Proceedings of - Sites at Penn State

1 Copyright © 2020 Solar Turbines Incorporated

Proceedings of ASME Turbo Expo 2020: Turbomachinery Technical Conference and Exhibition

GT2020 June 22-26, 2020, London, England

GT2020-16252

EFFECT OF CENTRAL PILOT FLAME ON SELF-EXCITED COMBUSTION INSTABILITY IN A VARIABLE LENGTH CAN COMBUSTOR

Daniel Doleiden1, Jihang Li1, James Blust2, Jacqueline O’Connor1

1Pennsylvania State University, University Park, PA, USA 2Solar Turbines Incorporated, San Diego, CA, USA

ABSTRACT Combustion instabilities in gas turbines are associated with

increased pollutant emissions and premature wear and failure of components. Combustion instabilities arise from a feedback loop between resonant combustor acoustics and flame heat release rate oscillations. In this study, we consider the stabilizing efficacy of a central pilot flame on a swirl-stabilized premixed main flame in a self-excited, single-nozzle combustor. In this configuration, the fuel is split between the main and pilot streams so that as the pilot flame fueling rate increases, the main flame fueling rate decreases and the overall equivalence ratio stays constant. The global equivalence ratio is fixed at 0.6, while the overall length of the combustor and percentage of main fuel diverted to the pilot flame are varied to map combustor stability across various operating points. Flame oscillations appear in three modes (modes I, III, and IV) at frequencies of 168 Hz, 365 Hz, and 220 Hz, at combustor lengths of 63.5 cm, 106.7 cm, and 149.9 cm, respectively. As pilot percentage is increased from 0% to 10%, two modes are stabilized, but the mode present at the longest combustor lengths (mode IV) is not. In addition, mode IV induces out-of-phase oscillations between the main and pilot flames, whereas pilot oscillation is not present in the other modes. This study utilizes high-speed chemiluminescence imaging and a one-dimensional acoustic model to compare the dynamic stability characteristics of mode IV to those of modes I and III and to gain insight about this mode’s response to piloting. Low-order acoustic modeling of the combustor suggests that the flame is located near a pressure minimum in modes I and III and near a pressure anti-node in mode IV. Thus, improvements in static stability may resultant from piloting are able to reduce heat release oscillations in modes I and III, but significant thermoacoustic driving remains in mode IV due to the flame’s location relative to the pressure antinode.

NOMENCLATURE φglobal Global equivalence ratio φmain Main flame equivalence ratio φpilot Pilot flame equivalence ratio

Lcomb Overall combustor length Π Percentage of fuel diverted to pilot Πair Percentage of air diverted to pilot P’ Acoustic pressure fluctuation Pmean Combustor mean pressure Q’ Heat release rate fluctuation finst Instability frequency

INTRODUCTION

Thermoacoustic instabilities in industrial gas turbine engines occur when fluctuating heat release rate and the resonant acoustic field couple, causing high-amplitude oscillations in the combustor [1],[2]. These instabilities lead to increased pollutant emissions, premature component wear, and increased potential for flame flashback or blow-off [2]. Fuel-lean operation, commonly used in gas turbines to reduce thermal NOx emissions, has been shown to increase susceptibility to thermoacoustic instability [3]. Thus, a large literature has been devoted to the understanding and mitigation of these instabilities.

Instability mitigation techniques are generally divided into active and passive categories. Active mitigation techniques aim to directly interrupt the instability cycle by modulating reactant flow on a timescale commensurate with the period of the acoustic cycle [4]. Passive methods, such as fuel staging, may apply a uniform parameter change on much longer timescales to alter the acoustic response of the combustor [5]. Other passive methods reduce the sensitivity of the reactant mixing process to acoustic oscillations via passive damping [6] or alter the acoustic response of the combustor itself using acoustic damping devices [7]. Pilot flames, small secondary flames in proximity to larger main flames, can be implemented in an active or passive fashion to control thermoacoustic instabilities [8]. Pilot flames are generally either positioned in an annular ‘ring’ fashion around a main flame or as a thin central torch inside a main flame. Piloting is utilized in designs by several manufacturers of industrial gas turbine engines [5], [9].

The presence of a pilot flame affects both the static and dynamic stability of the main flame. Near lean blow-off (LBO)


conditions, pilot flames enhance static stability by providing a consistent source of heat and radical chemical species near the base of the flame, thus promoting flame anchoring and increasing static stability [10],[11],[12]. Under conditions where the main flame is already anchored, generation of heat and radical species by the pilot flame may enhance the dynamic stability of the flame. One such example is the combined experimental and large eddy simulation (LES) work of Sengissen et al. [13], who demonstrated that piloting prevents the formation of a precessing vortex core (PVC), stabilizing the main flame. A pilot flame may also weaken flame-vortex interaction, as shown by Lee et al. [14]. It should be noted, however, that the pilot flame can be unstable, as shown by the combined experimental and computational work of Fu et al. [15]. Here, a bluff-body stabilized partially-premixed pilot flame coupled with an acoustic mode of the combustor in the absence of the main flame. Additionally, previous results from the experiment considered in this study showed that the presence of a pilot flame can have a destabilizing effect [16].

Various studies of other types of burners with a central jet have also been conducted. Gupta et al. [17] used a double-concentric premixed burner consisting of a central nozzle with two swirled outer annuli to study the effects of radial swirl distribution. They observed the formation of a thin but intense reaction zone when the annuli were counter-swirled, along with non-symmetric temperature fluctuation. Russ et al. [18] established a physical model for frequency-dependent flame dynamics of steady-state premixed flames to which experimental data from a lean premixed (LP) and lean premixed pre-vaporized (LPP) burner was supplied. This burner consisted of a swirled central pilot lance surrounded by a concentric adjustable swirled LP/LPP main flame. Meier et al. [19] used a turbulent diffusion burner for confined swirled natural gas flames. This burner enabled probability density function measurements of temperature, mixture fraction, and presence of major chemical species to be taken, with the aim of validating commonly available models.

In the current experiment, the amplitude of the instability increases, decreases, or stays the same with the addition of the pilot flame depending on the operating condition and combustor length. Previous studies by Li et al. [16] discussed a two-dimensional stability map relating acoustic pressure fluctuation to percentage of total fuel flow diverted to the pilot flame (Π) and global equivalence ratio (φglobal). A pilot fuel flow percentage sweep was conducted across one of the stability boundaries of the stability map to study the abrupt increase in acoustic pressure fluctuation associated with the boundary. Without piloting, the main flame was not anchored and heat release primarily occurred in the outer recirculation zone. Addition of piloting caused the flame to propagate upstream towards the centerbody, shifting heat release to the inner shear layer. This shift exposed the flame to strong vortices, causing instability.

The goal of the current study is to understand the efficacy of a premixed pilot flame in mitigating instabilities associated with three of the acoustic modes of the combustor, termed Modes I, III, and IV. Previous work showed that at an overall equivalence

ratio of φglobal = 0.6, Modes I and III can be fully suppressed by piloting, but Mode IV does not respond to piloting [16]. High-speed CH* chemiluminescence images are used to construct deconvoluted time-averaged images, local phase images, and Rayleigh index images to study these phenomena and characterize the effects of the pilot flame on static and dynamic stability characteristics of the main flame. Reduced-order thermoacoustic modeling using OSCILOS [20] is used to understand the impact that acoustic mode shape has on the instability mechanisms at different combustor lengths. We find that the flame of Modes I and III is situated near an acoustic pressure minimum, while the flame of Mode IV is situated near a pressure anti-node. From the combination of experimental results and modeling, we hypothesize that improvements in static stability resultant from piloting are able to reduce the level of thermoacoustic driving in Modes I and III, but a high level of driving remains in Mode IV due to the flame’s location relative to the pressure anti-node of that mode.

EXPERIMENTAL OVERVIEW AND METHODS Experimental configuration

Experiments are carried out in an atmospheric self-excited swirled lean-premixed natural gas combustor. As shown in Figure 1, the combustor consists of an inlet section, an optically-accessible quartz liner, and a variable-length metallic section. The flame configuration includes a swirl-stabilized main flame and an axial jet pilot flame. The main and pilot flames are fully-premixed and their equivalence ratios (φmain and φpilot) are controlled independently. Compressed air is metered through a Sierra 570S thermal-mass flow meter before being electrically preheated (Tin=250 °C). Fuel for the main flame is metered through a Teledyne Hastings HFM-301 flowmeter, while fuel for the pilot flame is metered through a Teledyne Hastings HFM-D-301A flowmeter. Preheated compressed air and fuel are combined before being supplied to the inlet section to provide separate reactant streams for the main and pilot flames, with pilot fuel flow rate (Π) ranging from 0-10% of the total fuel flow rate and pilot air flow rate (Πair) fixed at 5.5% of the total air flow rate.

Figure 1: The experimental apparatus showing: a) the inlet section; b) the plenum; c) the quartz combustor; and d) the

metallic variable-length combustor.

Figure 2 shows the structure of the injector. Here, the main reactant premixture flows through an annular passage containing an axial swirler before being injected into the quartz combustion chamber at the dump plane. The pilot reactant premixture is fed through a central passage and injected along the central axis of


the main flame without swirling. Previous work using this combustor [16] was conducted with a technically-premixed pilot configuration, in which the pilot fuel stream was mixed with a passively entrained air stream from the main air passage. In the current fully-premixed configuration, pilot fuel and air streams are independently supplied and controlled. Although the technically-premixed pilot configuration is more representative of piloting techniques in industrial combustors, the fully-premixed pilot configuration enhances the research potential of the combustor by providing fine control of pilot mixture over a broad range of operating conditions.

Typical reactant flow rates for the operating conditions considered in this study result in a bulk velocity of 40 m/s at the dump plane. The quartz combustor liner seen in Figure 1 and Figure 2 is 15 cm in diameter and 30.5 cm long and allows the entire structure of the main and pilot flames to be accessed optically. The quartz combustor liner is coupled to a 12.3 cm inner diameter variable-length metallic section containing a water-cooled plug. A stepper motor-controlled traverser system (Isel-Automation) allows overall combustor length between the dump plane and the water-cooled plug to be varied between 63.5-149.9 cm. The values and ranges of operating parameters relevant to this study are tabulated in Table 1.

Figure 2: Schematic of the nozzle showing the flow paths of

the pilot mixture, the swirled main mixture, and the locations of the main and pilot flames inside the quartz

combustor wall.

Table 1: Test matrix Global equivalence ratio (φglobal) 0.6 Combustor length (Lcomb) 63.5-149.9 cm Pilot fuel flowrate as percent of total fuel flowrate (Π)

0, 3, 6.5, 8, 10%

Pilot air percentage 5.5% Air inlet temperature (Tin) 250 °C

Diagnostics

Water-cooled dynamic pressure transducers (PCB) are mounted in recesses in the dump plane, air nozzle, and main fuel system. The global chemiluminescence is measured using a photomultiplier tube (PMT) fitted with a 432 nm narrow band-pass filter. Pressure and global chemiluminescence are sampled

simultaneously at a rate of 8132 Hz. The dynamic structure of the flame is evaluated via CH* chemiluminescence images captured with a Photron Fastcam SA4 camera coupled to an Invisible Vision UVi intensifier. This combination allows images of the entire structure of the main and pilot flames to be captured with a frame rate of 4000 frames per second, an exposure time of 200 μs, and a resolution of 79 pixels per inch. For each operating condition, 8 seconds of pressure and global chemiluminescence data are taken, and one second (4000 frames) of high-speed images are taken. Data analysis

The stability of the combustor is characterized quantitatively by determining the root-mean-squared (RMS) level of the acoustic pressure fluctuation at the dump plane (P’RMS) within a ±10 Hz band centered about the instability frequency for each operating condition. The combustor is considered to be unstable when the normalized pressure RMS level (P’RMS/Pmean) exceeds 0.01.

Because chemiluminescence imaging is an integrated line-of-sight technique, the raw projection images are deconvoluted to obtain revolved, radially-weighted emission images. To process the images, frequency filtering is applied on a per-pixel basis to remove noise associated with turbulent fluctuations while retaining coherent fluctuations. Background image subtraction and median filtering further reduce spatial high-frequency noise. A Hankel-Fourier operator [21] is applied to perform an inverse Abel transform, which yields an emission image representing local heat release intensity on a 2-D plane. Radial weighting applies a factor of 2πr to produce revolved images representative of CH* emission integrated over the azimuthal coordinate. Revolved images are used to compute time-averaged flame structure and local heat release rate fluctuation RMS level (Q’RMS), which contribute to evaluation of the static and dynamic stability of the flames.

The dynamic stability of the flames is further characterized using the local Rayleigh index, which quantifies the degree to which a flame drives the instability. Here, Rayleigh index is defined as:

𝑅𝑅𝑅𝑅 = � 𝑃𝑃′(𝜏𝜏)𝑅𝑅𝑅𝑅𝑅𝑅𝑅𝑅′ (𝑟𝑟, 𝑧𝑧, 𝜏𝜏)𝑑𝑑𝜏𝜏𝑇𝑇

0

where RI represents the Rayleigh index, P’ represents the instantaneous combustor pressure at the dump plane, and I’Rev represents the fluctuating component of the deconvoluted local CH* intensity.

OSCILOS [20], an open-source combustion instability simulation package, was used to investigate the acoustic characteristics of the combustor. Eigenfrequencies, instability growth rates, and acoustic pressure mode-shapes were calculated in OSCILOS by modeling the combustor as an acoustic network of constant-diameter sections. Acoustic waves are modeled as plane waves and a choked inlet condition and a closed end outlet condition (representing the moveable plug) were imposed. Crocco’s n-τ model [22] is used to simulate the response of the flame to acoustic waves. Because it was not possible to readily determine the value of n, results for two n-values are supplied to


show insensitivity to the amplification factor. The value of τ is calculated by dividing the axial distance between the center of heat release (CoHR) of the main flame and the dump plane by bulk convective velocity. Test matrix

As shown in Table 1, the combustor is operated at a fixed global equivalence ratio (φglobal) of 0.6 while the ratio of pilot fuel flowrate to total fuel flowrate (Π) is varied between 0%-10% across five increments. Doing so enables the thermal output of the combustor to be held constant for each operating condition. The overall combustor length between the dump plane and the water-cooled moveable plug is varied in 2.54 cm increments between 63.5 cm and 149.9 cm for each increment of Π. The heat release rate (Q) and equivalence ratios of the main and pilot flames as functions of Π are provided in Table 2. Here, it can be seen the thermal output of the pilot flame is relatively low compared to that of the main flame.

Table 2: Heat release rate and equivalence ratio of the main

and pilot flames at various pilot fuel percentages. Π φmain Qmain [kW] φpilot Qpilot [kW]

0% 0.600 98.25 0 0 3% 0.534 95.3 0.284 2.95

6.5% 0.515 91.9 0.614 6.39 8% 0.506 90.4 0.756 7.86

10% 0.495 88.4 0.946 9.83

RESULTS AND DISCUSSION Unpiloted dynamic stability map

At the operating conditions and combustor length range considered in this study, we observe three distinct self-excited instability modes, termed Modes I, III, IV. Mode I (finst = 168 Hz) activates at combustor lengths shorter than 76.2 cm, Mode III (finst = 366 Hz) activates at combustor lengths between 101.6-119.4 cm, and Mode IV (finst = 226 Hz) activates at combustor lengths longer than 139.7 cm. The modes are defined at the peak amplitudes, although the system is unstable (P’RMS/Pmean>0.01) at combustor lengths both higher and lower than the mode location. An additional instability mode, termed Mode II, activates at equivalence ratios not considered in this study. The dynamic stability map of the combustor in Figure 3 shows the RMS level of the acoustic pressure oscillation of the unpiloted main flame (Π = 0%) as a function of combustor length between the dump plane and moveable plug (Lcomb). Mode IV is the loudest of the modes, followed by Mode III and Mode I, which have similar instability amplitudes. These results match those from previous experiments in this facility [16],[23], indicating that these instability modes are repeatable over the course of years of testing.

Unpiloted flame dynamics

The dynamics of the unpiloted, unstable main flames are considered by examining the time-averaged flame structure, heat release rate RMS, and Rayleigh index images constructed from high-speed CH* chemiluminescence images shown in Figure 4.

As shown in Figure 4(a), the time-averaged flame brush of Mode I indicates that the main flame is anchored to the centerbody of the nozzle; the main flame impinges against the quartz liner, leading to recirculation of hot radical species in the outer recirculation zone (ORZ). The heat release rate RMS level for this mode suggests that heat release rate oscillation primarily occurs in the inner shear layer (ISL) of the confined flow, likely as a result of vortex shedding from the dump plane, though a simultaneous velocimetry study was not performed and thus direct conclusions about the flow field cannot be drawn.

Figure 3. Dynamic stability map for the unpiloted flame

showing instability Modes I, III, IV.

Figure 4. Time-averaged flame structure (top row), heat

release rate RMS (middle row), and Rayleigh index images (bottom row) for instability Modes I (a), III (b), and IV (c).

Phase-resolved images computed from high-speed CH* chemiluminescence imaging of Mode I in Figure 5a show that the main flame lingers near the dump plane for much of the oscillation period, then the flame quickly translates downstream as vortices are shed. At this end of the oscillation, the flame temporarily lifts off from its attachment point on the centerbody, causing a significant spike in heat release rate oscillation. The Rayleigh index plot of Mode I shows regions of in-phase


oscillation between local CH* emission intensity and dump plane pressure fluctuation with shades of orange; regions having an out-of-phase relationship are denoted with shades of blue. Here, almost the entire flame drives the instability, with a minor out-of-phase region associated with the base of the ORZ.

Figure 5: Phase-resolved images of of the unpiloted main

flame of Modes a) I; b) III; and c) IV.

The time-averaged flame structure in Mode III, in Figure 4b, is more compact than that of Mode I, with the center of heat release shifted upstream. The heat release rate RMS level for this mode shows that the region of significant heat release rate oscillation is concentrated near the combustor liner, with little to no oscillation occurring towards the base of the flame. Phase-resolved images computed from high-speed video of this condition (Figure 5b) show that the main flame does not lift off during the oscillation period as it does in Mode I. The longitudinal oscillation of the flame is roughly sinusoidal with time as a result of regular vortex shedding along an anchored flame. In the Rayleigh index image, a region of damping near the base of the flame develops and more of the ORZ region drives the instability instead of damping, unlike Mode I.

The time-averaged flame structure of Mode IV, in Figure 4c, is similar to that of Mode I. The structure of the heat release rate RMS is also similar except that less oscillations are seen near the wall than in Mode I. Phase-resolved images computed from high-speed video of this condition (Figure 5c) also reveal a similarity of the flame dynamics between Modes I and IV, showing lingering of the main flame near the dump plane followed by rapid displacement of the flame as vortices are shed from the dump plane and the flame temporarily blows off from the centerbody attachment point. The main flame angle also flares outward at the end of the oscillation period, perhaps indicating increased vortex strength in this mode, which would be driven with the increased amplitude of the pressure fluctuation in the combustor, as seen in Figure 3. The Rayleigh

index plot of Mode IV shows a similar structure to that of Mode I, with stronger driving near the base of the flame and extension of the driving region through the ORZ to the dump plane.

The phase between oscillations in global heat release rate (measured from the PMT) and dump plane acoustic pressure is tabulated in Table 3. Here, all phases are within 90°, which indicates by the Rayleigh criterion that thermoacoustic driving results in the instability. However, heat release rate oscillations lead acoustic pressure oscillations at the dump plane (P’comb) and at the air nozzle (P’air) in Modes I and III, while an inverse relationship is observed for Mode IV. The negative phase difference between Q’ and P’comb in Mode IV indicates that the coupling mechanism in this mode likely differs from those of Modes I and III, despite similarities in flame dynamics between Modes I and IV. A readily apparent trend is not observed for acoustic pressure in the fuel system (P’fuel); previous measurements in the fuel system indicate that fuel-system acoustic oscillations are relatively weak and so equivalence ratio coupling is not dominant at these conditions.

Table 3. Phases between heat release rate fluctuation and

pressure in the combustor, air nozzle, and main fuel circuit.

Mode Q’-P’comb Q’ – P’air Q’-P’fuel I 53° 55° -128°

III 58° 66° 226° IV -22° -17° 99°

To better understand the relationship between flame

dynamics and combustor acoustics, a one-dimensional thermoacoustic model study outlined in Experimental Setup and Methods was performed for Modes I, III, and IV. The flame was represented as a 1-D heat source whose position corresponds to the downstream location of the center of heat release of the main flame. Heat flux magnitude is determined from the reactant flow rates, lower heating value of the fuel, and an assumed 100% combustion efficiency for the fuel-limited reaction. The characteristic time delay (τ) used in the linear n-τ flame model was calculated by dividing the downstream location of the CoHR by the bulk velocity of the combustor. The results of this study are presented in Figure 5 as acoustic mode shapes, which represent the spatial pattern of pressure fluctuation amplitude. A shaded blue region denotes the air nozzle, while the solid black line denotes the dump plane. The vertical red line indicates the average location of the flame for all three cases; it only varies by a maximum of 8.8 cm between cases as denoted by the vertical red shaded region.

Due to the difficulty of estimating the gain factor, n, modeling results for two values of n are provided. These values do not impact the overall spatial characteristics of the mode shapes, only the amplitude of the pressure fluctuation. While these modeled acoustic mode shapes have not been verified with experimental measurements, close agreement between the calculated frequencies and those measured in the experiments provides some confidence in the mode shapes from this model. The experimental and modeled modal frequencies for Mode I are


168 Hz and 171.7 Hz/181.6 Hz, for Mode III are 365 Hz and 376.1 Hz/401.5 Hz, and Mode IV are 220 Hz and 212.8 Hz/225.4 Hz, respectively; all modeled frequencies are quoted for both n=2 and n=5.

Figure 6. Acoustic mode shapes of modes I, III, and IV from

1D acoustic modeling.

The pressure mode shapes of Modes I and III exhibit a single pressure minimum inside or shortly downstream of the air nozzle, while the pressure mode shapes of Mode IV exhibit two minima: one inside the air nozzle and one inside the variable-length combustor section. Thermoacoustic driving is present in all cases because the flame is located at a region of finite pressure fluctuation, but the presence of the flame at a local maximum in Mode IV indicates that the instability is likely more robust than the other two modes as the pressure fluctuation amplitude is high throughout the entire flame region. This mode shape is largely driven by the length of the combustor, which allows for more variations in acoustic pressure fluctuation amplitude than the shorter combustor sections. This difference in the shape of Mode IV is also likely responsible for the difference in the phase between pressure and heat release rate oscillations as compared to those of Modes I and III.

The hypothesis that Mode IV is more “robust,” or driven more strongly by the pressure field due to its proximity to a pressure fluctuation maximum, has been seen in work by Anderson and coworkers [24]. This work focused on experiments and modeling of longitudinal instabilities in single-element rocket injectors and the role of acoustic coupling between the combustion chamber and the injector. In this study, the authors varied the combustor length (like in the current study) to understand the impact of the geometry on the mode shape. They found that high-amplitude, high growth-rate instability occurs when the pressure anti-node is located approximately 6.4 cm downstream of the dump plane, which is also the length of the recirculation zone in the step between the injector and the combustor, where the flame is stabilized in this recirculation zone. The authors also noted that high amplitudes occur when

oscillations in the fuel injector constructively interfere to further drive oscillations at the injector face. Anderson and coworkers used an acoustic model with similar fidelity to OSCILOS, where the instability frequencies from the models and the experiments matched closely as did pressure measurements in a few locations in the injector and combustor. Results from this study indicate that pressure anti-nodes occurring in the region of high heat release rate oscillations from the flame can drive high-amplitude, low damping-rate instabilities, as is seen in Mode IV in the current combustor.

Suppression efficacy of piloting

A fully-premixed, non-swirled axial pilot flame was used to stabilize the main flame with pilot fuel flowrate percentages ranging between 0-10% of the total fuel flowrate. The pilot air flowrate was held constant at 5.5% of total air flowrate for all operating points. In a similar fashion to Figure 3, the dynamic stability map in Figure 6 shows dump plane acoustic pressure fluctuation RMS level as a function of combustor length. Each colored line corresponds to an individual pilot percentage. Without piloting (Π = 0%), Modes I, III, and IV are visible as in Figure 3. As Π is increased, Modes I and III stabilize fully. Mode IV does not stabilize even with heavy piloting (Π = 10%), although there is some reduction in the instability amplitude with piloting. The inability of high levels of piloting to completely stabilize Mode IV, as well as the associated flame dynamics, are the focus of the remainder of the investigation.

Figure 7. Dynamic stability map at a range of piloting

levels, Π=0-10%.

Figure 8 presents time-averaged flame structure and Rayleigh index plots for all operating conditions at various levels of piloting. These images are arranged into columns corresponding to Mode I, III, and IV in descending order of Π. The time-averaged main-flame structure of Mode I narrows and elongates as pilot fuel diversion reduces the main flame equivalence ratio and thus decreases the flame speed and heat output of the main flame.


Figure 8: Time-averaged flame structure (top) and Rayleigh index images (bottom) for Modes I (a), III (b), IV (c). Units of the colorbars are image counts of the intensified camera.


The CoHR moves towards the combustor liner as the flame stabilizes in a region of lower local velocity to satisfy the kinematic condition. Flame brush thickness decreases as piloting stabilizes the main flame, reducing the flame edge oscillations. At Π = 8%, the structure of the pilot flame becomes evident. The pilot flame is a narrow jet flame with a high aspect ratio and a region of concentrated heat release extending from the dump plane to the middle of the main flame brush.

The Rayleigh index plots for this mode show that the driving region of the main flame exhibits relatively similar structural changes to the flame brush and that damping regions develop along the base of the main flame and near the ORZ as Π increases. The Rayleigh index of the pilot flame indicates that it does not oscillate significantly in this mode, with only light damping evident at Π = 10%. These observations suggest that the pilot flame indirectly improves the dynamic stability of the main flame by reducing lift-off, thus altering the dynamics of the main flame, rather than significantly participating in the dynamics of the flame oscillations. The time-averaged flame structure of Mode III responds very similarly to changes in piloting as that of Mode I, with a similar main flame brush at Π = 8%. The pronounced damping region near the base of the main flame suggests the presence of convective pockets consistent with the aforementioned hypothesis of vortex shedding as an instability mechanism in this mode. The pilot flame develops a split damping-driving structure between Π = 6.5-10%. High-speed video shows that this manifests as periodic out-of-phase displacement between the base of the flame and the flame tip. This phase-cancellation between different parts of the flame as a mechanism of instability suppression has been seen in flame transfer function characterization of this combustor main flame, with simultaneous out-of-phase oscillation of the inner shear layer [23] as well as other studies of swirl-stabilized flames [25].

In contrast to Modes I and III, the flame brush of Mode IV is not significantly altered by piloting, as expected since the main flame is not stabilized by piloting in this mode. The Rayleigh index plots of this mode show that the entire main flame drives the instability for all values of Π. Interestingly, the pilot flame is seen to drive the instability at Π = 8% and damp the instability at Π = 10%, suggesting a shift in phase offset occurs between these pilot percentages. Phase-resolved images computed from high-speed video, shown in Figure 9, indicate that the entire pilot flame oscillates out-of-phase with the main flame, with the highest degree of oscillation occurring toward the tip of the pilot flame. This out-of-phase behavior is likely also the result of the difference in acoustic mode shape in Mode IV as compared to Modes I and III. The flow rate through the pilot injector is likely driven by the pressure oscillations in the combustor, which, according to the mode shapes in Figure 5, are out-of-phase with the oscillations in the injector. The clear gap in CH* emission intensity between the main and pilot flames in the time-averaged flame structure images of Figure 8 show that direct flame-flame interaction does not occur. Thus, direct flame-flame interaction does not appear to contribute to the suppression efficacy of piloting.

Figure 9: Phase-resolved images of the main and pilot

flames in Mode IV (Π = 10%).

Figure 10 shows the flame fluctuation level in both the main and pilot flames at different piloting levels. The normalized Q’RMS of the pilot is generally low compared to that of the main flame, indicating that the pilot flame does not have a significant dynamical contribution to the overall heat release rate oscillation. The pilot flame doesn’t oscillate significantly except at Π = 10%, in which the normalized Q’RMS level of the pilot flame is considerably greater for Mode IV than for Modes I or III. This observation again differentiates Mode IV from Modes I and III in that the pilot is not only ineffective at stabilizing the main flame, it also begins to destabilize as a result of the system instability.

Figure 10. Normalized heat release rate oscillation

magnitudes for main and pilot flames as a function of pilot.

Suppression mechanisms of piloting We made several comparisons to understand the mechanism

by which piloting suppresses instability in Modes I and III but not in Mode IV. First, the time-averaged flame structures of stable unpiloted main flames were compared to determine the sensitivity of the main flame brush to changes in combustor flow and pressure field associated with variations in combustor length. This comparison determines whether combustor length changes main-flame structure and hence the ability for the pilot to suppress the flame. If the flame structure varies significantly with length, then this may be a reason why the pilot is unable to suppress the instability in Mode IV. The time-averaged flame


structures of Figure 9a and Figure 9b show nearly identical stable unpiloted main-flame structure in the combustor length intervals between Modes I - III and III – IV (Lcomb = 91.4 cm and 132.1 cm, respectively). This suggests that difference in combustor pressure distribution associated with overall combustor length does not alter the main flame structure and that this is not the reason for differences in pilot efficacy in Mode IV.

Figure 11: Time-averaged stable main-flame structure at combustor lengths intermediate to (a) Modes I-III and (b) Modes III – IV; time-averaged main-flame structure for piloted and unpiloted main flames of similar main-flame

equivalence ratio (фmain) in (c,d) Mode I, (e,f) Mode III, and (g,h) Mode IV.

To determine the effects of the pilot flame on the macrostructure of the main flame, comparisons of flame structure were made for piloted and unpiloted main flames of Modes I, III, and IV with identical main flame equivalence ratios (φmain). This comparison illustrates whether the pilot changes the flame shape or location, which could also change its thermoacoustic coupling potential. For φglobal = 0.6 with Π = 10%, φmain ≃ 0.58. At the same combustor length as Mode I (63.5 cm) but without a pilot flame, the φmain = 0.58 main flame is lifted, as shown in Figure 9c. The pilot flame causes the main flame to anchor by providing a source of heat and radical species near the base of the main flame (Figure 9d). At the same combustor length as Mode III (106.7 cm), the unpiloted φmain = 0.58 main flame is attached, as shown in Figure 9e; piloting causes the flame to more strongly attach to the centerbody, as evidenced by the stronger chemiluminescence signal near the base of the flame in Figure 9f. Thus, the presence of the pilot flame is mainly responsible for altering the static stability of Modes I and III, which in turn likely has a significant effect on the dynamic stability. Similarly, in Mode IV, the pilot flame enhances flame stabilization at the centerbody, as seen by the higher chemiluminescence signal in Figure 9h.

Putting together the results of the imaging study as well as the 1D modeling, we hypothesize that the difference in piloting efficacy is a result of the location of the flame relative to the pressure mode shape in Mode IV as compared to that of Modes I and III. According to the acoustic modeling, the flame is located near a pressure minimum, though not a node, in Modes I and III, meaning that the thermoacoustic driving level is not as high as in Mode IV, where the flame is located near a pressure anti-node. Improvements to the flame anchoring by the pilot in Modes I and III likely reduce the level of heat release rate oscillations for a given acoustic input, resulting in lower thermoacoustic driving and a suppression of the instability. These results are similar to those from Miller et al. [24] . While the pilot does help to improve the static stability of the flame in Mode IV, the thermoacoustic driving is still relatively high due to the flame’s location in the acoustic mode and cannot be as easily suppressed.

Evidence for the pilot’s mechanism – suppression of instability through reduction of heat release rate oscillations by improvement of flame anchoring – was also seen in measurements of the flame transfer function in previous work on this facility by Li et al. [23]. Here, it was shown that heat release rate fluctuation in the inner shear layer had the highest impact on the gain of the flame transfer function and that piloting served to reduce the size of the positive heat release index regions within the ISL while simultaneously generating regions of negative heat release index. It was also hypothesized that the pilot changed the static stability characteristics of the flame, which lead to the difference in flame transfer function and hence instability in the self-excited case. CONCLUSION

The instability suppression efficacy of an axial, fully-premixed pilot flame on a swirl-stabilized, premixed main flame was studied in an optically-accessible, variable length natural gas combustor equipped with a moveable end plug. Main flame and pilot flame equivalence ratios are independently controlled and the ratio of pilot fuel flowrate to overall fuel flow rate is varied between 0-10%. Three distinct acoustic modes, termed Modes I, III, IV, occur as the overall length between the combustor dump plane and end plug is varied. Stabilization of Modes I and III occurs at Π > 6.5%. Mode IV is not stabilized by piloting, although dump plane acoustic pressure fluctuation RMS levels do decrease as piloting percentage increases. Additionally, this mode exhibits a negative phase relationship between heat release rate fluctuation and acoustic pressure fluctuation at the dump plane, whereas Modes I and III exhibit a positive phase difference. This suggests that the flame dynamics of Mode IV may differ from those of the other modes.

The static and dynamic stability of the main and pilot flames were investigated by studying images of time-averaged CH* intensity, CH* intensity fluctuation RMS level, and Rayleigh index images, all of which were constructed from revolved Abel-transformed high-speed image sets. These images show that piloting enhances static stability in Modes I and III, likely by providing a source of heat and radicals near the base of the main


flame. This may also influence the dynamic stability of the main flame.

To investigate the acoustic properties of the combustor, a low-order thermoacoustic simulation study was conducted. The pressure mode shapes generated by this study show the formation of a single pressure minimum in Modes I and III in the region of the flame, whereas Mode IV has two pressure minima where the flame is located near the pressure maximum in between. The presence of the flame near a maximum in Mode IV makes suppression more difficult; while the pilot is able to reduce the amplitude of the instability, it cannot fully suppress it. Future acoustic measurements should be made to confirm this hypothesis about the mode shape, but agreement between the modeled instability frequencies and the measured ones provides some confidence in the hypothesis.

The results of this study have implications for design of combustor piloting systems in the context of a given acoustic mode shape. In particular, the main mechanism by which the pilot suppresses instability seems to be through changes to the flame static stability, which reduces the thermoacoustic driving of the self-excited system. However, pilot flame efficacy varies based on the location of the flame relative to the acoustic mode, and so the same level of pilot fuel flow rate will not suppress instability for all tones. Further investigation of this mechanism will be conducted using high-speed OH-PLIF in order to investigate the detailed flame dynamics in the flame-stabilization region with and without pilot flames at these three modes. ACKNOWLEDGEMENTS The authors are grateful for the financial support provided by Solar Turbines Incorporated with program monitor Dave Voss.

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