A.-L Zimmermann
email: anna.zimmermann@ltf.mw.tum.de

R H M Giepman

Q T Nguyen

C Aalburg

V Gümmer

S Y Mpo

O Tatin

Design of a static test rig for advanced seals and air bearing testing

Keywords:

Today's industrial gas turbines are required to cope with strong fluctuations caused by a strongly varying feed of renewable energies into the grid. ese transient operating conditions result in high temperature gradients and consequently lead to increased axial and radial displacements of turbine parts. Such flexible operations need to be supported by novel sealing technologies. is paper presents a new test facility for investigating advanced seals under 2D static conditions. It facilitates detailed experimental studies of the static pressure distribution on the seal air bearing faces, measurements on the leakage flow through the seal and measurements on the air bearing force balance. e clearance between the rotor and the seal can be set very accurately and it is furthermore possible to apply a predetermined amount of eccentricity to the seal / rotor combination.

INTRODUCTION

e increasing share of renewables on the electrical grid makes the supply side of the grid less predictable than it used to be. Large fluctuations may occur, which are to be compensated by the traditional sources of electricity, like gas turbines [ ]. Today's gas turbines are, however, designed for high performance operation in a base-load regime and their design is not well focused on quickly changing load requirements. Flexible operation cycles result in high temperature gradients coupled with large axial and radial displacements of turbine parts and are currently limited by the tight clearance between the rotor and stator. Novel technologies thus need to be introduced to balance demand peaks by providing flexible operation, while still increasing the performance of gas turbines. Current sealing concepts are particularly affected, so that advanced seal design concepts need to be invented, optimized and tested at engine-representative conditions to match the latest requirements on turbomachinery.

Adaptive seals present a promising approach to allow transient operations and reduced leakage flows. ey ensure minimal clearances and can handle a wide range of operating conditions [ ].

e seal is spring-mounted allowing it to follow the rotor's axial movements and small feed holes are present on the seal's surface. ese feed holes inject highpressure air in the rotor / stator gap, effectively creating a hydrostatic gas bearing between both components. Performance of these gas bearings and consequently the design and optimization have been the subject of various studies. Of special interest here are: the stiffness, damping, load carrying capacity, flow rate and stability of the air bearing [ ].

As a necessary first phase in designing an adaptive seal the static characteristics of air bearings are to be understood [ ].

is determines the general feasibility. Subsequently, a rotating test rig is developed [ ], which is outside the scope of this paper, to address the dynamic behaviors of the seal. Nishio et al. [ ] investigated both static and dynamic characteristics of air bearings with feed holes of less than 0.05 mm in diameter and focused on the effect of the surface roughness on the bearing characteristics. Fourka et al. [ ] developed a numerical approach to predict the stability of air bearings, which was supported by some experimental test results. A similar research approach with a similar test facility was developed by Franssen et al. [ ]. All test facilities mentioned allow the investigation of the load capacity, but there has been no study covering the exact pressure distribution on the air bearing surface. Such measurements are, however, of great interest since the pressure in the end translates back to the air bearing's stiffness, which guarantees a non-contact operation of the seal. Also, for the sake of validating and improving the numerical codes [ -] used for simulating the behavior of adaptive seals it would be helpful to have accurate pressure distributions available.

In this paper a new test facility is presented which allows to study advanced seals within a static test environment. e static characterization is of fundamental interest to assure a successful application of the seal. e gained findings can provide a deeper understanding of the air bearing behavior and can be fed into dynamic tests as a next step towards . TEST RIG . Rig capabilities e new test facility was developed to test advanced seal designs under various Reynolds and Mach numbers, representing realistic gas turbine conditions. A 2D straight segment of such an advanced seal or air bearing can be fi ed into the rig while its modular design supports a quick exchange of the interchangeable test build. Consequently, various types of test builds can be tested without great effort and within short periods of time. e rig allows for the horizontal and vertical gap adjustment to simulate different axial and radial clearances, respectively. It is also possible to set a certain eccentricity of the air bearing in the radial direction. Numerical studies of an optimized inlet and outlet design were done to ensure uniform flow conditions across the span of the seal.

e leakage across the seal is one of the key outcomes of the planned measurement campaigns. Special effort was therefore placed on sealing all parasitic leakage paths, in particular: 1) leakage out of the housing and 2) leakage around the test build, instead of through the test build. Additionally, the rig is capable of force measurements to quantify the load acting on the bearing face. Selected measures were planned Pressurized air enters the rig through the air inlet and is guided to the inside of the rig. e extended geometry of the inlet duct hereby ensures a uniform inlet flow across the entire span of the seal. A er the air flow passes the test build it exits the rig through the extended duct and outlet, whereby the test build could either be a seal or a bearing. e test build is interchangeable and always consists of two units: A 2D rotor and a seal model. A shallow pocket with radius r p is embedded around the feed hole exit. Since the most interesting area to be examined is the bearing face, it is schematically illustrated in Fig. b). e air flow comes from the high pressure area P in upstream the seal and is fed through the feed hole into the bearing clearance with the axial gap width x BF . e flow impinges on the opposite bearing surface of the rotor and escapes continuously to the low pressure area P out downstream the seal.

e rotor and the seal are equipped with a total of 66 pressure taps to allow for an accurate reconstruction of the flow field a erwards. Metal tubes with an outer diameter of 1.1 mm or 1.6 mm are on the models and the rig's sidewalls and are connected through plastic pressure tubing. ese are then hooked up to the pressure scanners outside the rig. Furthermore, the rotor features 3 flush-mounted proximity probes, which are used for tracking the seal / rotor clearance across its full-span.

. . Gap adjustment

Axial gap.

e traversing mechanism illustrated in Fig. is used to set the axial gap width at the bearing face between seal and rotor. e rotor model can be mounted on the slider, which is axially guided by two rods. A stepless adjustment of the axial gap width can be realized by using the screw mechanism comprising two counterrotating nuts screws engaged on the threaded rod. e distance between rotor and seal is measured by 3 proximity probes, which have an axial range up to 2 mm, and the two dial gauges outside the rig.

e gauges also help to detect and remove any model tilt. To ensure a fixed position of the rotor during testing, multiple features are present and can be applied: First, by tightening the two counterrotating nuts against each other, the rotor movement can be blocked. Second, two eccentric rollers on both rig sidewalls can be used to additionally push the rotor / slider unit against the bo om plate. ird, a pair of limit stops can be clamped between rotor and the rig's backwall to facilitate the gap adjustment and to block any kind of rotor movement during rig operation. e seal model is a ached firmly to the backwall of the rig and is furthermore supported by 3 vertical rods to minimize any deformation during rig operation (see Fig. ).

e length of the rods can be varied stepwise by adding shims of different thicknesses. is setup allows for a variation of ±0.38 mm in the radial gap size. Additionally, a thin shim can be inserted between the rotor and slider to have the rotor and seal surface tilted with respect to each other.

. . Sealing features

e test rig has been designed to provide accurate mass flow measurements of the leakage flow across the seal / rotor combination. erefore, it is important to remove or properly seal all parasitic leakage flow paths: ) air leaking out of the housing, ) air leaking around the seal / rotor housing.

Well-placed sealing rings are used to block the air flow out of the housing and an ingenious sealing mechanism at both sidewalls, see Fig. , minimizes the air leaking between the test build and the housing walls. Eventually, two rubber shims added to a pressure fi ing can be pushed against the test build to consequently seal all critical areas. e pressure fi ing is hereby linearly guided by a guiding rod and the two . . Rigid design A structural analysis of all components was made to ensure a stiff design to minimize the deformation of the bearing surfaces while testing. e expected deformations at the bearing surfaces in axial direction were calculated to be less than 2.5 µm at maximum operation pressures. ese deformations are substantially smaller than the machining tolerances of 13 µm and can be neglected.

. MEASUREMENT SYSTEM

A comprehensive measurement system to measure mass flow rate, static pressure, clearance and force was built up and allows for a detailed characterization of the air bearing flow. e measured error of the flow meter device is indicated with ±0.5%of full scale. Furthermore, the mass flow measurements are affected by, first, unwanted leakage paths around the seal / rotor combination inside of the test rig and, second, by leakage from the inside to the outside of the test rig. Multiple leak tests were performed to quantify the remaining leakage rate of the rig. During all tests either the rig's outlet or the seal was fully closed and via the inlet the rig was pressurized up to a certain pressure level. e rate of depressurization could subsequently be used to infer the leakage flow: e leakage inside the rig around the seal / rotor combination is calculated to be less than 0.2%, and less than 0.02 % from the inside to the outside of the rig. Tab. sums up all proportions of uncertainty of the mass flow measurement.

Table . Uncertainty of mass flow measurements

S U [%]

Flow meter accuracy ±0.5 In-rig leakage (around seal / rotor) 0.2 In-to-out rig leakage 0.02

Overall uncertainty -0.5 ... 0.72

Repeatability of mass flow measurements. A number of operation conditions were repeated multiple times and evaluated with regard to the mass flow repeatability. At mass flow rates of ∼ 0.045 kg/s the repeatability error is ±1.6%. is value decreases for mass flow rates smaller than 0.045 kg/s.

. Static pressure

Before starting the experimental tests, a pre-study based on CFD simulations was done to estimate the distribution of static pressures at the rotor's bearing surface. e feed hole flow impinges at the center of the segment, more specifically the origin of the coordinate system and from there on spreads out into both y-and z-directions.

Due to a symmetrical spread of the flow it is sufficient to only investigate the flow within one quarter of the segment in detail, see Fig. b).

e bearing face flow can basically be divided into two flow regimes: A radial flow expands in z-direction from the high pressure (∼ P in ) at the feed hole region to the low pressure (∼ P out ) at the upper or bo om edge of the bearing face; due to the presence of an adjacent feed hole outside the segment, a circumferential flow expands from the high pressure at the feed hole to a stagnation pressure in between the feed holes. Consequently, the flow in the la er regime first follows the y-direction but is then deflected into the z-direction and eventually merges with the flow in the radial regime.

e CFD results were used to optimize the distribution of the pressure taps across the bearing surface such that all important flow features can be captured: the feed hole, expansion and stagnation region as well as the area away from the feed hole. Since the numerical result predicted 4 nearly identical quarters around the feed hole, the authors decided to investigate only one quarter of the segment in detail. e black markers in Fig. a) and b) illustrate the final pressure tap pa ern on the rotor's bearing surface. As it can be seen, the first quarter possesses the largest amount of pressure taps, while some additional taps are located in the other quarters to confirm the symmetry assumption. More pressure taps were also fi ed on the seal to measure pressures upstream and downstream of the bearing face and to check for flow uniformity.

e pressure measurements were performed with multiple Scanivalve pressure scanners of type DSA3218 in different operating ranges.

Uncertainty of pressure measurements. First, the vendor given accuracy of the pressure scanner modules must be captured.

e accuracy here is stated with ±0.05 % of the full scale. Second, the pressure measurements are affected by the flatness tolerances (∼ 13 µm) that can be obtained for the tested seal / rotor combination, the diameter of the taps (∼ 0.3 mm) and their positioning accuracy (∼ 0.05 mm), respectively. To assess the different impacts, CFD calculations were performed for the desired experimental conditions. e results are illustrated in Fig. a)-c). As it can be seen, very accurate pressure measurements are possible away from the feed hole and the expansion region, but they are difficult in its near vicinity. Hence, it is preferable to define 3 uncertainty ranges: Pressure measurements close to the expansion region e experiments were repeated multiple times on different days. Subsequently, the measurement data could be analyzed with a special focus on its repeatability. By so doing, 3 different ranges were defined again: e repeatability error is highest near the expansion region (±0.7 %), intermediate near the feed hole region (±0.3 %) and smallest away from the feed hole and expansion region (±0.2 %).

. Clearance e axial gap x BF between the seal's and rotor's bearing face is measured by 3 Capacitec's HPT-150 non-contact probes, which are distributed along the bearing face span of the rotor. In doing so, a uniform gap width over the entire bearing face width can be ensured and quantified. e capacitive sensors were delivered including a calibration record wherein an accuracy of ±0.5 µm was specified. To achieve and maintain high accuracy, the probes are calibrated in regular intervals.

e calibration is executed separately for every probe and in accordance with the following protocol:

Calibration protocol. Figure shows the calibration unit, which was designed to calibrate the proximity probes embedded into the rotor model. e unit consists of 3 main modules: a robust holding block ensures a safe stand on the model surface; a distance rod with integrated limit stop allows for the precise positioning of the calibration bolt above the proximity probe; a micrometer screw gauge measures the adjusted clearance between proximity probe and calibration bolt. First, the unit is positioned right next to one proximity probe and the micrometer screw gauge is zeroed. Using this zero value the holder can be subsequently placed above the probe to start the calibration. While moving up the calibration bolt stepwise, the corresponding voltage is acquired. e calibration procedure is therea er repeated for all proximity probes, whereby the zero se ing is not changed in between the probes. In general, the calibration result confirms, that Uncertainty of clearance measurements. Evaluating the total uncertainty of the clearance measurement, multiple contributions need to be considered. While the sensor itself has a systematic error of ±0.5 µm, the error due to the calibration needs to be looked at as well. e calibration error comes from 2 different sources: First, a constant shi of the inputcalibration curve is expected as a result of zeroing the micrometer screw gauge. Repeated zero calibrations by several users show that the wall location can be determined with an accuracy of ±13 µm. Second, interpolating the calibration measurement points to a linear calibration curve adds a systematic error of maximum ±2.5 µm. While the listed uncertainties are caused by the calibration of the probes, an additional contribution comes from the test setup itself and has to be added to the total uncertainty of the clearance measurement. is uncertainty is found in a possible tilt of the rotor inside the rig while testing. Such a tilt can cause a nonuniform gap over the model span and needs to be quantified. On average the occurring tilt error is in the range of ±5 µm. Summing up all contributions, see Tab. , a mean uncertainty on the clearance measurement of about ±21 µm is recorded. By evaluating multiple repeated calibrations, the repeatability error of the gap measurements was assessed. It is maximum ±15 µm.

. Force

A load cell of type LCM204 from Omega is installed in the traversing mechanism of the rig (cf. Fig. ) to measure the force acting on the rotor. Due to large contact surfaces between the rotor assembly and the test rig housing systeminherent friction forces are expected. For this reason, the rig setup has several mechanisms implemented to minimize the effect of friction. Almost all mechanisms, which are originally applied for sealing purposes, can be removed: e two rubber shims of the sidewall sealing mechanism can be replaced by flush mounted metal side plates and all affecting rubber sealing elements can be removed. Additionally, the eccentric rollers can stay in a released position. Even though all these measures come along with an increased leakage around the seal, they are accepted for the single event of force measurement. However, a remaining friction force is still expected, which is why a procedure is introduced to estimate and correct for the effects of friction on the force measurement. Once the friction force is known, it can be used to quantify the force F BF acting on the bearing face, which is a crucial outcome of this study. e force can be derived from the force balance equation

F BF = F in -F out -F cav -F LC -F f , ( )
where F in represents the force due to pressure P in acting on the inlet side of the seal and F out represents the force due to the outlet pressure P out . Additionally, in between the rotor and the seal one can distinguish between the force F cav established in the air duct cavity and the force F BF coming from the bearing face. All forces are illustrated in Fig.

. e friction force can be estimated by opening the seal / rotor combination. e pressure on the le -end of the rotor then equals P in , whereas on the right it equals P out . In a hypothetical no-friction case one would expect to record the following force by the load cell:

F id = A • (P in -P out ) , ( )
where A represents the projected area. e difference between the ideal force F id and the recorded force by the load cell F LC can then be used as an estimate for the friction force

F f = F id -F LC . ( )
Since the friction force is dependent on the differential pressure across the seal, the calibration needs to be performed for all operating pressure conditions P in and P out . Some effects of hysteresis were encountered when performing this calibration procedure, that is, the friction forces differed slightly between a calibration campaign where the pressures were progressively increased or decreased. To compensate this effect, two separate calibrations curves are defined: One curve to quantify F f for an ascending pressure mission and another curve to describe F f a descending pressure mission.

Uncertainty of force measurements. F BF is the most crucial force in this study. It is gained from Eq. and depends on various quantities and their uncertainties. A first and major uncertainty hereby comes from the estimation on the effect of friction forces. e estimation is based on a semi-empirical approach and we consequential assume an uncertainty of ±(7 ... 16) % at a conservative evaluation. Second, the uncertainty caused by the load cell itself has to be considered. According to the load cell's specification the different sources (linearity, hysteresis and repeatability) contribute to a total uncertainty of ±0.3 % of the full scale output. is affects the result of F BF by maximum ±13 %. Since F BF is a calculated and not directly measured quantity, both the uncertainty coming from the pressure measurements and the uncertainty resulting from the manufacturing must be looked at as well.

e pressure measurements are very accurate so that their impact on the force result is only small with ±0.2 %. Furthermore, for each linear dimension a maximum tolerance of ±13 µm due to machining precision is known. is affects the result of F BF by less than ±0.5 %. Summing up all shares, see Tab. , the overall uncertainty of F BF is in the range of ±(20.7...29.7) %. e friction force calibration was repeated multiple times within different days. Based on these tests, the repeatability of the load cell measurement could be assessed. e mean repeatability was determined to be in the range of ±(4 ... 21) %. is uncertainty propagates to the uncertainty of F BF , which is maximum 13.5 %.

. TEST RESULTS

Fig. compares 4 representative results

of the reconstructed pressure field on the bearing face for various differential pressures ∆P := (P in -P out )/P in and axial clearances x := x BF /r p . Herein, r p is the radius of the feed pocket, which is embedded around the feed hole exit. A cubic spline interpolation was used to reconstruct the pressure measurement points.

e result basically corresponds to what has been expected and predicted by CFD. Near the feed hole region, the static pressures are maximum or rather close to P in and decrease from there in z-direction towards the low pressure P out at top edge of the bearing face (∼ radial flow regime). In y-direction the pressures first decrease and increase again while ge ing closer to the stagnation region (∼ circumferential flow regime). A direct comparison between Fig. a) and c), respectively and d), shows, that the static pressure distribution changes with the axial gap width: e static pressures near the feed hole region are higher at small gap widths than at large gap widths. Furthermore, for small gap widths the expansion region around the feed hole is more extended and the overall pressure level is increased. What can also be observed is that the static pressure development qualitatively does not change significantly with the differential pressure, see antitatively, the overall static pressure at the bearing face obviously increases with an increase of the differential pressure.

Figure . Comparison between FBF and FBF,int

Based on the static pressure measurements the force acting on the bearing face can be determined. e reconstructed field of static pressures is integrated over the area of the bear-ing face such that a erwards this integrated force F BF,int can be compared to the force F BF coming directly from the load cell. By doing so, the accuracy of the force measurements can be assessed. Fig. represents the comparison between FBF,int := F BF,int /(P in • A BF ) and FBF := F BF /(P in • A BF ) for the case of ascending differential pressures and a non-varying axial gap width. e difference between both force results is maximum ±10 %. For descending differential pressures, the result looks alike and is therefore not shown here. e normalized force decreases with normalized differential pressure because the force increase is less than the increase of differential pressure. A linear relationship between the normalized force and the differential pressure is furthermore observable.

. CONCLUSION is paper presented a newly designed and built test rig for studying advanced sealing technology. e rig allows for a quick exchange of different seal / rotor designs and provides detailed pressure measurements on the bearing surface, accurate mass flow measurements and load cell measurements to investigate the seal / rotor force balance. Additionally, the rig allows for a very precise se ing of the major seal operating parameters: the axial gap width can be adjusted during operation; for se ing the radial gap width a quick disassembly is necessary; high differential pressure ratios, which are derived and scaled from real engine conditions, can be applied accurately.

e research on the rig results in high quality test data, which can be used to investigate the aerodynamic characteristics of advanced seals. A first focus was placed on the development of static pressures on the bearing surface. At increasing axial gap widths and non-varying differential pressures across the seal the pressure level on the bearing surface decreases. e expansion region furthermore weakens with increasing gap width. At increasing differential pressures and non-varying axial gap widths the pressure level increases quantitatively, but its distribution stays qualitatively alike.

e static pressure results can additionally be used to determine the force acting on the bearing face so that the accuracy of the direct force measurement can be evaluated. Upcoming research on the rig will cover important quantities such as the stiffness or stability of the air bearing. e gained knowledge can subsequently be fed into rotating tests with a focus on a dynamic seal characterization and it can also be used to validate numerical models for the air bearing flow. is in the end allows for a goal-oriented optimization of new seal designs.

Figure .

 . Figure . Design overview

Fig. shows the

  Fig. shows the design overview of the new test facility. Pressurized air enters the rig through the air inlet and is guided to the inside of the rig.e extended geometry of the inlet duct hereby ensures a uniform inlet flow across the entire span of the seal. A er the air flow passes the test build it exits the rig through the extended duct and outlet, whereby the test build could either be a seal or a bearing. e test build is interchangeable and always consists of two units: A 2D rotor and a seal model. Fig.a) represents a sectional view of an exemplary test build. As it is schematically shown, the seal possesses two different types of flow ducts: Equally distributed feed holes serve as supply ducts to establish the air film at the bearing face, whereby multiple air ducts, which are also equally distributed over the span of the model, are in place to ensure a low static pressure in

Figure .

 . Fig. shows the design overview of the new test facility. Pressurized air enters the rig through the air inlet and is guided to the inside of the rig.e extended geometry of the inlet duct hereby ensures a uniform inlet flow across the entire span of the seal. A er the air flow passes the test build it exits the rig through the extended duct and outlet, whereby the test build could either be a seal or a bearing. e test build is interchangeable and always consists of two units: A 2D rotor and a seal model. Fig.a) represents a sectional view of an exemplary test build. As it is schematically shown, the seal possesses two different types of flow ducts: Equally distributed feed holes serve as supply ducts to establish the air film at the bearing face, whereby multiple air ducts, which are also equally distributed over the span of the model, are in place to ensure a low static pressure in

  Figure . Axial traversing mechanism

Figure .

 . Figure . Mechanisms to adjust the radial gap width

  e data acquisition is accomplished by a NI Compact-DAQ, which is equipped with multiple input and output NI modules. All modules have a 16-bit resolution and 200 kS/s aggregate sampling rate.. Mass flow e test rig is connected to the local GE screw compressor, which delivers up to 0.75 kg/s pressurized air with maximum 10 bar absolute pressure. e mass flow measurement is done by one of two parallel Coriolis flow meters of type Promass 80F by Endress+Hauser, both installed upstream of the rig, and each flow meter is adjusted for a different mass flow range. While the first meter is for mass flow rates up to 0.1 kg/s, the second covers mass flow rates between 0.1 kg/s and 0.75 kg/s. e inlet mass flow rate can be either mass flow or pressure controlled. Downstream of the rig a control valve is installed to also control the outlet mass flow. By a controlled closing of the valve the backpressure required downstream of the seal can be set accurately.Uncertainty of mass flow measurements.

Figure .

 . Figure . Expected pressure distribution on the bearing surface of the rotor for representative operating conditions gained from CFD simulations

Figure .

 . Figure . Uncertainty on the pressure measurements caused by a) the flatness tolerance on the bearing surface, b) the diameter of the pressure taps and c) the positioning accuracy of the pressure taps

Figure .

 . Figure . Calibration unit

Figure .

 . Figure . Force balance in axial direction

Figure .

 . Figure . Pressure distribution on the bearing face for varying differential pressures and axial gap widths

  Fig. a) and b), respectively c) and d).

Table .

 . Rig capabilities

	Maximum pressure:	10 bar
	Maximum mass flow rate: 0.75 kg/s
	Horizontal traverse:	up to 58 mm
	Vertical adjustment:	at least ±0.38 mm
	Measurements:	Mass flow, clearance,
		static pressure, force
	into the test rig concept to minimize system-inherent friction
	forces during the force measurement. A summary of the
	main rig capabilities is given in Tab. .
	. Design overview

Table .

 . Uncertainty of static pressure measurements

	S	U	[%]
	DSA accuracy	±0.05
	Flatness tolerance
	-Away from feed hole ±1.5
	-At expansion region	±6
	-At feed hole region	±3
	Diameter of pressure taps
	-Away from feed hole ±0.01
	-At expansion region	±5
	-At feed hole region	±0.2
	Positioning of pressure taps
	-Away from feed hole ±0.1
	-At expansion region	±10
	-At feed hole region	±1
	Overall uncertainty
	-Away from feed hole ±1.66
	-At expansion region	±21.05
	-At feed hole region	±4.25

Table .

 . Uncertainty of clearance measurements

	S	U	[µm]
	Proximity probe accuracy ±0.5
	Zero adjustment	±13
	Calibration interpolation	±2.5
	Tilt of the rotor	±5
	Overall uncertainty	±21
	Table . Uncertainty of force measurements
	S	U	[%]
	Estimation of friction force	±(7 ... 16)
	Load cell accuracy	±13
	Machining precision of A BF	±0.2
	Pressure measurement accuracy ±0.5
	Overall uncertainty	±(20.7...29.7)

Repeatability of clearance measurements.

ACKNOWLEDGMENTS

is project has received funding from the European Union's Horizon 2020 research and innovation programme under grant agreement No. 653941. e authors would also like to thank various researchers at GE Global Research for useful discussions and inputs during the rig design.

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-.


Edouard Delangle

Bertrand Lecordier

Corine Lacour

Armelle Cessou
email: armelle.cessou@coria.fr

Back-supported stratified flame propagation in lean and nonflammable mixtures

In an effort to reduce pollutant emissions and increase energy efficiency, partially premixed combustion has been integrated into many new combustion technologies. The present study investigated lean back-supported flames in a stratified combustion regime. Outwardly propagating flames were observed following ignition under laminar stratification conditions generated in a constant volume vessel. The quantitative analysis of the flame properties relied on simultaneous PIV measurements to obtain local flame burning velocities and stretch rates and used anisole-PLIF measurements to calculate the equivalence ratio. Simultaneous OH-PLIF and OH-gradient measurements were used to differentiate between the burned gas boundaries and the active flame front. This differentiation was necessary to investigate the nonflammable mixture. Simultaneous OH-and anisole-PLIF measurements were used to estimate the thermal flame thickness. Two flame families were investigated: in family A the flame was ignited in a lean mixture (φ=0.6) with a rich stratification; in family B the mixture in the chamber was nonflammable. In rich mixtures ignition compensated for the non-equidiffusive effects of the lean propane flame. Both a flammable and a nonflammable mixture were investigated to evaluate the time scales of the back-supported propagation for the given stratification. The enhanced combustion regime allowed the flame to propagate with an active flame front, even in the nonflammable mixture for a few milliseconds before the flame extinguished.

Introduction

In addition to providing a canonical description of premixed and nonpremixed combustion, many studies have addressed partially premixed combustion regimes, in which the oxidant and the fuel species are inhomogeneously mixed before reacting. These conditions are introduced intentionally to reduce pollutant emissions and increase energy efficiency in a variety of applications ranging from gas turbine combustors [1,[START_REF] Gicquel | Large Eddy[END_REF] to stratified-charge direct-injection engines [START_REF] Yang | Tailoring[END_REF]4].

Moreover, identical conditions are often encountered in practical devices in which the combustion regime involves mixture inhomogeneities (local extinction in two-phase combustion). Consequently, there is renewed interest in seeking a better understanding of partially premixed combustion, as reviewed by Masri [5]. The present paper examined stratified combustion in a premixed combustion regime with a compositionally inhomogeneous mixture in which the flame propagated successfully [5,6]. For lean back-supported flames, stratified combustion results in a higher burning rate [7][START_REF] Richardson | Combustion Theory and Modelling[END_REF][START_REF] Jiménez | [END_REF], active heat release [6], a reduced flame thickness [START_REF] Richardson | Combustion Theory and Modelling[END_REF], extended flammable limits [7,10], and a better stretch resistance [11]. Experimental and numerical investigations have concluded that these effects occur because of increased temperature of burned gases and species flux from products [7,12]. The present study experimentally investigated whether higher burning rates and thermal flame thinning enhance combustion. A time scale of these potential benefits is proposed for a lean back-supporting flame. The outwardly propagating flame configuration is a well-established and extensively used method for measuring burning velocities in homogeneous mixtures; the method is suitable for the direct determination of the stretch rate [13]. In this configuration, the burning velocity is often deduced from the flame speed (S " ) and the displacement speed of the burned gas, based on assuming a stationary burned gas and a known burned gas temperature. For inhomogeneous mixtures these assumptions are no longer valid, but the burning velocity [14] can be measured directly by determining the displacement speed with respect to the unburned gas 𝑈 $ , defined as 𝑆 & -𝑈 ( [15,16], where U * is the maximum velocity upstream of the flame. Here, outwardly propagating flames were observed following ignition under laminar stratification conditions generated in a constant volume vessel. The flame was first ignited in a richer mixture, and the analysis is conducted once the flame propagates in the lean homogenous mixture, as has been done previously [11]. This sequence was chosen to establish whether prior burning of a richer mixture affects the flame properties in the lean homogeneous mixture. The analysis was restricted to measuring flame segments that were primarily exposed to the curvature stretch effect with limited flow non-uniformity influences and with almost the same evolution from the point of ignition throughout the stratified mixture. The study used lean mixtures in which propane-air flames are sensitive to non-equidiffusive effects, which made it possible to also investigate flame stretch rates. As done previously, simultaneous PIV and anisole-PLIF measurements were used to analyse the flame properties [11]. Local flame burning velocities and stretch rates were derived from flame speed measurements, fresh gas velocities, and the principal curvature radii based on a PIV algorithm proposed in a previous paper [14]. The investigation of nonflammable mixtures requires burned gas boundaries to be differentiated from the active flame front. Simultaneous OH-PLIF measurements showed that the OH-gradient could be used to distinguish burned gas interfaces from active flame fronts.

Two flame families were investigated: family A, for which the flame was ignited in a lean mixture (j = 0.6) with a rich stratification, corresponding to a mixture bordering on the flammability limit; and family B, for which the mixture in the chamber (j = 0.4) was below the flammability limit. Flames were also ignited in several richer mixtures ranging from j = 0.8 to 1.5.

Balusamy et al. [11] have shown that the properties of back-supported stratified flames depend on the ignition conditions of the stratified layer. In richer mixtures, ignition compensates for the non-equidiffusive effects of a lean propane flame. The increase in flame propagation observed in lean homogeneous mixtures is caused as much by the flame's resistance to stretch as by an increase, in flame speed and burning velocity. The decay time of this influence of the stratification, called the memory [12], was determined in the present study. It is due to the fact that the flame properties are considered at the equivalence ratio in fresh gas, which differs from the effective equivalence ratio of the flame front (peak heat release) [6]. The back-supporting behaviour of the flame thus appears as a memory effect that reflects the amount of fuel previously consumed by the flame. Aiming to complement previous research [11], the present study used various mixture stratification amplitudes with both flammable and nonflammable mixtures to evaluate the flame's ability to propagate in these mixtures through a back support structure.

Experimental set-up

To investigate flame propagation in a laminar stratified mixture field, a constant volume combustion chamber was used in combination with PIV, anisole-PLIF, and OH-PLIF techniques.

Combustion facilities

Experiments were conducted in a 60×60×160 mm 3 constant volume combustion chamber equipped with three UV quartz windows. The chamber was initially filled with a lean homogeneous propane-air mixture with either j = 0.6 or j = 0.4 at atmospheric pressure. The stratified mixture field was created by injecting a laminar jet of a richer mixture into the chamber with an injection system, providing controllable and repeatable mixture stratifications. Two flame families were investigated (Tab. 1). The stratified mixture was spark-ignited by energizing two thin electrodes with a diameter of 0.4 mm placed at the centre of the combustion chamber. Previous experiments in this combustion chamber have shown that spark energy and electrode heat loss did not affect flame propagation. The limited volume of the chamber had little influence on flame propagation for the investigated time frame (pressure increase < 40 mbar), and due to the direct measurement procedure used for its determination, the burning velocity measurement is non-affected by the limited chamber volume [14].

Tab. 1-Investigated flame families: mixture at rest in chamber (𝜑 ,-); equivalence ratio at ignition (𝜑 .( )

Laser diagnostics

The optical setup is presented in Figure 1. The PIV setup consisted of a double-cavity Nd:YAG laser (Twins Ultra, Quantel 30mJ@532nm), and CCD camera (Hamamatsu C9300 4Mp in 12 bits) that was equipped with a Nikkor lens (135mm, f/d max = 2, f/8). The anisole-PLIF setup used a fourth harmonic generator with an Nd:YAG laser (Quantel, YG180) that delivered 30 mJ to ensure a linear regime. A dye laser (Quantel, TDL90 20 mJ) pumped by a second harmonic generator with an Nd:YAG laser (Quantel, Brillant B) was used as the OH-PLIF laser source. Both LIF radiation measurements were collected with ICCD cameras (Roper Scientific PIMAX512), equipped with a 98 mm UV lens (CERCO 2178, SODERN, f/2.8) for anisole-PLIF and a 105 mm lens (Nikkor 105, f/4.5) for OH-PLIF. Dichroic mirrors were used to optically combine the three laser beams, and the overlapping laser sheets were formed by combining a spherical (f = 1 m) lens with two cylindrical lenses (f = -25 mm, f = +100 mm).

The airflow was seeded with olive oil particles, and the propane flow was seeded with anisole (0.015% in volume in the stoichiometric propane-air mixture). OH-PLIF images were corrected for laser sheet inhomogeneities by collecting a LIF reference image obtained from homogeneous images of a rich propane-air mixture seeded with anisole. PIV and LIF radiations were separated with a dichroic mirror, which had a cut-off wavelength of 350 nm. Fluorescence was collected for anisole in the spectral range of 275 to 350 nm with a 10 mm thick liquid filter (0.6% vol. of toluene in iso-octane) and with a combination of Schott filters (UG11 and WG280, 3 mm). The time delay between the two PIV pulses was adjusted to the velocity magnitude. The PLIF laser pulses were fired at intervals of 2 µs and positioned in the middle of the PIV time delay at intervals of 2 µs. A calibration grid was used to determine the high-order mapping functions of the three cameras. This calibration ensured the perfect overlap of the three measurements, especially for the OH-PLIF off-axis acquisition, which viewed the laser sheet at a slightly tilted angle, and is essential for providing a good accuracy of the flame thickness measurement described afterwards.

Active flame front identification

Once the flame reached the nonflammable mixture, j ch = 0.4, it became necessary to differentiate between the burned gas boundaries and the active flame front. For the interface of fresh and burned gas, numerous studies have proposed to use OH-gradient levels to identify flame fronts [17][18][START_REF] Bayley | 15th Int Symp on Applications of Laser Techniques to Fluid Mechanics[END_REF]. After obtaining the OH outline, the OH signal was extracted along its normal. This normal profile was interpolated with a complementary error function (erfc) before calculating the noise-free gradient. To ensure that OH-gradient levels are suitable for determining active flame fronts, OH-gradient mapping was first compared with OH×CH2O PLIF to verify that this simple diagnostic was sufficiently able to localize local flame extinctions. This preliminary experiment is performed in a Bunsen burner conical flame, where a local extinction was produced by a small air jet. In Figure 2, a threshold value of OH-gradient is visible where the flame is extinguished. The value of this threshold value was determined for homogeneous outwardly propagating flame. The OH gradient has been measured for homogenous outwardly propagating flames at various equivalence ratios of up to j = 0.6, while in the present configuration, the flame kernel could not self-propagate below j = 0.62. Figure 3 shows the PDFs of the OH gradient according to equivalence ratio. The peak values decreased from the stoichiometric mixture to the lean composition. The flames with the two lowest equivalence ratios (j = 0.6 and j = 0.62) for which the homogenous flame could not selfpropagate, presented PDFs clearly separated from the others. Thus, the critical threshold was defined as 17 000 a.u, after correcting the laser sheet profile.

Flame thickness determination

The equivalence ratio was measured based on anisole-PLIF, following a shot-to-shot laser sheet correction according to the procedure outlined by Balusamy et al. [11]. The OH-PLIF marked high temperature burned gas (T>1500 K), whereas anisole pyrolysis was noticeable around 800 K [START_REF] Nowakowska | [END_REF]. After calibrating the mapping functions of the three cameras, the distance between these two isotherms, determined by superpositioning the simultaneous OH-PLIF and anisole-PLIF measurements, was used to define a thickness labeled 𝛿 012 . The local distance between these two isotherms, which varies when flame has crossed stratification (Figure 4), is measured all around the flame outline from the superposition of the simultaneous OH-PLIF and anisole-PLIF.

Figure 5 These new measurements are combined with the measurements of burning velocity and stretch rates measured all around the stratified flames from the procedure of Balusamy et al [14], providing a direct measurement based on the difference between the local flame speed and the fresh gas velocity at the entrance of the preheat zone.

Results and discussion

The flame properties were analysed after the flame began propagating in the lean homogeneous mixture filling the combustion chamber that is, for 𝑡 ≥ 2 𝑚𝑠. To limit the analysis to conditions comparable to outwardly propagating flames in a homogeneous mixture, the conditional analysis relied on the same criteria as used by Balusamy et al. [11]: Measurements were conditioned on the equivalence ratio in the chamber, (𝜑 ,-= ± 0.05) and the moderate tangential velocity (-5% to 5% of the 𝑈 ( average). An additional condition limited the investigation to flame segments propagating almost perpendicularly to the stratified layer: an angular sector of ±20° from the ignition location was chosen (Figure 6). This angular condition restricted the flame analysis to segments mainly exposed to the flame curvature effect with limited flow non-uniformity influences and with almost the same residence time in the homogeneous lean mixture. Given these new criteria, sampling was restricted to flame segments with an almost constant stretch rate at every moment after ignition. Consequently, the stretch rate was averaged at each time, and the flame properties were not extrapolated to a zero stretch rate, as done by Balusamy et al. [11].

Figure 6 : Combined measurements and angular sector of the conditional analysis

Family A: propagation in the lean flammable mixture Figure 7 presents the burning velocity according to time. For ignitions with j ig = 1.2 in family A at 0.34 m/s, the initial burning velocity of the flame segment propagating in the lean mixture (t = 2 ms) was significantly higher than expected for lean stretched flames with j = 0.6 and 0.7: the unstretched velocity was 𝑈 $ °= 0.12 𝑚/𝑠 for j = 0.6 and 𝑈 $ °= 0.20 𝑚/𝑠 for j = 0.7. The stretch effect resulted in decreased burning velocities in lean propane-air mixtures. In the rich mixture, ignition initially led to high values of 𝑈 $ , despite the high stretch rate (Figure 8).

For the two cases with j ig = 1.2, the velocity 𝑈 $ showed identical time evolutions with an asymptotic trend of the exponential decrease for t > 6ms, where 𝑈 $ was almost constant (Figure 7 and Figure 8). The asymptotic values were close to those of the unstretched velocity, with 0.20 m/s for j ch = 0.7, and 0.14 m/s for j ch = 0.6. Exponential modelling of 𝑈 $ (𝑡) (Tab. 2), from t=2 ms to the asymptotic behaviour, shows that the decay time was 1.6 𝑚𝑠 for j ch = 0.7, and 1.9 ms for j ch = 0.7. The asymptotic behaviour was attained more rapidly for j ch = 0.6 because of the lower stratification amplitude (j ig = 0.6 to j ch = 0.7).

-Exponential modelling of the burning velocity decay

Once the asymptotic behaviour was attained (t>5 ms), a slight increase in 𝑈 $ was observed for j ch = 0.7 corresponding to the decreasing stretch effect observed for increasing flame radii. For j ch = 0.6, the minimum value of 𝑈 $ was very close to the value of 𝑈 $ °, and a slight decrease was observed, in agreement with previous observations by Balusamy et al. [11]. This behaviour illustrates the competition between the stretch effects, which tended to reduce the burning rate and the backsupport behaviour of the flame. After igniting in the homogeneous mixture with j ch = 0.6, outwardly propagating spherical flames were no longer able to selfpropagate due to overly intense stretch effects occurring with small flame radii. In the stratified case (j ig = 1.2 to j ch = 0.6) the flame succeeded to propagate despite its small radius, as confirmed by the OH gradient values (Figure 3), which were higher than the defined extinction threshold (17 000 a.u.). The flame was sustained by the memory effect, but was slowed down by the stretch effects. The gain provided by the memory effect decreased with time, and the influence of the stretch effect also decreased, resulting in a relatively constant burning velocity: for the investigated time range, the two effects were balanced, with a slightly faster decay observed for the memory effect.

Overall, these results show that ignition in a slightly rich mixture resulted in a flame that rapidly propagated in the lean mixture despite the high stretch rates. This velocity decreased as the flame propagated in the lean mixture, with a decay time of slightly less than 2 ms. This high burning velocity was related to the discrepancy between the local equivalence ratio in the flame front (peak heat release) and in fresh gas. The flame propagation then adopted an asymptotic behaviour with a burning velocity close to the fundamental velocity. The physical properties of the burned gas (here, mainly the high temperature) compensated for the non-equidiffusive effects of the lean stretched propane-air flame, which resulted in a leaner mixture in front of the flame with a lower burned gas temperature. These observations were made for stratifications that led to few partial oxidation products in burned gas (peak equivalence ratio of 1.2). For j ig = 1.5 to j ch = 0.6 partial oxidation products are more abundant in burned gas. At 2 ms, just after the flame had left the stratified layer, 𝑈 $ was equal to 0.33 m/s, which is roughly the same value as for j ig = 1.2 to j ch = 0.6 and for j ig = 1.2 to j ch = 0.7. Because stretching had an effect on the flame, these identical velocities possibly indicate either an identical equivalence ratio within the flame front for the three cases, or a much richer flame whose burning velocity was increased by the stretch effects. The time decay was slower and differed from the exponential decay observed for the two other cases of family A. The concavity of the curve 𝑈 $ (𝑡) differed from the ignition cases with j ig = 1.2, demonstrating that the burning velocity increased more strongly between 3 and 6 ms. This additional combustion enhancement is visible in Figure 8, in which the case with j ig = 1.5 to j ch = 0.6 represents a faster burning velocity for stretch rates in a time range of 3 to 6 ms. These results show that the back-support behaviour of the flame was enhanced by ignition in a richer stratified layer: the excess in partial oxidation products in burned gas enriched the flame front for longer and/or led to a secondary reaction in the burned gas [23]. Due to the relative change in the curve concavity 𝑈 $ (𝑡), a decay time cannot be determined through exponential modelling. However, the same asymptotic behaviour as for j ig = 1.2 to j ch = 0.6 was attained after a slightly longer time (t>6 ms); this behaviour was observed for all measured properties (Figure 7).

For the three cases of family A, the flame thickness 𝛿 012 showed the same time evolution. At t= 2 ms, despite the lean mixture in fresh gas, the flame was relatively thin with a thickness close to the minimum value measured in near-stoichiometric homogeneous flames for the cases j ig = 1.2 to j ch = 0.7 and j ig = 1.5 to j ch = 0.6 the flame was slightly thicker for the case j ig = 1.2 to j ch = 0.6 with 𝛿 012 = 400 µ𝑚, corresponding to j = 0.85 for a homogeneous flame (Figure 9). These smaller thickness values are consistent with high values of 𝑈 $ . Then, 𝛿 012 increased over time to attain a roughly constant value close to that measured in the homogeneous flame with j ch = 0.6 and 0.7 (Figure 5).

Family B: propagation in a lean nonflammable mixture

Flames of family B, for which the lean mixture was below the flammability limit, were investigated to determine whether this combustion enhancement also occurred in nonflammable mixtures and over which time scales. For these four flames, the OH-LIF gradient was used to differentiate the border of hot burned gas from the active flame front (Figure 9). At t=2 ms the OH-LIF gradient values were much higher than the extinction threshold (17 000 a.u./mm). Depending on the stratification amplitude, the gradient's progressive decay and the extinction occurred at different times greater than 4 ms and increased with the equivalence ratio at ignition. Thus, ignition in a richer stratified mixture allowed the flame to propagate in the nonflammable mixture: the flame ignited in the lean mixture (j ig = 0.8) as well as in the rich mixture (j ig = 1.5). Figure 7 shows the burning velocities over time for active reaction zones (with an OH-LIF gradient higher than 17 000 a.u/mm). The conditions for which the flame extinguished (open symbols in Figure 7), indicate that the burning velocity was non-zero due to measurement inadequacies in this case. At 2 ms, the case j ig = 0.8 to j ch = 0.4 had a burning velocity of 23 cm/s, a value close to the velocity measured for j ig = 0.8 to j ch = 0.8 (Figure 7), which is evidence of enhanced combustion in the nonflammable mixture. This increase was observed even if the flame was ignited in a lean stratified layer (peak equivalence ratio of 0.8). In this case, enhanced combustion could only have resulted from an increase in the burned gas temperature. This enhancement decreased over time, and flame extinction occurred between 4 and 5 ms. The three other flames (j ig = 1.0 to j ch = 0.4; j ig = 1.2 to j ch = 0.4 and j ig = 1.5 to j ch = 0.4) had nearly identical burning velocities of around 0.3 m/s at 2 ms. The burning velocities decreased over time; the slowest decay time was observed with the richer stratification. Exponential modelling of the decay with 𝑈 T = 0 𝑚/𝑠 took into account the extinction occurrence. The following decay times were determined: 4 ms for j ig = 1.5 to j ch = 0.4; 2.8 ms for j ig = 1.2 to j ch = 0.4; and 2 ms for j ig = 1.0 to j ch = 0.4 (Tab. 2). For j ig = 0.8 to j ch = 0.4 the decay time was slightly longer (2.4 ms), or at least roughly equivalent, considering the measurement uncertainty. This result demonstrates the combined effects of stretch and stratification. For rich propane-air mixtures, stretching tended to increase the burning velocity. Thus, in these cases, combustion was enhanced by both stratification and stretching, which led to higher burned gas temperatures. During propagation in the lean mixture, these positive influences decreased simultaneously. For ignition in the lean mixture with j ig = 0.8 to j ch = 0.4, the burned gas temperature increased as a result of ignition in the stratified mixture, whereas stretch tend to decrease this temperature. During propagation in the lean mixture, the effect of stretch on combustion decreased progressively, leading to a slower decay. However, the increase in decay time remained moderate, and the flame was rapidly extinguished (t<5 ms) because of the minor presence of a back-support structure for the stratification in this case.

The combustion enhancement was also noticeable with the small values of 𝛿 012 at 2 ms (Figure 5). The flame thickness was around 500 µm, and was as thin as 400 µm for j ig = 1.5 to j ch = 0.4. These values are close to the value measured in a homogeneous flame with j = 0.8. The flame became thicker over time, and extinction occurred at 𝛿 012 > 1.2 𝑚𝑚. For the entire duration, the richer the stratification was, the thinner the flame became. Additionally, for a given stretch rate, the fastest burning velocity occurred in the richer stratification (Figure 8). These results show that a richer stratification resulted in longer lasting, increasingly enhanced combustion in the lean nonflammable mixture.

Conclusion

Simultaneous PIV, anisole-PLIF, and OH-PLIF measurements were used to experimentally investigate the possibility of lean-burn combustion. Outwardly propagating flames were ignited under laminar stratification conditions generated in a constant volume vessel. After ignition in a rich stratified layer, the flame properties (burning velocity, stretch, thermal thickness 𝛿 012 ) were analysed in a lean flammable mixture bordering on the flammability domain (family A) and in a nonflammable mixture (family B).

The behaviour of flames in family A showed that the lean-mixture flames benefited from ignition in the rich stratified layer, which led to faster burning velocities despite the high stretch rate that usually inhibits combustion in such a lean mixture (𝜑 = 0.6). The burning velocity decreased toward fundamental values, with a decay time close to 2 ms for the investigated configuration. The memory effect compensated for the stretch effect, resulting in a fundamental burning velocity at a non-zero stretch rate. For family B, the flame maintained an active flame front in the nonflammable mixture over time, which increased with the equivalence ratio in the stratification. The combustion enhancement was also apparent in the flame thickness values 𝛿 012 : the flame became thinner in the early stage of propagation in the lean mixture. The thickness values were depended on the effective equivalence ratio in the flame front and on the flame thickness, whereas the burning velocity decreased. Stratified flames were used to construct a database suitable for validating numerical simulations and investigating the influence of-in contrast to homogeneous combustion-higher burned gas temperatures and excess in partial combustion products, which could enrich the flame front and compensate for the negative thermodiffusive effects of the lean propane flame (Le<1).

Figure 1 :

 1 Figure 1: Experimental set-up

Figure 2 :

 2 Figure 2: simultaneous OH-gradient and CH2OxOH LIF in a Bunsen burner conical flame, where a local extinction is made by injection a small air jet

  shows a comparison of 𝛿 012 values obtained for different homogeneous flames with the thermal thickness 𝛿 3-= 𝑇 & -𝑇 6 ∇𝑇 89: , calculated by modelling the unstretched 1D propane-air flame carried out by Cantera and the chemical scheme provided by Jerzembeck et al. [21]. The similarity of the resulting values shows that 𝛿 012 can be used to evaluate the local flame thickness.

Figure 3 :

 3 Figure 3 : PDF of OH-PLIF gradient normal to the PLIF outline

Figure 5 :

 5 Figure 5 : Thickness according to equivalence ratio: comparison of δ =>? with the thermal thickness given by the plane flame calculation for propane-air

Figure 7 .

 7 Figure 7. Burning velocity according to time: family A (red), B (blue), and homogeneous flames (green) with 𝜑 = 0.8. Extinguished flame conditions (open symbols).

Figure 8 .

 8 Figure 8. Burning velocity according to stretch rate

Figure 9 .

 9 Figure 9. OH-PLIF gradient according to time

Acknowledgment

We gratefully acknowledge the support of the French Agence Nationale de la Recherche [National Research Agency] (ANR) and the French Centre National de la Recherche Scientifique (CNRS) through the ACTING-CO2 program, provided in collaboration with the French engine research grouping Groupement Scientifique Moteurs (GSM). We would like to thank Dr G. Lartigue for his help in modelling and for the fruitful discussions.


E Delangle

B Lecordier

C Lacour

A Cessou

Simultaneous measurements of PIV, anisole-PLIF and OH-PLIF for investigating back-supported stratified flame propagation in lean and nonflammable mixtures

Keywords: PIV, anisole-PLIF, OH-PLIF, combustion, laminar stratified flames, nonflammable mixture, memory effect

In an effort to reduce pollutant emissions and increase energy efficiency, partially premixed combustion has been integrated into many new combustion technologies. The present study investigated lean back-supported flames in a stratified combustion regime. This strategy leads to hybrid combustion regimes, ranging between fully premixed and fully non-premixed reactants, with a large panel of flame structures and properties requiring to be characterized. Outwardly propagating flames were observed following ignition under laminar stratification conditions generated in a constant volume vessel. The quantitative analysis of the flame properties relied on simultaneous PIV measurements to obtain local flame burning velocities and stretch rates and used anisole-PLIF measurements to calculate the equivalence ratio. Simultaneous OH-PLIF measurements were used to differentiate between the burned gas boundaries and the active flame front. This differentiation was necessary to investigate the nonflammable mixture. The OH-gradient measurement proved to be suitable for distinguishing burned gas interfaces from active flame fronts. Simultaneous OH-and anisole-PLIF measurements were used to estimate the thermal flame thickness. Two flame families were investigated: in family A the flame was ignited in a lean mixture (φ=0.6) with a rich stratification; in family B the mixture in the chamber was nonflammable. In rich mixtures ignition compensated for the non-equidiffusive effects of the lean propane flame and reinforced the flame's stretch resistance. Both a flammable and a nonflammable mixture were investigated to determine the time scales of the back-supported propagation for the given stratification. The enhanced combustion regime allowed the flame to propagate with an active flame front, even in the nonflammable mixture. Combustion continued for a few milliseconds before the flame extinguished. The richer the stratification, the longer the combustion lasted in the nonflammable mixture.

Introduction

Partially premixed combustion is intentionally introduced in many new combustion technologies to reduce pollutant emission or increase energetic efficiency. This strategy leads to hybrid combustion regimes, ranging between fully premixed and fully non-premixed reactants, with a large panel of flame structures and properties requiring to be characterized. The paper investigates stratified flames, which concerns premixed combustion with compositionally inhomogeneous mixture where the flame succeeds to propagate (Masri 2015). In the purpose of investigating lean back-supported flames by stratified combustion and of focusing on flames propagating towards lean non-flammable mixtures, the present paper presents a combination of simultaneous diagnostics: PIV, anisole-PLIF, OH-PLIF, and proposes a new measurement of flame thickness directly linked to the thermal flame thickness. These simultaneous diagnostics provide various flame properties like burning velocity, stretch rate, equivalence ratio and velocity fields in fresh gases, flame thickness and identification of the active flame front. These quantities are of major interest for investigating lean back-supported flames, where stratified combustion leads to enhancement of burning rate [START_REF] Jiménez | Numerical Simulation and Modeling for Lean Stratified Propane-Air Flames[END_REF][START_REF] Marzouk | Dynamic response of strained premixed flames to equivalence ratio gradients[END_REF][START_REF] Richardson | Effects of equivalence ratio variation on lean, stratified methane-air laminar counterflow flames[END_REF], heat release [START_REF] Zhou | The behaviour of laminar stratified methane/air flames in counterflow[END_REF], thinner flame thickness [START_REF] Richardson | Effects of equivalence ratio variation on lean, stratified methane-air laminar counterflow flames[END_REF], extended flammable limits [START_REF] Kang | Departure from quasi-homogeneity during laminar flame propagation in lean, compositionally stratified methane-air mixtures[END_REF]Kyritsis 2007, Marzouk, Ghoniem and[START_REF] Marzouk | Dynamic response of strained premixed flames to equivalence ratio gradients[END_REF] and better resistance to stretch [START_REF] Balusamy | Laminar propagation of lean premixed flames ignited in stratified mixture[END_REF]. Even if experimental and numerical investigations conclude that these effects are due to additional temperature of burned gas and species flux from products [START_REF] Marzouk | Dynamic response of strained premixed flames to equivalence ratio gradients[END_REF]Najm 2000, Pires da Cruz 2000), these flames still suffered from incomplete understanding.

The flames investigated consist in outwardly propagating flames. This flame configuration is chosen since it is a well-established and extensively used method for measuring burning velocity in homogeneous mixture and is convenient for the direct determination of stretch rate [START_REF] Egolfopoulos | Advances and challenges in laminar flame experiments and implications for combustion chemistry[END_REF]. In this configuration the burning velocity is often deduced from the measurement of the flame speed, Sb displacement speed of the burned gas, using the assumptions of stationary burned gas and of the knowledge of the burned gas temperature. For inhomogeneous mixtures these assumptions are not valid anymore, but the direct measurement of burning velocity [START_REF] Balusamy | Direct measurement of local instantaneous laminar burning velocity by a new PIV algorithm[END_REF]) can be performed from the determination of the displacement speed with respect to the unburned gas, Un, determined as Sb-Ug [START_REF] Andrews | The burning velocity of methane-air mixtures[END_REF]Bradley 1972, Bradley 1996), where Ug is the maximum velocity upstream of the flame. The flames are outwardly propagating after ignition in the middle of laminar stratifications generated in a constant volume vessel. The analysis is conducted once the flame propagates in the lean homogenous mixture after having been ignited in richer mixture as in a previous work [START_REF] Balusamy | Laminar propagation of lean premixed flames ignited in stratified mixture[END_REF]. OH-PLIF is implemented for investigation in nonflammable mixture, where the distinction between burned gas boundaries and active flame front is required. It is shown that OH-gradient measurement is suitable for distinguishing burned gas interfaces from active flame front. Simultaneous OH and anisole PLIF provides an estimation of the thermal flame thickness. The interest of all the laser diagnostic measurements is shown by presenting the flame properties for various amplitudes of the stratified layers.

Experimental setup

The optical setup is presented in Figure 1. The PIV setup consisted of a double-cavity Nd:YAG laser (Twins Ultra, Quantel), delivering pulses of 30 mJ at 532 nm, and a full-frame interline CCD camera (Hamamatsu C9300) that was equipped with a Nikkor lens (135mm, f/dmax = 2, f/8) and provided a 12-bit image pair of 2048×2048 pixels. The anisole-PLIF setup used a fourth harmonic generator with an Nd:YAG laser (Quantel, YG180) that delivered 30 mJ of power to ensure a linear regime. A dye laser (Quantel, TDL90) (20 mJ) pumped by a second harmonic generator with an Nd:YAG laser (Quantel, Brillant B) was used as the OH-PLIF laser source.

Both LIF radiation measurements were collected with intensified CCD cameras (Roper Scientific PIMAX512), equipped with a 98 mm UV lens (CERCO 2178, SODERN, f/2.8)) for anisole-PLIF and a 105 mm lens (Nikkor 105, f/4.5) for OH-PLIF. Dichroic mirrors were used to optically combine the three laser beams, and the overlapping laser sheets were formed by combining a spherical (f = 1 m) lens with two cylindrical lenses (f = -25 mm, f = +100 mm).

Fig 1 Experimental setup

The airflow was seeded with olive oil particles, and the propane flow was seeded with anisole (0.015% in volume in the stoichiometric propane-air mixture). OH-PLIF images were corrected for laser sheet inhomogeneities by collecting a LIF reference image obtained from homogeneous images of a rich propane-air mixture seeded with anisole. PIV and LIF radiations were separated with a dichroic mirror, which had a cut-off wavelength of 350 nm. Fluorescence was collected for anisole in the spectral range of 275 to 350 nm with a 10 mm thick liquid filter (0.6% vol. of toluene in iso-octane) and with a combination of Schott filters (UG11 and WG280, 3 mm). The time delay between the two PIV pulses was adjusted to the velocity magnitude. The PLIF laser pulses were fired at intervals of 2 µs and positioned in the middle of the PIV time delay at intervals of 2 µs. A calibration grid was used to determine the high-order mapping functions of the three cameras (Figure 2). This calibration ensured the perfect overlap of the three measurements, especially for the OH-PLIF off-axis acquisition, which viewed the laser sheet at a slightly tilted angle. Experiments were conducted in a 60×60×160 mm 3 constant volume combustion chamber equipped with three quartz windows, allowing optical access in the visible and UV spectra. The chamber was initially filled with a lean homogeneous propane-air mixture with either ϕ=0.6 or ϕ=0.4 at atmospheric pressure. The stratified mixture field was created by injecting a laminar jet of a richer mixture into the chamber with an injection system, providing controllable and repeatable mixture stratifications. Two flame families were investigated (Table 1): Family A, where the flame was ignited in a rich stratification with a lean mixture of φch=0.6, corresponding to a mixture at the border of the flammability limit. Family B, where the mixture in the chamber was φch=0.4, below flammability limit, and flames were ignited in several richer mixtures from φ_ig=0.8 to φ_ig=1.5.

The stratified mixture was spark-ignited by energizing two thin electrodes with a diameter of 0.4 mm placed at the center of the combustion chamber. Previous experiments on homogeneous mixtures in this combustion chamber have shown that spark energy and electrode heat loss did not affect flame propagation. The limited volume of the chamber had little influence on flame propagation for the investigated time frame (with a pressure increase of less than 40 mbar), and due to the direct measurement procedure used for its determination the burning velocity measurement is non-affected by the limited chamber volume [START_REF] Balusamy | Direct measurement of local instantaneous laminar burning velocity by a new PIV algorithm[END_REF].

Results and discussion

Active flame front identification

Once the flame reached the nonflammable mixture, ϕch=0.4, it became necessary to differentiate between the burned gas boundaries and the active flame front. For the interface of fresh and burned gas, numerous studies have proposed to use OH-gradient levels to identify flame fronts [START_REF] Bayley | Curvature measurements of a lean partiallypremixed model gas turbine combustor 15th[END_REF][START_REF] Sadanandan | Simultaneous OH-PLIF and PIV measurements in a gas turbine model combustor[END_REF][START_REF] Sweeney | Autonomous extraction of optimal flame fronts in OH planar laser-induced fluorescence images[END_REF]. After obtaining the OH outline, the OH signal was extracted along its normal. This normal profile was interpolated with a complementary error function (erfc) before calculating the noise-free gradient (Figure 3).

Fig 3 example of erfc function extraction from a local

To ensure that OH-gradient levels are suitable for determining active flame fronts, OH-gradient mapping was first compared with OH×CH2O PLIF to verify that this simple diagnostic was sufficiently able to localize local flame extinctions. This preliminary experiment is performed in a Bunsen burner conical flame, where a local extinction was produced by a small air jet. In Fig. 4, a threshold value of OH-gradient is visible where the flame is extinguished. The value of this threshold value was determined for homogeneous outwardly propagating flame. Then, the OH-gradient threshold was defined (in arbitrary units) to provide the critical limit below which the flame extinguished. The OH gradient has previously been measured for homogenous outwardly propagating flames at various equivalence ratios of up to ϕ = 0.6., while in the present configuration, the flame kernel could not self-propagate below ϕ = 0.62. Figure 5 shows the probability density functions (PDF) of the OH gradient according to equivalence ratio.

Flame thickness determination

The equivalence ratio was measured based on anisole-PLIF, following a shot-to-shot laser sheet correction according to the procedure outlined by Balusamy et al. [START_REF] Balusamy | Laminar propagation of lean premixed flames ignited in stratified mixture[END_REF]. The OH-PLIF marked high temperature burned gas (T>1500 K), whereas anisole pyrolysis was noticeable around 800 K [START_REF] Nowakowska | Detailed kinetic study of anisole pyrolysis and oxidation to understand tar formation during biomass combustion and gasification[END_REF]. After calibrating the mapping functions of the three cameras, the distance between these two isotherms, determined by superpositioning the simultaneous OH-PLIF and anisole-PLIF measurements, was used to define a thickness labeled . The local distance between these two isotherms is measured all around the flame outline from the superposition of the simultaneous OH-PLIF and anisole-PLIF. It varies with the stratification (Figure 6). Consequently, the stretch rate was averaged at each time, and the flame properties were not extrapolated to a zero stretch rate, as done by Balusamy et al. [START_REF] Balusamy | Laminar propagation of lean premixed flames ignited in stratified mixture[END_REF]. Exponential modeling of (Table 2), from to the asymptotic behavior, shows that the decay time was for and for . The asymptotic behavior was attained more rapidly for because of the lower stratification amplitude ( ).

Once the asymptotic behavior was attained (t>5 ms), a slight increase in was observed for , corresponding to the decreasing stretch effect observed for increasing flame radii. For longer able to self-propagate due to overly intense stretch effects occurring with small flame radii. In the stratified case the flame succeeded to propagate despite its small radius, as confirmed by the OH gradient values (Figure 11), which were higher than the defined extinction threshold (17 000 a.u./mm). The flame was sustained by the memory effect, but was slowed down by the stretch effects. The gain provided by the memory effect decreased with time, and the influence of the stretch effect also decreased, resulting in a relatively constant burning velocity: for the investigated time range, the two effects were balanced, with a slightly faster decay observed for the memory effect.

Fig 11 OH-PLIF gradient according to time

Overall, these results show that ignition in a slightly rich mixture resulted in a flame that rapidly propagated in the lean mixture despite the high stretch rates. This velocity decreased as the flame propagated in the lean mixture, with a decay time of slightly less than 2 ms. This high burning velocity was related to the discrepancy between the local equivalence ratio in the flame front (peak heat release) and in fresh gas. The flame propagation then adopted an asymptotic behavior with a burning velocity close to the fundamental velocity. The physical properties of the burned gas (here, mainly the high temperature) compensated for the non-equidiffusive effects of the lean stretched propane-air flame, which resulted in a leaner mixture in front of the flame with a lower burned gas temperature.

These observations were made for stratifications that led to few partial oxidation products in burned gas (peak equivalence ratio of 1.2). For partial oxidation products are more abundant in burned gas. At 2 ms, just after the flame had left the stratified layer, was equal to 0.33 m/s, which is roughly the same value as for and for demonstrating that the burning velocity increased more strongly between 3 and 6 ms. This additional combustion enhancement is visible in Figure 10, in which the case represents a faster burning velocity for stretch rates in a time range of 3 to 6 ms. These results show that the back-support behavior of the flame was enhanced by ignition in a richer stratified layer: the excess in partial oxidation products in burned gas enriched the flame front for longer and/or led to a secondary reaction in the burned gas [START_REF] Haworth | Numerical Simulation of Turbulent Propane-Air Combustion with Nonhomogeneous Reactants[END_REF]. Due

to the relative change in the curve concavity , a decay time cannot be determined through exponential modeling. However, the same asymptotic behavior as for was attained after a slightly longer time (t>6 ms); this behavior was observed for all measured properties.

For the three cases of family A, the flame thickness showed the same time evolution. At , despite the lean mixture in fresh gas, the flame was relatively thin with a thickness close to the minimum value measured in near-stoichiometric homogeneous flames for the cases and the flame was slightly thicker for the case with corresponding to for a homogeneous flame (Figure 1). These smaller thickness values are consistent with high values of . Then, increased over time to attain a roughly constant value close to that measured in the homogeneous flame with and (Figure 7).

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  International Symposium on the Application of Laser and Imaging Techniques to Fluid Mechanics・LISBON | PORTUGAL ・JULY 4 -7, 2016

Fig 2

 2 Fig 2 regular grid superimposed to raw grid image (left), corrected grid image (right)

Fig 4

 4 Fig 4 simultaneous OH-gradient and CH2OxOH LIF in a Bunsen burner conical flame, where a local extinction is made by injection a small air jet.

Fig 5

 5 Fig 5 Probability density function of OH-PLIF gradient normal to the PLIF outline

Fig 6

 6 Fig 6 superposition of the simultaneous OH-PLIF and anisole-PLIF for the homogeneous laminar jet at , 4 ms after ignition (left) and for the case (right)

18thFig 7

 7 Fig 7 Thickness according to equivalence ratio: comparison of δLIF with the thermal thickness given by the plane flame calculation for propane-air, the distance between the isotherm 650K and the line of ∇[OH] max in the flame modeling

Fig 8

 8 Fig 8 Combined measurements and angular sector of the conditional analysis

Fig 9 Fig 10

 910 Fig 9 Burning velocity according to time: family A (red), family B (blue), and homogeneous flames (green) with ϕ=0.8. Extinguished flame conditions (open symbols)

,

  the minimum value of was very close to the value of , and a slight decrease was observed, in agreement with previous observations by Balusamy et al. (Balusamy, Cessou and Lecordier 2014). This behavior illustrated the competition between the stretch effects, which tended to reduce the burning rate and the back-support behavior of the flame. After igniting in the homogeneous mixture with , outwardly propagating spherical flames were no 18th International Symposium on the Application of Laser and Imaging Techniques to Fluid Mechanics・LISBON | PORTUGAL ・JULY 4 -7, 2016

18th

  International Symposium on the Application of Laser and Imaging Techniques to Fluid Mechanics・LISBON | PORTUGAL ・JULY 4 -7, 2016. Because stretch had an effect on the flame, these identical velocities possibly indicate either an identical equivalence ratio within the flame front for the three cases, or a much richer flame whose burning velocity was increased by the stretch effects. The time decay was slower and differed from the exponential decay observed for the two other cases of family A. The concavity of the curve differed from the ignition cases with ,

Fig 12

 12 Fig 12 Flame thickness according to time

Family B: propagation in a lean nonflammable mixture Flames of family B, for which the lean mixture was below the flammability limit, were investigated to determine whether this combustion enhancement also occurred in nonflammable mixtures and over which time scales. For these four flames, the OH-LIF gradient was used to differentiate the border of hot burned gas from the active flame front (Fig 11). A , the OH-LIF gradient values were much higher than the extinction threshold (17 000 a.u./mm).

Depending on the stratification amplitude, the gradient's progressive decay and the extinction occurred at different times (all of which were greater than 4 ms) and increased with the equivalence ratio at ignition. Thus, ignition in a richer stratified mixture allowed the flame to propagate in the nonflammable mixture: the flame ignited in the lean mixture ( ) as well as in the rich mixture ( ). Figure 9 shows the burning velocities over time for active reaction zones (with an OH-LIF gradient higher than 17 000 a.u/mm.). The conditions for which We would like to thank Dr Ghislain Lartigue for his help in modeling the flame and for the fruitful discussions.