtained from the usual model which is the same for the three magnetization levels. The harmonic model impedance corresponding to each magnetization level does not influence the fundamental current amplitude. It operates only to characterize the harmonics. This good result on the fundamental component is confirmed by the spectral analysis of Fig 12. In

Fig 1 :

 1 Fig 1: Induction machine harmonic model.Fig 2 : Phase to neutral impedance norm and argument versus frequency at constant voltage. Fig 3 : Evolution of the magnetic induction and the permeability according to the magnetic field (sheet FeV40050HA). Fig 4 : Evolution of the equivalent inductance of the IM versus frequency (V = cst).Fig 5 : Phase to neutral impedance norm and argument versus frequency at constant V/f ratio. Fig 6 : Measurements device Fig 7 : Evolution of the equivalent inductance of the IM versus frequency (V/f = cst).Fig 8 : Phase to neutral impedance norm and argument versus frequency at constant DC current. Fig 9 : Evolution of the equivalent inductance of the IM versus frequency (Idc = cst).Fig 10 : Phase to neutral impedance norm and argument. Tests (iac+Idc), harmonic model. Fig 11 : Line currents simulated by the harmonic model and the measurements (Idc=cst) Fig 12 : The harmonics of the line currents simulated by the harmonic model and the measurements (Idc=cst). Fig 13 : Line currents simulated by the harmonic model and the measurements Fig 14 : The harmonics of the line currents simulated by the harmonic model and the measurements.

Fig 2 :

 2 Fig 1: Induction machine harmonic model.Fig 2 : Phase to neutral impedance norm and argument versus frequency at constant voltage. Fig 3 : Evolution of the magnetic induction and the permeability according to the magnetic field (sheet FeV40050HA). Fig 4 : Evolution of the equivalent inductance of the IM versus frequency (V = cst).Fig 5 : Phase to neutral impedance norm and argument versus frequency at constant V/f ratio. Fig 6 : Measurements device Fig 7 : Evolution of the equivalent inductance of the IM versus frequency (V/f = cst).Fig 8 : Phase to neutral impedance norm and argument versus frequency at constant DC current. Fig 9 : Evolution of the equivalent inductance of the IM versus frequency (Idc = cst).Fig 10 : Phase to neutral impedance norm and argument. Tests (iac+Idc), harmonic model. Fig 11 : Line currents simulated by the harmonic model and the measurements (Idc=cst) Fig 12 : The harmonics of the line currents simulated by the harmonic model and the measurements (Idc=cst). Fig 13 : Line currents simulated by the harmonic model and the measurements Fig 14 : The harmonics of the line currents simulated by the harmonic model and the measurements.

Fig 5 :

 5 Fig 1: Induction machine harmonic model.Fig 2 : Phase to neutral impedance norm and argument versus frequency at constant voltage. Fig 3 : Evolution of the magnetic induction and the permeability according to the magnetic field (sheet FeV40050HA). Fig 4 : Evolution of the equivalent inductance of the IM versus frequency (V = cst).Fig 5 : Phase to neutral impedance norm and argument versus frequency at constant V/f ratio. Fig 6 : Measurements device Fig 7 : Evolution of the equivalent inductance of the IM versus frequency (V/f = cst).Fig 8 : Phase to neutral impedance norm and argument versus frequency at constant DC current. Fig 9 : Evolution of the equivalent inductance of the IM versus frequency (Idc = cst).Fig 10 : Phase to neutral impedance norm and argument. Tests (iac+Idc), harmonic model. Fig 11 : Line currents simulated by the harmonic model and the measurements (Idc=cst) Fig 12 : The harmonics of the line currents simulated by the harmonic model and the measurements (Idc=cst). Fig 13 : Line currents simulated by the harmonic model and the measurements Fig 14 : The harmonics of the line currents simulated by the harmonic model and the measurements.

Fig 8 :

 8 Fig 1: Induction machine harmonic model.Fig 2 : Phase to neutral impedance norm and argument versus frequency at constant voltage. Fig 3 : Evolution of the magnetic induction and the permeability according to the magnetic field (sheet FeV40050HA). Fig 4 : Evolution of the equivalent inductance of the IM versus frequency (V = cst).Fig 5 : Phase to neutral impedance norm and argument versus frequency at constant V/f ratio. Fig 6 : Measurements device Fig 7 : Evolution of the equivalent inductance of the IM versus frequency (V/f = cst).Fig 8 : Phase to neutral impedance norm and argument versus frequency at constant DC current. Fig 9 : Evolution of the equivalent inductance of the IM versus frequency (Idc = cst).Fig 10 : Phase to neutral impedance norm and argument. Tests (iac+Idc), harmonic model. Fig 11 : Line currents simulated by the harmonic model and the measurements (Idc=cst) Fig 12 : The harmonics of the line currents simulated by the harmonic model and the measurements (Idc=cst). Fig 13 : Line currents simulated by the harmonic model and the measurements Fig 14 : The harmonics of the line currents simulated by the harmonic model and the measurements.

Fig 10 :

 10 Fig 1: Induction machine harmonic model.Fig 2 : Phase to neutral impedance norm and argument versus frequency at constant voltage. Fig 3 : Evolution of the magnetic induction and the permeability according to the magnetic field (sheet FeV40050HA). Fig 4 : Evolution of the equivalent inductance of the IM versus frequency (V = cst).Fig 5 : Phase to neutral impedance norm and argument versus frequency at constant V/f ratio. Fig 6 : Measurements device Fig 7 : Evolution of the equivalent inductance of the IM versus frequency (V/f = cst).Fig 8 : Phase to neutral impedance norm and argument versus frequency at constant DC current. Fig 9 : Evolution of the equivalent inductance of the IM versus frequency (Idc = cst).Fig 10 : Phase to neutral impedance norm and argument. Tests (iac+Idc), harmonic model. Fig 11 : Line currents simulated by the harmonic model and the measurements (Idc=cst) Fig 12 : The harmonics of the line currents simulated by the harmonic model and the measurements (Idc=cst). Fig 13 : Line currents simulated by the harmonic model and the measurements Fig 14 : The harmonics of the line currents simulated by the harmonic model and the measurements.

List of tables

'


Santos Sousa

ATMOSPHERIC PRESSURE GENERATION OF O 2 (a 1 ∆ ∆ ∆ ∆ g ) BY MICROPLASMAS

Keywords: numbers: 52.50.Dg, 52.70.Kz, 52.80.Tn

The generation of singlet oxygen states (O 2 (a 1 ∆ g )) by microplasmas has been studied experimentally. In the present paper, it is shown that micro-cathode sustained discharges (MCSD's) can be used to produce high fluxes of O 2 (a 1 ∆ g ) at atmospheric pressure. In He/O 2 /NO mixtures, O 2 (a 1 ∆ g ) number densities higher than 10 16 cm -3 can be generated by this 3-electrode configuration and transported over distances of some tens of cm. In fact, at total flow rates up to 30 ln/min, O 2 (a 1 ∆ g ) fluxes above 10 mmol/h were measured in the MCSD afterglow, at 26 cm downstream. As a result, MCSD's appear to be very efficient and suitable tools for the continuous production of large amounts of O 2 (a 1 ∆ g ) at atmospheric pressure, which could give rise to a wide range of new applications, namely biological. The effect of different parameters such as gas flows and mixtures, and discharge current are discussed in the paper.

INTRODUCTION

Singlet delta oxygen, O 2 (a 1 ∆ g ), the first electronically excited low-energy (0.98 eV) state of molecular oxygen, possesses an extremely long radiative lifetime of more than 73 min. Moreover, this metastable state of molecular oxygen is very resistant against homogeneous and heterogeneous relaxation processes, which makes it possible for O 2 (a 1 ∆ g ) to be used far away from its place of production. As so, this unique molecular state can be used as an energy donor for a large spectrum of applications, including removal of air pollutants [1], degradation of synthetic polymers [2], medicine [3], laser excitation [4], or combustion [5]. Furthermore, O 2 (a 1 ∆ g ) is also well known to produce cytotoxic effects and to be an important agent in biophysical and biochemical processes [START_REF] Krinsky | Singlet Oxygen[END_REF][START_REF]Singlet O[END_REF][START_REF]singlet oxygen, UV A and ozone[END_REF][9][10][11]. Up to now, for biological applications, O 2 (a 1 ∆ g ) is mainly produced through the photosensitization technique [11] resulting in rather low O 2 (a 1 ∆ g ) concentrations, of a few hundreds parts per billion, in gas flow at atmospheric pressure [12]. Thus, new plasma sources of high fluxes of O 2 (a 1 ∆ g ) molecules are of high potential interest for biological applications, but at the mandatory condition that these sources can be operated at atmospheric pressure and the O 2 (a 1 ∆ g ) can be transported to targets placed at some tens of cm from the source.

While it is well known for many years that O 2 (a 1 ∆ g ) can be efficiently produced in oxygen discharges operating at a pressure of few mbars [13], its production by high pressure discharges, and specifically at atmospheric pressure, is still very challenging, as described in the comprehensive topical review about the physics and the engineering of O 2 (a 1 ∆ g ) production in low temperature plasma, recently published by Ionin et al [14]. In the previous years, the production of O 2 (a 1 ∆ g ) by high pressure electrical discharges has attracted much attention because of its potential application for the pumping a Electronic address: joao.santos-sousa@u-psud.fr of the oxygen-iodine laser [15][16][17][18][19]. In order to efficiently produce the high O 2 (a 1 ∆ g ) densities needed for these applications, it is now established that an electric discharge source must satisfy two conditions: i) it must be stable at high pressure and high power loading and ii) it should operate at low reduced electric field [START_REF] Hill | Proceedings of the International Conference on LASERS[END_REF]. So far, high-pressure, non-self-sustained discharges operating in short duration pulsed mode seemed to be the most attractive option for generating stable discharges at high pressure. Different kinds of non-self sustained discharges have been extensively investigated for O 2 (a 1 ∆ g ) production: i) the e-beam sustained discharges [14,17,[START_REF] Vagin | [END_REF], ii) the crossed spiker-sustainer discharges [22,23], and iii) the controlled avalanche method [24]. However, we have recently demonstrated [25] that the so-called micro-cathode sustained discharge (MCSD), operating in DC mode, can also be used to produce large amounts of O 2 (a 1 ∆ g ). Detailed numerical simulations of the MCSD [START_REF] Bauville | Proceedings of the 38th AIAA Plasmadynamics and Lasers Conference[END_REF][START_REF] Makasheva | [END_REF] have shown that the MCSD is basically a high pressure positive column with a low value of the reduced electric field and gas temperature. It must be pointed out that, to this date, the MCSD is the only DC discharge technique which has been proven capable of generating high O 2 (a 1 ∆ g ) concentrations at atmospheric pressure [START_REF] Puech | Proceedings of the SPIE International High Power Laser Ablation Symposium[END_REF][START_REF] Sousa | [END_REF]. As shown by modeling and experiments [30][31][32], the MCSD exhibits unrivalled discharge stability properties which directly result from the stability of the micro-hollow cathode discharge (MHCD) which acts as a plasma cathode in such a device [33,34]. The stability of the MHCD results from the transition from an abnormal discharge mode, which occurs at low current when the discharge is confined inside the hole of the MHCD, to a normal discharge mode, characterized by a plasma extension on the backside of the MHCD cathode, as the current increases [35][36][37]. As a result, this transition prevents, at high pressure and current, the development of the classical thermal instability which induces the discharge constriction and the glow to arc transition in many kinds of discharges.

EXPERIMENTAL SET-UP

As shown in Figure 1, our microplasma reactor basically consists of a micro-hollow cathode discharge (MHCD), a discharge concept first developed by Schoenbach and co-workers [38], acting as a plasma cathode to generate a stable glow discharge of larger volume between the MHCD and a third, 2.5 cm diameter, planar electrode, placed at a distance of 8 mm from the MHCD. This 3-electrode configuration is the so-called MCSD, which was initially proposed by Stark and Schoenbach [33,34]. The MHCD itself was made of 100 µm thick molybdenum electrodes glued on each side of a 500 µm thick alumina plate. An 800 µm diameter hole was pierced through this sandwich. The MHCD cathode is connected to a negative power supply through a 440 kΩ resistor, while the MHCD anode (anode A1) is directly grounded. Facing the anode A1 is the third electrode (anode A2), which is positively biased through a 440 kΩ resistor. The device is placed into a stainless steel chamber, which is evacuated down to 10 -3 Torr before the gas mixture is introduced. The total pressure in the chamber is measured with a Baratron gauge, while the helium, oxygen and nitric oxide (NO) flows are separately regulated with flow-meters. Optical windows (not shown in Figure 1) allow CCD imaging of the spatial discharge development and optical emission spectroscopy.

After passing through the discharge chamber, the gas flow is evacuated through a pumping line. A calibrated detection cell equipped with quartz windows can be positioned at different locations on the pumping line, allowing the measurement of the radiative emission of O 2 (a 1 ∆ g ) at 1.27 µm. This emission provides a measure of the O 2 (a 1 ∆ g ) number density in the flowing afterglow. The distance between this cell and the discharge reactor can be varied in the range 20-70 cm. The detection system consists of a calibrated liquid-nitrogen-cooled InGaAs detector (Judson model J22D-M204) and a narrow band (19 nm) interference filter centred at 1.27 µm. The signal is amplified by an op-amp with a gain of 10 7 and monitored with a numerical oscilloscope, which is also used to record the electrical parameters of the discharge. A part of the infrared emission is picked up with an optical fiber connected to a spectrometer equipped with a 512 pixel InGaAs detector (OMA-V from Roper Scientific). Figure 2 presents a comparison between spectra recorded at 26 cm downstream in the afterglow with the MCSD operating in pure oxygen at a pressure of 9 mbar and a discharge current of 2 mA (Figure 2 (a)) and in a He/O 2 /NO mixture at atmospheric pressure (2.2% of O 2 and 185 ppm of NO) with a discharge current of 5 mA (Figure 2 (b)). Even if the O 2 (a 1 ∆ g ) number density is 25 times higher in the case of the He/O 2 /NO mixture at atmospheric pressure (9.8 10 15 cm -3 ) than in the case of the low pressure of pure oxygen (3.8 10 14 cm -3 ), the comparison of the normalized spectra points out that they are essentially the same, in spite of the very different discharge conditions. This insures that, in He/O 2 /NO mixtures, the infrared signal measured in the flowing afterglow is only due to the radiative emission of O 2 (a 1 ∆ g ) and is not contaminated by spurious signals originating either from important collision-induced emission or NO* 2 de-excitation, which could add continua to the O 2 (a 1 ∆ g ) spectrum. Therefore, the signal measured by the Judson detector can be used with confidence to measure the O 2 (a 1 ∆ g ) density. Knowing the sensitivity of the detector, the transmission curve of the interference filter, the geometrical parameter of the optical detection system, and the radiative transition probability of O 2 (a 1 ∆ g ), (A=2.256 10 -4 s -1 ) [39], the absolute O 2 (a 1 ∆ g ) number density can be deduced from the amplitude of the infrared signal. In that procedure, a Monte Carlo simulation was used to calculate the probability that a photon emitted in the detection cell will reach the InGaAs element of the Judson detector. As a result, in our device, the O 2 (a 1 ∆ g ) number density is correlated to the intensity of the Judson signal by the ratio 3.77 10 15 O 2 (a 1 ∆ g ) cm -3 / mV.

The measurement of the ozone (O 3 ) number density is based on ultraviolet (UV) absorption. As a UV source we used a deuterium lamp emitting a broadband spectrum between 180 and 400 nm. In the present work, the measurements were mainly performed by using the calibrated detection cell, 8 cm long, as an absorption cell. However, in some experiments a longer cell, 50 cm long, was used to get more accurate results at low O 3 density. The O 3 number density was deduced from the ratio of the intensity (I) transmitted when the discharge is ON to the intensity (I 0 ) transmitted when the discharge is OFF. The O 3 density is deduced from the relation

ln(I 0 /I) = [O 3 ] . σ . L (1),
where σ is the absorption cross-section and L is the length of the absorption path. The O 3 absorption cross-section was measured by Molina and Molina [40]. After passing through the absorption cell, the transmitted light intensity is detected with a 75 cm focal length spectrometer (Acton 2750) and a 1340 pixel detector (Pixis from Roper Scientific) providing a resolution of 0.044 nm per pixel in the UV range. Figure 3 shows the variation of ln(I 0 /I) versus wavelength in the range 220-270 nm, for a discharge operating in a He/O 2 mixture (1% of O 2 ) at atmospheric pressure with a current of 3 mA. From the absorption cross-section given in [40], an O 3 density of 2.15 ± 0.06 10 15 cm -3 was obtained over all the wavelength range. For the different experimental conditions of the present work, we measured intensity values integrated between 253 and 255 nm, and we recorded the temporal evolution of these values with the discharge alternatively ON and OFF, as shown in Figure 4. As a result, O 3 number densities in the range 10 13 -10 16 cm -3 were measured.

The O 2 (a 1 ∆ g ) and O 3 number densities presented in this paper were measured in the MCSD afterglow, at 26 cm downstream.

RESULTS AND DISCUSSION

The O 2 (a 1 ∆ g ) number density is a very sensitive function of the gas velocity, due to the occurrence of fast quenching processes. In fact, the O 2 (a 1 ∆ g ) molecules are presumably quenched by atomic oxygen (O-atom) and O 3 for long residence time [14,41]. Therefore, the higher the gas flow, the lower the destruction of the O 2 (a 1 ∆ g ) states before their detection in the afterglow, as pointed out by Figure 5.

Actually, for binary He/O 2 mixtures at atmospheric pressure, the MCSD allowed the production of an O 2 (a 1 ∆ g ) density of about 10 15 cm -3 whenever the helium flow was higher than 15000 sccm. In order to get greater O 2 (a 1 ∆ g ) concentrations, we could increase the oxygen flow. However, this can't be done endlessly due to insufficient energy deposition per oxygen molecule, P/Q, for higher oxygen flows. Thus, as shown in Figure 5, besides the helium flow, the O 2 (a 1 ∆ g ) generation is maximized whenever the oxygen concentration is about 1-2%.

In these experimental conditions, characterized by atmospheric pressure and low oxygen partial pressure, the main O 2 (a 1 ∆ g ) destruction channel is expected to be the three-body reaction:

O 2 (a 1 ∆ g ) + O + He O 2 (X 3 Σ g ) + O + He
(2) [41] Thereby, one solution to increase the O 2 (a 1 ∆ g ) number density is to add, in the He/O 2 mixture, an Oatom scavenger in order to reduce the quenching processes. As so, we have studied the influence of adding NO molecules at low concentration in the He/O 2 mixture at atmospheric pressure. As pointed out by Figures 6 and7, adding small concentrations of NO induces, on the one hand, a large increase of the O 2 (a 1 ∆ g ) number density (up to 50 times more), and, on the other hand, a huge reduction on the O 3 number density (up to its apparent complete destruction). As shown in Figure 8, the O 3 molecules and the O-atoms are quenched by NO through the reactions: [14,41,42] O + NO + He NO 2 + He (4) [START_REF] Hicks | Proceedings of the 38th AIAA Plasmadynamics and Lasers Conference[END_REF][START_REF] Benavides | Proceedings of the 38th AIAA Plasmadynamics and Lasers Conference[END_REF] So, the observed increase in the O 2 (a 1 ∆ g ) density probably results from a decrease in the O-atom and O 3 densities. In fact, the O 2 (a 1 ∆ g ) states are expected to be quenched by the O-atoms in the discharge and by the O 3 molecules in and after the discharge. Since the effect of NO is to quench the O-atom and the O 3 , injecting NO in the gas flow is likely to reduce the loss of O 2 (a 1 ∆ g ) molecules. However, this assumption remains to be validated by measurements of the O-atom and NO x densities, in the discharge chamber and in the afterglow, and for different discharge conditions. We intend to make these measurements by using, for the O-atom density, either the NO titration method or the two photon laser-induced fluorescence technique, as we have already done for MCSD operating in pure oxygen [START_REF] Bauville | Proceedings of the 38th AIAA Plasmadynamics and Lasers Conference[END_REF], and, for the NO x density, through the UV absorption approach or the Fourier transform infrared spectroscopy technique.

O 3 + NO NO 2 + O 2 (3)
In Figures 6 and7, we observe that the huge reduction on the O 3 number density is much greater than the small concentration of NO added. Indeed, as pointed out by Figure 7 (a), by only introducing 0.15 sccm of NO in the gas mixture, the O 3 is presumably totally destroyed, with its density passing from 3.0 10 15 cm -3 to less than 1.0 10 13 cm -3 (under the limit of detection). As previously mentioned, besides destroying O 3 , the NO also quenches O-atom. With increasing pressure, O 3 starts to be produced effectively in a three-body recombination of atomic and molecular oxygen (see Figure 8):

O + O 2 + He O 3 + He (5) [41] This is especially true at atmospheric pressure, when almost all O-atom becomes O 3 [START_REF] Takezawa | Proceedings of the 40th IEEE -Industry Applications Society Annual Meeting[END_REF][START_REF] Eliasson | [END_REF]. So, a reduction in the O-atom density also leads to a decrease in the O 3 density in the gas flow. Furthermore, as NO is not lost in the process, but it is rather recycled in collisions (see Figure 8) [42] the earlier referred effects and reactions are somehow catalytic and, thus, cause a huge decrease on the O 3 density.

O + NO 2 NO + O 2 (6)
On the other hand, in Figures 6 (a The further increase of the O 2 (a 1 ∆ g ) density can only be explained by a decrease on the O-atom density and a consequent reduction of the O 2 (a 1 ∆ g ) molecules loss. The stabilization of the O 2 (a 1 ∆ g ) density for higher NO flow could mean that there are no more O-atom to be quenched. But, as already mentioned, these hypotheses need to be verified by measuring the O-atom density in and after the discharge. Besides that, NO and NO 2 molecules also having a quenching effect on O 2 (a 1 ∆ g ) can induce as well a stabilization of the O 2 (a 1 ∆ g ) density. Moreover, a decrease of the O 2 (a 1 ∆ g ) density is expected for greater concentrations of NO. Indeed, a reduction of the number of O 2 (a 1 ∆ g ) molecules with increasing flow of NO has been observed for other discharge conditions [START_REF] Sousa | [END_REF]. Therefore, the optimum NO gas flow is a sensitive function of both the oxygen partial pressure and the helium flow.

From Figures 6 and7, we can also infer about the influence of the helium flow and the oxygen partial pressure on the O 2 (a 1 ∆ g ) and O 3 densities. For a constant oxygen partial pressure, the O 2 (a 1 ∆ g ) initial density (for Q(NO) = 0 sccm) is larger when the helium flow is higher (while no difference is observed for O 3 number density). But the increase, versus the NO concentration, of the O 2 (a 1 ∆ g ) density is smaller as the helium flow is higher. A greater gas flow means a shorter transit time and, thus, a reduction of the quenching effect before reaching the detection cell. This immediately explains the increase observed for the O 2 (a 1 ∆ g ) initial densities. As the O-atom and the O 3 are also less quenched by NO for higher gas flow, the addition of small concentrations of NO becomes less effective in increasing the O 2 (a 1 ∆ g ) number density. Even if we do not have measurements of the Oatom density, Figures 6 and7 demonstrate that the reduction of the O 3 density with increasing NO flow is less important for higher gas flow. As shown in Figures 6 (b) and 7 (b), for a constant helium flow, besides the decrease of the energy deposited per oxygen molecule, P/Q, we do not notice large differences with increasing oxygen partial pressure, apart from the O 3 initial density. The increase of the O 3 density with higher oxygen concentration is expected as O 3 is effectively produced through the reaction (5). To clearly understand the influence of the helium flow and the oxygen partial pressure on the O 2 (a 1 ∆ g ) density, we should observe Figure 9.

As pointed out by Figure 9, for a discharge current of 3 mA, the largest O 2 (a 1 ∆ g ) densities are obtained for helium flows of about 6000 sccm, regardless of the oxygen partial pressure. The maximal O 2 (a 1 ∆ g ) density, of about 7.5 10 15 cm -3 , was observed for an oxygen partial pressure of 20 mbar. The shape of the curves showed in Figure 9 results from a competition between two effects. Increasing the total flow leads, on the one hand, to a decrease of the residence time in the afterglow, and, on the other hand, to a decrease of the residence time in the discharge and to an increase of the oxygen flow. The first effect causes a reduction of the losses of O 2 (a 1 ∆ g ) during transport. The last two effects induce a decrease of the P/Q, resulting in a reduction of the production of O 2 (a 1 ∆ g ). For a given oxygen partial pressure, i.e. a given oxygen gas flow, the P/Q value can be increased by increasing the discharge current. Figure 10 shows that, for oxygen and NO partial pressures respectively equal to 23 and 0.19 mbar the O 2 (a 1 ∆ g ) number density increase is nearly linear for discharge currents up to 5 mA, while at higher currents the O 2 (a 1 ∆ g ) density begins to saturate. As a result, O 2 (a 1 ∆ g ) number densities higher than 1.0 10 16 cm -3 have been produced by MCSD operating at atmospheric pressure.

CONCLUSIONS

In the present paper, we have shown that the MCSD's can be efficiently used to generate stable DC glow discharges, free from the glow-to-arc transition, at atmospheric pressure in binary He/O 2 and ternary He/O 2 /NO mixtures. Our results show that O 2 (a 1 ∆ g ) number densities higher than 10 16 cm -3 can be obtained at atmospheric pressure and transported over some tens of cm for helium flows in the range 2000-30000 sccm. It should be emphasized that the densities of the various excited states, O 2 (a 1 ∆ g ) and O 3 , impacting targets positioned up to 50 cm downstream are very reproducible. Throughout the course of experiments performed during several months, the variation of the O 2 (a 1 ∆ g ) and O 3 concentrations is lower than 10%, for given values of gas flows and discharge current. Moreover, it should be pointed out that the ratio between the O 2 (a 1 ∆ g ) and the O 3 densities can be finely tuned through the values of the discharge current and the NO concentration. Thus, MCSDs operating at atmospheric pressure in He/O 2 /NO mixtures appear to be very suitable tools for different applications, and particularly for biological ones [START_REF] Sousa | Proceedings of the 61st Annual Gaseous Electronics Conference[END_REF].

  ) and 7 (a), the increase of the O 2 (a 1 ∆ g ) number density is much larger than the initial concentration of O 3 (for Q(NO) = 0 sccm). As clearly pointed out by Figure6(b) and 7 (b), the increase of the O 2 (a 1 ∆ g ) density is strongly correlated with the reduction of the O 3 density. As a matter of fact, the rise of the number of O 2 (a 1 ∆ g ) molecules is equal to the decrease found on the number of O 3 molecules. Actually, this is also observed in Figures6 (a) and 7 (a), as the increase of the O 2 (a 1 ∆ g ) density until the O 3 complete destruction is equal to the O 3 initial concentration.

Figure 1 :

 1 Figure 1: Schematic of the experimental set-up.

Figure 2 :

 2 Figure 2: Infra-red spectra measured in the flowing afterglow (26 cm downstream from the MCSD) in a discharge operating (a) in pure O 2 at a pressure of 9 mbar and a discharge current of 2 mA, or (b) in a He/O 2 /NO mixture at atmospheric pressure (2.2% of O 2 and 185 ppm of NO) and a discharge current of 5 mA.

Figure 3 :

 3 Figure 3: Evolution of the absorbance versus the wavelength for a discharge operating at atmospheric pressure in a He/O 2 mixture (1% of O 2 ) and a discharge current of 3 mA.

Figure 4 :

 4 Figure 4: Temporal evolution of the transmitted intensity of the deuterium lamp integrated from 253 to 255 nm. For times lower than 10 s or greater than 45 s, the discharge is OFF, while it is ON for times between 10 and 45 s. Upper red curve: He/O 2 mixture with 625 ppm of O 2 ; lower blue curve: He/O 2 mixture with 1% of O 2 .

Figure 5 :

 5 Figure5: Evolution of the O 2 (a 1 ∆ g ) number density versus the oxygen flow, obtained at atmospheric pressure, for a discharge current of 3 mA, in He/O 2 mixtures, for different helium flows: full squares=3500 sccm, full stars=7000 sccm, full circles=14000 sccm, full diamonds=21000 sccm, and crosses=28000 sccm.

Figure 6 :

 6 Figure 6: Evolution versus the NO flow of the O 2 (a 1 ∆ g ) (full diamonds) and O 3 (full squares) number densities obtained at atmospheric pressure, for a discharge current of 3 mA, in He/O 2 /NO mixtures, with constant oxygen partial pressure of 10 mbar and helium flows of (a) 8000 sccm and (b) 25000 sccm.

Figure 7 :

 7 Figure 7: Evolution versus the NO flow of the O 2 (a 1 ∆ g ) (full diamonds) and O 3 (full squares) number densities obtained at atmospheric pressure, for a discharge current of 3 mA, in He/O 2 /NO mixtures, with constant oxygen partial pressure of 20 mbar and helium flows of (a) 6250 sccm and (b) 25000 sccm.

Figure 8 :

 8 Figure 8: Schematic of the quenching of O-atom and O 3 by NO.

Figure 9 :

 9 Figure 9: Evolution of the O 2 (a 1 ∆ g ) number density versus the helium flow, obtained at atmospheric pressure, for a discharge current of 3 mA, in He/O 2 /NO mixtures, with constant NO flow of 1.4 sccm and different oxygen partial pressures: full squares=3 mbar, full stars=10 mbar and full circles=20 mbar.

Figure 10 :

 10 Figure 10: Evolution of the O 2 (a 1 ∆ g ) number density versus the discharge current, obtained at atmospheric pressure in a He/O 2 /NO mixture with O 2 and NO partial pressures respectively equal to 23 and 0.19 mbar, and an helium flow of 3500 sccm.

Fig. 10

 10 Fig. 1

ACKNOWLEDGMENTS

This work was partially supported by the French Agence Nationale de la Recherche. J.S.S thanks the Portuguese Fundação para a Ciência e a Tecnologia (Grant No. SFRH/28668/2006) for their financial support.