Exhaust CO emissions of a laminar premixed propane–air flame interacting with cold gas jets

Combustion and Flame 210 (2019) 374–388

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Exhaust CO emissions of a laminar premixed propane–air flame interacting with cold gas jets Jacob E. Rivera∗, Robert L. Gordon, Davy Brouzet, Mohsen Talei Department of Mechanical Engineering, The University of Melbourne, Parkville 3010, Australia

a r t i c l e

i n f o

Article history: Received 9 July 2019 Revised 2 September 2019 Accepted 3 September 2019

Keywords: Gas turbine exhaust emissions Carbon monoxide Effusion cooling Flame–wall interaction Flame-cooling-air interaction

a b s t r a c t This study investigates a laminar premixed flame interacting with cold gas jets, at different cooling jet mass flow fractions (m˙ ∗jet ) and diluent types, namely air and N2 . A novel burner and wall configuration is used to experimentally induce flame-cooling-air interaction (FCAI). Flame chemiluminescence imaging, exhaust temperature (Texh ) and exhaust CO emissions ([CO]exh ) measurements are conducted to characterise the flame shape and [CO]exh response to the cooling jets. Flame imaging reveals that the cooling jets greatly affect the flame shape. Measurements of [CO]exh demonstrate a direct correlation with Texh , as decreasing Texh is observed to occur with decreasing [CO]exh . Additionally, the air diluent case shows consistently lower [CO]exh values, relative to the N2 diluent case. Using a novel modelling approach, the cooling jets are simulated using one-dimensional (1D) fully resolved simulations (FRS). The effect of jet dilution, jet cooling and exhaust gas cooling are independently and jointly investigated in these simulations. The FRS results support the experimentally observed behaviour, and show that exhaust gas cooling and exhaust gas oxygenation produce decreased CO concentrations. Using a chemical reactor network (CRN), the jet mixing process is modelled by a perfectly stirred reactor (PSR), while the exhaust gas cooling process is modelled by a plug flow reactor (PFR). The CRN modelling shows that the jet mass flow rates dictated by m˙ ∗jet , the dilution time (tdil ) assumed for cooling jet mixing, and the exhaust gas cooling residence time (tcool ), play an important role in determining the [CO]exh . An equilibrium analysis illustrates that the relationship between [CO]exh , Texh and exhaust O2 , is due to the thermodynamically favoured equilibrium states. Timescale analyses demonstrate that appropriate modelling of jet mixing, and accounting for the rate of exhaust gas cooling, are important for estimations of [CO]exh . © 2019 The Combustion Institute. Published by Elsevier Inc. All rights reserved.

1. Introduction Dry low emission gas turbines typically feature increased power densities to achieve higher thermal efficiencies for reduced fuel consumption [1]. This increases the likelihood of the flame interacting with the combustor liner, which can have potential consequences on engine exhaust emissions. While there have been some studies on pure flame-wall interaction (FWI) of back-side cooled walls [2,3], relatively few studies have explored how the flame is affected by effusion cooling [4], a common combustor liner cooling method. In effusion-cooled walls, holes in the combustor liner allow the cooling air to form a protective layer between the liner and the flame, including the hot exhaust [5]. The mass flow rate of this ∗

Corresponding author. E-mail addresses: [email protected] (J.E. Rivera), [email protected] (R.L. Gordon), [email protected] (D. Brouzet), [email protected] (M. Talei).

cooling air has been previously correlated with gas turbine exhaust CO emissions ([CO]exh ) [6]. However, it is unclear whether the measured [CO]exh comes from flame-cooling-air interaction (FCAI), or diversion of primary combustor air. This paper aims to address this by focusing on FCAI. Due to the difficulty of experimentally and numerically investigating FCAI, few studies have been done regarding this phenomenon. Early effusion cooling studies used turbulent boundary layer models to explore their cooling potential [7,8]. Subsequent studies have focused on heat transfer effectiveness [9], with few investigating how the jets interact with the flame. Bizzari et al. [10] studied the evolution of the jets in a wind-tunnel type configuration, and demonstrated the difficulty involved in simulating the complex flow field induced by this phenomenon. Moreover, the addition of swirl [11] and combustion [12] into the problem further increases the challenges associated with studying this phenomenon. One study has used a single-sector model gas turbine combustor, with an effusion-cooled wall, to experimentally characterise

https://doi.org/10.1016/j.combustflame.2019.09.001 0010-2180/© 2019 The Combustion Institute. Published by Elsevier Inc. All rights reserved.

J.E. Rivera, R.L. Gordon and D. Brouzet et al. / Combustion and Flame 210 (2019) 374–388

the flow field under different operating conditions [13]. It was illustrated that, under specific operating conditions, heavy impingement of the flame onto the effusion-cooled wall occurs, leading to local penetration of the cooling jets into the flame brush. As shown by Yahagi and Makino [14], this can lead to local extinction of the flame around the jet. Flame extinction can potentially lead to incomplete combustion, and could have consequences for [CO]exh [15]. However to the authors’ knowledge, no study has attempted to directly link this behaviour with [CO]exh . Understanding this link would aid in building models to mitigate the potential impact of FCAI on emissions. Therefore, this work investigates flames interacting with cold gas jets. Flame chemiluminescence (FCL) imaging was conducted to qualitatively determine the response of the flame shape to increasing jet mass flow rates (m˙ jet ) and diluent types, namely air and N2 . Exhaust temperature (Texh ) and exhaust CO emissions ([CO]exh ) measurements were then conducted to determine the individual CO response to the cooling and diluting effect of the jets. The fundamental mechanisms that affect the observed behaviour was further investigated using one-dimensional (1D) fully resolved simulations (FRS) of the flame under the influence of the cooling jets and wall heat transfer. Source terms were added to the Navier–Stokes equations solved by the FRS, to individually induce the cooling and dilution effect of the jets, as well as the exhaust gas cooling occurring due to wall heat transfer. The FRS results were compared to the experimentally observed behaviour to confirm the direct contribution of the cooling jets and wall heat transfer on CO oxidation, and to reveal the primary controlling parameters. The conclusions of the study were extended to greater practical relevance by exploring the impact of residence time on CO oxidation under these conditions. This was conducted using chemical reactor networks (CRN), which provided scope to investigate residence times relevant to gas turbine combustors. These observations provide a step towards understanding the mechanisms governing the impact of FCAI on [CO]exh .

375

Fig. 1. Schematic cross section view of the test section, with blue solid arrows highlighting the major features. The flow directions are indicated in black dashed arrows, while the indicative flame position is shown in red lines. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

2. Experimental methods 2.1. Burner configuration The burner used in these experiments is the same as that used by Rivera et al. [16] to study FWI, except for the jet cooled wall section, and is hence only briefly discussed here. The burner is a laminar fully premixed propane–air burner, with the test section shown in Fig. 1. The reactants are metered by two MKS thermal flow controllers, M1559A and M100B for air and fuel, respectively, and are mixed 3 m upstream of the plenum. The plenum contains several layers of steel meshes, a honeycomb flow straighter and a contracting nozzle leading to the test section [17]. The test section is enclosed in a fused silica tube to reject outside air disturbances while still allowing optical access [18]. The support cap on top of the fused silica tube integrates the cooling jet section, as well as the emissions probes, and is mounted on three adjustable rails that can control the height of the jets relative to the nozzle outlet. A pilot flame is used to light the flame, and is extinguished once the flame has stabilised. Fig. 2a shows the jet section. The jet section is the main difference between this and the configuration used in a previous purely FWI study [16]. It consists of three centrally mounted concentric stainless steel tubes (Fig. 2b), with the two outer-most tubes delivering the cooling water in and out. This allows cooling of the assembly to be decoupled from the effect of the jet exit velocity, which becomes necessarily high and turbulent otherwise. The inner-most tube delivers the cooling jets, with either air or N2 as the diluent, to the jet outlet (Fig. 2c). The jet outlet section is made of 3D printed Inconel, and is welded to the upper (Fig. 2b) and

Fig. 2. Schematic cross section view of the jet section (a), as well as a top-down section view of the top (b), middle (c) and bottom (d) of the jet section. The cooling water path is indicated by black arrows, while the cooling jet path is indicated with red arrows. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

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lower (Fig. 2d) stainless steel tubes. In the middle of the jet outlet shown in Fig. 2c, coolant water paths turn from annulus-shaped to slot-shaped to allow for the four axisymmetric 3 mm diameter jet outlets. An image of the jet outlets in the test section is shown in Fig. 3, highlighting the jet exit and jet exit plane. Below the jet outlets, Fig. 2d shows the water paths returning to their annular shape, as the water then drains through the outer-most tube. The coordinate system used in this study is shown in Fig. 3, and is cylindrical in nature. The origin is set at the nozzle exit, with θ oriented out of the page, r horizontally left and x vertically up. The location of the tip of the flame (xflame ) is approximately 30 mm away from the nozzle exit, therefore x f lame = 30 mm. Distances relative to the flame tip is denoted by x∗ , where x∗ = x − x f lame . The equivalence ratio (φ ) used for this propane–air flame was 0.85, using a bulk inlet mass flow rate (m˙ in ) of 0.83 g/s through the nozzle. This corresponds to a mean unburnt gas velocity of u = 1.6 m/s, estimated based on component gas densities at 298 K and 1 atm. A hydraulic Reynolds number of ReD = 825 was calculated based on the inlet velocity and the annulus radius of 7.74 mm between the nozzle and the tube. The jet mass flow rates (m˙ jet ) used for this experiment range from m˙ jet = 0 to 0.024 g/s. This is characterised using the jet mass flow fraction (m˙ ∗jet ), defined as:

m˙ ∗jet =

m˙ jet m˙ jet = , m˙ in + m˙ jet m˙ total

(1)

where m˙ total is the total mass flow rate. The water cooling rate (Q˙ c ) used for this experiment is ≈ 150 W, set by controlling the cooling water mass flow rate (m˙ c = 1.5 g/s), and measuring the water temperatures (Tc ) at the coolant inlet (Tc,in = 298 K) and outlet (Tc,out = 328 K). The experimental residence time (τ exp ) was estimated to be approximately 250 ms. This was conducted by constructing a 1D model of the experimental domain, and accounting for the velocity changes due to changes in area, changes in density across the flame front and changes in density due to exhaust gas cooling. The estimated τ exp will be used as a basis for connecting the numerical results with the experimentally observed behaviour. 2.2. Flame imaging A Photron Fastcam SA1.1, with a Sigma 150 mm f/2.8 VIS macro lens, was used for image acquisition. The camera operated

at a repetition rate of 60 Hz, with an exposure time of 16 s and a full frame resolution of 1024 × 1024 px. Darkfield and lightfield corrections were used to account for the dark noise and non-uniform pixel response of the camera sensor. Bicubic dewarping to account for the fused silica tube was conducted [19], based on a chequerboard target, which also gave pixel distance of 67.4 m. For each condition tested, the camera was set to capture 200 images which showed flame surface unsteadiness of approximately ± 0.11 mm. This was calculated from the root-mean-square-error of the images, and allows ensembleaveraging for the qualitative comparisons presented in this study. A sample of an ensemble-averaged image of the flame is shown in Fig. 3.

2.3. CO emissions measurement A Thermofisher Model 48c was used to analyse the sample gas. This uses the Non-Dispersive Infrared (NDIR) technique to measure dry CO concentration. Two internally water-cooled emission probes were used to capture the sample gas. The first is termed the “end-gas” (EG) probe, while the second is termed the “semi-local” (SL) probe, with measurements from these two probes denoted as [CO]EG and [CO]SL , respectively. The locations for these two probes were xEG = 240 mm and xSL = 70 mm for the EG and SL probes, respectively. Relative to xflame , these correspond to x∗EG = 210 mm and x∗SL = 40 mm. The SL probe was attached to the burner via a movable plate, which allowed measurements to be conducted at a variety of radial locations, from r = 6.4 to 21.4 mm. Additionally, the cooling jet section could be rotated to allow the SL probe to measure at two azimuthal locations (θ SL ), one directly above and one in-between two jets, corresponding to θ SL ≈ 0◦ and 45◦ , respectively. Due to the different concentration ranges of the EG and SL measurements, two calibration gases were used. For the EG measurements 100.6 ppm CO − N2 was used, while 10 0 0 ppm CO − N2 was for the SL measurements. This maximises the measurement sensitivity across the different ranges. The calibration was conducted using the analyser itself, which linearises over the measurement range resulting in an uncertainty of ± 0.1 ppm. The Texh was measured using thermocouples attached to the tip of each probe, and was radiation-corrected based on the method outlined in Markides [20]. For the EG probe, TEG = 430 K, while for the SL probe TSL varies with radial location.

Fig. 3. Cross-section (a) and top-down (b) schematics of the test section coordinate system, including a sample flame image at a high jet flow rate. Highlighted are the jet exit, jet exit plane, and characteristic secondary flamelets produced by the jet. Note that the contrast has been significantly enhanced to highlight features of the flame. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

J.E. Rivera, R.L. Gordon and D. Brouzet et al. / Combustion and Flame 210 (2019) 374–388

where λ is the gas-phase thermal conductivity, Cw is the heat transfer coefficient, T is the local gas temperature and Tw is the wall temperature.

3. Numerical methods 3.1. Governing equations To separate some of the physical processes occurring in the experiment, fully resolved simulations (FRS) were conducted in 1D, to allow the simulation of a larger parameter space using detailed chemical kinetic models. The code used to conduct the 1D FRS in this study is NTMIX-CHEMKIN, an accurate high-order fully parallelised flow solver designed to perform direct numerical simulations (DNS) of flames with detailed chemical kinetic models [21,22]. This code has been previously used in DNS studies involving FWI [23]. The code solves the fully compressible Navier–Stokes mass, momentum, energy and species conservation equations. It uses an 8th order explicit central differencing scheme and a 3rd order Runge–Kutta integrator for spatial and time derivatives, respectively [24]. Species production and molecular transport terms are solved using the CHEMKIN and TRANSPORT packages [25]. Diffusion velocities are modelled using mixture-based species-specific diffusivities, calculated from temperature-based binary diffusivities [26]. Additionally, the code accounts for the Soret and Dufour effects. To simulate the effects of the cooling jet and exhaust gas cooling, source terms were added into the governing equations. The full conservation equations that are solved in 1D are as follows:

∂ρ ∂ρ u ∂ m˙ jet + = , ∂t ∂x ∂x ∂ρ u ∂ρ uu ∂ p ∂τ + =− + , ∂t ∂x ∂x ∂x ∂ q˙ jet ∂ρ et ∂ (ρ et + p)u ∂ (uτ ) ∂ q + = − − q˙ w + , ∂t ∂x ∂x ∂x ∂x ∂ m˙ jet,α ∂ρYα ∂ρYα u ∂ρYα Vα + =− + Wα ω˙ α + , ∂t ∂x ∂x ∂x

(2)

3.3. 1D cooling jet model The mass, species and energy source terms of the 1D cooling jet model correspond to the m˙ jet , m˙ jet,α and q˙ jet terms, respectively, and are based on a Gaussian function. The source term for the mass conservation equation can be expressed as follows:

∂ m˙ jet = Am f g ( x ), ∂x

(3)

(4)

(5)

3.2. Exhaust cooling model The exhaust gas cooling present in the experiment is due to wall heat transfer along the quenching wall. This was simulated using the model developed by Kosaka et al. [27], and is defined as:

(6)

(7)

where Am is the amplitude of the Gaussian function for m˙ jet , with fg (x) defining the Gaussian function as follows:



fg (x ) = exp



− ( x − B )2 , 2C 2

(8)

where B is the coordinate of the centre of fg (x), and C determines the width of fg (x). The m˙ jet source term is defined based on Eq. (1) as follows:



m˙ jet = m˙ in



m˙ ∗jet 1 − m˙ ∗jet

,

(9)

based on m˙ in and m˙ ∗jet . The constant Am can be obtained by equating the integral of Eqs. (7)–(9), giving the definition for Am as follows:



where u refers to the fluid velocity while Yα refers to the mass fraction of species α . The total energy density per unit mass is denoted by et . The fluid mass density, thermodynamic pressure, viscous stress term and the heat flux term are, respectively, shown by ρ , p, τ and q. The rate of production, molecular weight and diffusion velocity of species α are represented by ω˙ α , Wα and Vα , respectively. The variables x and t represent the spatial and temporal coordinates, respectively. The terms in Eqs. (2)–(5), denoted by the subscript jet and w, are the source terms added to replicate the cooling jets and wall heat transfer, respectively. Specifically, the m˙ jet is the mass flux due to the cooling jet, q˙ jet is the energy flux due to the jet temperature (Tjet ) and m˙ jet,α describes the species composition of the jet. No source term was added into the momentum equation as the jet was injected at the same velocity as the bulk flow. Finally, the source term q˙ w replicates the effective energy loss due to wall heat transfer, resulting in exhaust gas cooling. These source terms were implemented into the governing equations and were verified using analytical solutions of hot air flowing over a 1D flat plate. The verification is detailed in the Supplementary material.

q˙ w = λCw (T − Tw ),

377

Am = m˙ in

m˙ ∗jet 1−

 

m˙ ∗jet

L 0

fg (x )dx ,

(10)

The source term for the species conservation equation can be expressed as follows:

∂ m˙ jet,α = AmY jet,α fg (x ), ∂x

(11)

where Yjet,α is the mass fraction of species α in the jet. This keeps Yα consistent with the species conservation equation shown in Eq. (5). The source term for the energy conservation equation can be expressed as follows:

∂ q˙ jet = Aq f g ( x ), ∂x

(12)

where Aq is the amplitude of the Gaussian function for q˙ jet , and defines the magnitude of heat transfer between the local gas and the cooling jet. The Aq term can be found by constructing an integral enthalpy flow rate (H˙ ) balance over the whole domain as follows:

H˙ total = H˙ in + H˙ jet = m˙ in hin + m˙ jet h jet = m˙ total htotal ,

(13)

where total, in and jet correspond to the total, inlet and jet mass flow rates (m˙ ) and specific enthalpies (h), respectively. The integral of Eq. (12) can be equated to H˙ jet as follows:

 H˙ jet = Aq

L 0



fg (x )dx = Am

 0

L



fg (x )dx h jet ,

(14)

where fg (x) is a constant given user-defined coefficients B and C. Therefore simplifying this expression produces the definition of Aq as follows:

Aq = Am h jet ,

(15)

where hjet is evaluated differently depending on the cases considered. This is because the effect of the cooling jet is two-fold. The first effect is due to the addition of O2 and N2 into the bulk flow, changing the chemical composition of the gas. The second is the

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reduction in temperature due to the cold jet temperature. To separate these two effects and isolate their individual impacts, two cases are considered. The first is where jet dilution only (DO) is considered, and the second is where both jet dilution and cooling (DC) is considered. Case A: Jet Dilution and Cooling (DC) In this case, both dilution and cooling due to the cooling jet is considered. To achieve this, hjet is evaluated at a defined temperature which is lower than the local gas temperature. Thus for this case, hjet is evaluated as follows:

Aq = Am h jet (T jet ),

(16)

where Tjet is the chosen jet temperature. Case B: Jet Dilution Only (DO) In this case, only dilution due to the cooling jet is considered, without any thermal effects. To achieve this, hjet is evaluated at the local gas temperature, which is 300 K in the unburnt gas region or the adiabatic flame temperature (TAd ) in the burnt gas region. Thus for this case, hjet is evaluated as follows:

Aq = Am h jet (T ),

(17)

where T is the local gas temperature. 3.4. Numerical configuration The reacting simulations were conducted using C3 H8 as the fuel at φ = 0.85, with the USC C1 − C3 chemical kinetic mechanism [28], containing 70 species and 462 reactions. The domain length was set at L = 240 mm, to mimic the distance of the EG probe from the nozzle, which is also at xEG = 240 mm. A grid resolution of 16 μm (or 26 points per flame thermal thickness) was used across the flame from the inlet at x = 0 mm to x = 60 mm, stretching linearly to 500 μ m from there to the outlet at x = L. A nonreflecting outlet was imposed based on the NSCBC condition [29], while species, temperature and velocities were imposed at the inlet. The unburnt gas temperature was set at 300 K, with the pressure set at 1 atm and velocity set equal to the laminar flame speed of 0.32 m/s. The flame was initialised using a freely propagating flame (FPF) profile generated using the Chemkin package [30] and the same chemical kinetic mechanism. 4. Laminar flame-cooling-air interaction In this section, flame imaging was conducted to determine the response of the flame shape to the cooling jets, under increasing m˙ jet and to different diluent types. Measurements of Texh and [CO]exh were then conducted under these conditions to determine how they respond to the cooling effect of the jets, as well as to the dilution they impose on the flow.

Table 1 Summary of experimental cases presented. Case

x∗jet (mm)

Diluent

m˙ ∗jet (%)

Q˙ c (W)

UJD AJD DJD

−5 0 5

Air/N2 Air/N2 Air/N2

0–2.5 0–2.5 0–2.5

150 150 150

The experimental results presented in this section consists of several cases, summarised in Table 1. The three jet configurations consists of upstream jet dilution (UJD), at-flame jet dilution (AJD) and downstream jet dilution (DJD), corresponding to x∗jet < 0, x∗jet = 0 and x∗jet > 0, respectively. These three can be further split into two cases, one where air is used as the diluent and another where N2 is used as the diluent. For each of these cases, a range of m˙ ∗jet are used with a constant Q˙ c . 4.1. Cooling jet effects on flame shape Fig. 4 shows the effect of increasing m˙ ∗jet on the flame shape, for < 0. Fig. 4a shows that at m˙ ∗jet = 0%, the flame is an A-shaped flame subjected to pure FWI, as seen in Rivera et al. [16]. As m˙ ∗jet is increased to 0.56%, the tip of the flame starts to be affected by the cooling jet, where a secondary flamelet forms at the flame tip. Further increasing m˙ ∗jet to 0.88% shifts the orientation of the secondary flamelets radially outwards, as highlighted in Fig. 4b. At the highest m˙ ∗jet used (Fig. 4c), the secondary flamelets are oriented even more bulk flow-normally. Observing the jet outlet oriented at θ = 0◦ in Fig. 4c, the location of the secondary flamelet forms a region of low FCL signal, surrounded by a regions of moderate to high FCL signal. This resembles the locally extinguishing flames observed by Yahagi et al. [31]. The low FCL signal could be due to either low FCL species concentrations, or due to low local gas temperatures. The flame shape observed for the x∗jet = 0 case (not shown here) is similar to what was observed for the x∗jet < 0 case. However, since the jet is at the tip, the only secondary flamelet shapes observed resembles Fig. 4b and c. For x∗jet > 0, the shape of the flame remains the same as m˙ ∗jet = 0% for all values of m˙ ∗jet , as expected since the jets are above the flame tip. To analyse the effect of different diluent types on the flame shape, similar to Fig. 4. These splines were traced manually through the middle of the edges of the secondary flamelets allowing qualitative comparison of the flamelet sizes. The splines were traced for the air and N2 diluent cases, for m˙ ∗jet = 0.67% to 2.79%, where the secondary flamelets were visible. These are shown in Fig. 5 for the upstream jet dilution configuration, for both the air x∗jet

Fig. 4. Ensemble-averaged flame images at varying m˙ ∗jet . All results shown are for the N2 diluent case, at x∗jet = −5 mm. Note that the contrast has been significantly enhanced to highlight features of the flame. Secondary flamelets are traced in red and indicated by arrows. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

J.E. Rivera, R.L. Gordon and D. Brouzet et al. / Combustion and Flame 210 (2019) 374–388

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Fig. 5. Comparison of the left secondary flamelets generated by air (a) and N2 (b) cooling jets, at m˙ ∗jet ∈ [0.67%, 2.79%], for x∗jet < 0. Curves estimated using traced splines of the middle of the secondary flamelets. Note that the vertical axis has been shifted to the jet exit plane to aid comparison.

and N2 diluent cases. Comparing Fig. 5a to b, it can be observed that the splines for the N2 diluent case are longer in general than those for the air diluent case, for the same m˙ jet . This could suggest that the N2 diluent case results in lower local flame speeds in the vicinity of the secondary flamelet, relative to the air diluent case. This is expected behaviour since longer flame lengths are generally associated with lower flame speeds [32]. 4.2. Exhaust CO emissions response to cooling jets Fig. 6 shows [CO]SL measurements at different m˙ ∗jet and θ SL , for = 0 at x∗SL = 40 mm and r = 6.4 mm. To highlight the differences, the results in Fig. 6 were shifted by their values at m˙ ∗jet = 0%, summarised in Table 2. Fig. 6 shows that the air diluent case consistently results in lower overall CO, relative to the N2 diluent case, with an average difference of 46 ppm. This could suggest that the presence of the additional O2 in air favours increased CO oxidation. Fig. 6 also shows that as m˙ ∗jet is increased, [CO]SL first decreases, then increases to an average of 243 ppm, a similar value x∗jet

Fig. 6. Measurements of [CO]SL for different m˙ ∗jet at x∗SL = 40 mm, r = 6.4 mm and with x∗jet = 0 mm. Results are shown for θSL = 0◦ (red) and 45◦ (green) cases, as well as for air (dots) and N2 (solid) jets. The [CO]SL values have been shifted by the values at m˙ ∗jet = 0% to aid comparison. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article). Table 2 Baseline [CO]SL values. Case

Air diluent

N2 diluent

θSL = 0◦ θSL = 45◦

236 ppm 238 ppm

231 ppm 232 ppm

to the 234 ppm exhibited m˙ ∗jet = 0%. Measurements at different θ SL in Fig. 6, show that the θSL = 45◦ exhibits larger overall [CO]SL relative to the θSL = 0◦ case by an average of 22 and 34 ppm, for the N2 and air diluent cases, respectively. Radial profiles of TSL and [CO]SL for x∗jet = 0 are shown in Fig. 7, for two different θ SL measurement positions at x∗SL = 40 mm. The two θ SL positions of 0◦ and 45◦ represent the probe situated above and between the cooling jets, respectively. As shown in Fig. 7a, as m˙ ∗jet increases for the θSL = 0◦ case, the temperature maxima shifts radially outwards, and to lower temperatures. In contrast, Fig. 7b shows that for the θSL = 45◦ case, the temperature remains relatively similar between the different m˙ ∗jet values for r < 10 mm, after which the m˙ ∗jet = 2.5% case decreases in temperature. The TSL trends in observed in Fig. 7a and b are likely due to the way the cooling jet mixes with the bulk flow, however further investigations are required to quantify this effect. Comparing the TSL with the [CO]SL , Fig. 7c and d shows that the [CO]SL follows the same trends as TSL . For the θSL = 0◦ case, the [CO]SL maxima also moves radially outwards and to lower values as m˙ ∗jet is increased, matching the TSL trends shown in Fig. 7a. Similarly, Fig. 7d shows that [CO]SL varies significantly less between the cases where r < 10 mm, also matching the TSL behaviour shown in Fig. 7b. This demonstrates a direct correlation between TSL and [CO]SL , where regions of lower TSL corresponding to lower [CO]SL . While the TSL profiles in Fig. 7a and b show little difference between the air and N2 diluent cases, the [CO]SL results demonstrate a significant impact of the diluent type on CO oxidation. Fig. 7c and d shows that the air diluent case consistently produces lower [CO]SL , relative to the N2 diluent case. This agrees with the [CO]SL measurements in Fig. 6, as well as [CO]EG measurements far downstream of the flame at xEG = 210 mm for the three x∗jet configurations shown in Fig. 8. Similar to the SL measurements, the EG measurements in Fig. 8 were also shifted by their values at m˙ ∗jet = 0%, summarised in Table 3. As Fig. 8 shows, far from the flame the difference in [CO]exh behaviour between the air and N2 diluent cases are still exhibited, where the air diluent case produces lower measured [CO]exh . Additionally, Fig. 8 shows significantly lower end-gas CO, with an average value of 8.3 ppm, compared to the semi-local CO, with an average value of 185.6 ppm. This agrees with earlier measurements, where it was suggested that exhaust gas cooling due to wall heat transfer promotes CO oxidation [16]. This also supports the direct correlation between TSL and [CO]SL , shown in Fig. 7. The fundamental mechanisms driving this correlation are investigated further in the following section.

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Fig. 7. Radial profiles of TSL (a,b) and [CO]SL (c,d) at different m˙ ∗jet , for the θSL = 0◦ (a,c) and θSL = 45◦ (b,d) cases of the x∗jet = 0 mm condition. Results are shown for air (dots) and N2 (solid) jets at x∗SL = 40 mm. Also shown is the location of the wall relative to the measurement locations.

Fig. 8. Measurements of [CO]EG at different m˙ ∗jet at x∗EG = 210 mm. Results are shown for x∗jet > 0 mm (red), x∗jet = 0 mm (green) and x∗jet < 0 mm (blue), for air (dots) and N2 (solid) diluents. The [CO]EG values have been shifted by the values at m˙ ∗jet = 0% to aid comparison. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

Table 3 Baseline [CO]EG values. Case

Air diluent

N2 diluent

x∗jet < 0 x∗jet = 0 x∗jet > 0

10 ppm 9 ppm 8 ppm

10 ppm 9 ppm 8 ppm

5. Isolation of primary flame-cooling-air interaction phenomena In this section, 1D FRS results are presented to confirm the direct contribution of cooling jets and wall heat transfer, to the experimentally observed behaviour. The individual impacts of the cooling and dilution imposed by the cooling jets, as well as of exhaust gas cooling, on [CO]exh are shown in this section. The mechanisms responsible for their behaviour are then investigated using calculations of chemical equilibria under these conditions. It is clear from Figs. 6–8, that the air diluent cases consistently produce lower [CO]exh measurements, relative to the N2 diluent

case. Moreover, [CO]exh shows a direct correlation with Texh , with decreasing Texh resulting in decreasing [CO]exh . To further investigate the underlying mechanisms for these experimentally observed behaviour, three configurations of 1D FRS were conducted. This allows the isolation of three phenomena based on the experimental conditions in Table 1. The three simulated configurations are: upstream jet dilution (UJD); downstream jet dilution (DJD); and exhaust gas cooling (EGC). For the jet dilution simulations, both for the upstream and downstream configurations, the wall heat transfer model based on Cw and Tw was not solved. Conversely, for the exhaust gas cooling simulations, the cooling jet model based on m˙ ∗jet , x∗jet and Tjet , was not solved. In this way, the individual impacts of these phenomena to CO concentration is isolated. For the EGC case, Cw was tuned to provide a temperature of 430 K at the outlet, matching the experimental TEG . The UJD and DJD configurations contain air or N2 cooling jets located at x∗jet < 0 and x∗jet > 0, respectively. However, in contrast to the experiments where xflame is located at the flame tip, xflame for the simulations are defined at the location of the peak temperature gradient, which in this case is at x f lame = 30 mm. The cooling jet cases can be further split into sub-cases, one considering jet dilution only (DO), and one considering jet dilution and cooling (DC). These two sub-cases are identical for the upstream jet dilution configuration, and thus only jet dilution is presented. Therefore the four cases discussed in this section are: upstream jet dilution – jet dilution only (UDO); downstream jet dilution – jet dilution only (DDO); downstream jet dilution – jet dilution and cooling (DDC); and exhaust gas cooling (EGC). These cases and their simulation parameters are summarised in Table 4. 5.1. Effects of upstream jet dilution The effect of upstream jet dilution, where x∗jet < 0, is shown in Fig. 9, for the upstream jet dilution – dilution only, air diluent case with m˙ ∗jet = 2.5%. This shows that the diluted case results in lower end-point temperature (Tend ) and CO (COend ). In addition, the gradient of both the temperature and CO in the CO burn-out region changes. This is expected, given that the addition of air upstream of the flame is equivalent to decreasing the equivalence ratio of the flame. Similarly, for the N2 diluent case, its behaviour is equivalent

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381

Table 4 Summary of numerical cases presented. Case

Cw (m−2 )

Tw (K)

m˙ ∗jet (%)

x∗jet (mm)

Diluent

Width (mm)

Tjet (K)

UDO DDO DDC EGC

– – – 0.25

– – – 300

0–2.5 2.5 2.5 –

-15.88 +15.88 +15.88 –

Air/N2 Air/N2 Air/N2 –

3.03 3.03 3.03 –

T(x) T(x) 300 –

Fig. 10. End-point temperature (green) and CO (red) for different m˙ ∗jet , for both air (dots) and N2 (solid) as the diluents, for the UDO cases. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

Fig. 9. Temperature (a) and CO (b) profiles for diluted (solid) and undiluted (dots) flames. Results are shown for the UDO case (x∗jet < 0), with air as the diluent. Insets focus on the region of interest. Note that data are shifted to account for the different flame speeds and aid interpretation. Table 5 Summary of Tend , Teq , COend and COeq for the UJD case at the m˙ jet = 2.5% condition. Diluent

Tend (K)

Teq (K)

COend (ppm)

COeq (ppm)

N2 Air

2107 2110

2096 2092

1837 1733

1912 1808

to inert gas dilution. Both of these effects have been previously shown to decrease both Tend and COend [33]. This is supported by comparing the equilibrium temperature (Teq ) and CO concentration (COeq ) values to Tend and COend , as shown in Table 5. As shown, the end-point and equilibrium values are of similar magnitudes. Values of Tend and COend for all m˙ ∗jet are shown in Fig. 10. This demonstrates that Tend for both air and N2 diluent cases decrease as m˙ ∗jet is increased, with COend matching this trend. However despite Tend for both air and N2 diluent cases exhibiting little difference, the air diluent case presents about 3% lower overall COend relative to the N2 diluent case. 5.2. Effects of downstream jet dilution The effect of downstream jet dilution, where x∗jet > 0, is shown in Fig. 11a and b, for temperature and CO respectively. These show the results for the DDO and DDC cases for air and N2 as

Fig. 11. Comparison of temperature (a) and CO (b) profiles for the DDC (red), DDO (green) and undiluted (blue) cases, for air (dots) and N2 (solid) diluents. Results are shown for x∗jet > 0 case. Insets focus on the region of interest. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

the diluents, as well as the undiluted flame solution. Fig. 11a shows the overlapping temperature profiles of the downstream jet dilution – jet dilution only case (DDO), and undiluted case, for both air and N2 diluent cases. This is due to the cooling jets being injected at the local gas temperature. In contrast, the downstream jet dilution – jet dilution and cooling case (DDC) shows a decrease in temperature due to the cooling jets being at T jet = 300 K, which reduces the total enthalpy of the resulting mixture.

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Table 6 Summary of Tend , Teq , COend and COeq for the different x∗jet > 0 cases, at the m˙ ∗jet = 2.5% condition. Case

Diluent

Tend (K)

Teq (K)

COend (ppm)

COeq (ppm)

Undiluted DDO DDO DDC DDC

– N2 Air N2 Air

2144 2143 2144 2110 2112

2144 2139 2138 2102 2115

2436 2393 2208 1898 1756

2436 2387 2068 1781 1558

The corresponding CO profiles are shown in Fig. 11b and summarised in Table 6. As these illustrate, the N2 diluent case of the downstream jet dilution – jet dilution only case, shows minimal change in post-flame CO, at ≈ 43 ppm less than the undiluted case. In contrast, the equivalent air diluent case results in 228 ppm less post-flame CO. When considering the change in temperature of the downstream jet dilution – jet dilution and cooling case, both the air and N2 diluent cases show lower post-flame CO, at an average of 609 ppm. The air diluent case also results in 142 ppm less COend relative to the N2 diluent case. This implies that the addition of O2 in the post-flame exhaust promotes CO oxidation, independently of a decrease in temperature. If the temperature does decrease, Fig. 11 shows that CO oxidation is promoted regardless of the diluent type. Both of these conclusions agree with the experimentally observed behaviour in Figs. 6–8, as well as the equilibrium values shown in Table 6. 5.3. Effects of exhaust gas cooling The effect of exhaust gas cooling (EGC) on the flame is illustrated in Fig. 12, which shows the temperature (Fig. 12a) and CO (Fig. 12b) profiles for both the adiabatic and EGC cases. Fig. 12a shows the exhaust gas cooling case decreasing in temperature,

Fig. 12. Comparison of temperature (a) and CO (b) profiles with distance, for the exhaust cooled (green) and adiabatic (red) cases. Insets focus on the region of interest. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

relative to the adiabatic case, as a result of the imposed heat loss. Correspondingly, Fig. 12b shows a significant deviation in the CO profiles between the EGC and adiabatic cases. It demonstrates that a decreasing exhaust gas temperature promotes significant postflame CO oxidation. The COend for Fig. 12b is 11 ppm, compared with the experimentally measured average [CO]EG of 8–10 ppm shown in Fig. 8. Fig. 12b also shows that the shape of the peak CO concentration remains the same between the adiabatic and EGC cases. This agrees with the timescale analyses conducted by Kosaka et al. [27] and Jainski et al. [34]. They found an order of magnitude difference in the timescales of CO production and oxidation, implying that the CO oxidation branch is more likely to be disrupted by heat transfer effects. In addition, the exhaust gas cooling result also agrees with the CRN simulations accounting for exhaust gas cooling conducted by Rivera et al. [16], who showed good agreement with experimentally measured [CO]exh results, for a flame undergoing pure FWI. This result further supports the direct correlation between Texh and [CO]exh that was experimentally observed, where decreasing Texh promotes CO oxidation. 5.4. Chemical equilibrium of cooling jet dilution and exhaust gas cooling To investigate the mechanism of the CO correlation with temperature and m˙ ∗jet , an equilibrium analysis was conducted. This was conducted using a constant pressure (1 atm) equilibrium reactor, fed with post-flame combustion products and varying levels of m˙ ∗jet , at different temperatures, and for both the air and N2 diluent cases. The COeq concentration is plotted in Fig. 13. To appropriately compare the CO results from the different m˙ ∗jet cases, these results are presented as dry concentrations, corrected to 15% O2 . This type of normalisation, a standard method for power-generation gas turbines [1], is required to compare the global and homogeneously

Fig. 13. COeq concentrations at different temperatures and m˙ ∗jet , for the air (a) and N2 (b) diluent cases. White iso-lines of CO concentration are shown in ppmvd to aid interpretation.

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mixed results of the equilibrium calculations. In contrast, this is inappropriate for the local FRS results and the potentially inhomogeneously mixed experimental results. For both the air and N2 diluent cases, COeq is less than 5% of the adiabatic value when the temperature decreases below ≈ 1800 K. As the temperature is decreased further, COeq approaches zero. This illustrates that as the temperature decreases, CO oxidation is thermodynamically favoured. Increasing m˙ ∗jet at a constant temperature, leads to lower COeq for the air diluent case, relative to the N2 diluent case. This demonstrates that exhaust gas oxygenation also thermodynamically favours CO oxidation. Thus, the experimentally and numerical observed decrease in CO concentration with decreasing temperature and increasing O2 can be attributed to this effect, and results in complete CO oxidation given sufficient residence time. 6. Decoupling of cooling jet dilution and exhaust gas cooling residence times As shown in the previous sections, decreasing exhaust temperatures and increasing exhaust O2 concentrations lead to decreasing CO concentrations due to thermodynamic favouring of CO oxidation. However, given the infinite-time nature of chemical equilibria, systems with short residence times may limit full CO oxidation. To explore the limitations of this behaviour, CRN simulations were conducted to determine the characteristic timescales of CO oxidation under the influence of cooling jets and exhaust gas cooling. The chemical kinetic timescales of the downstream jet dilution - jet dilution and cooling case (DDC) and the exhaust gas cooling (EGC) case, can be investigated through the use of standard idealised chemical kinetic reactors. In this study, a network consisting of a freely propagating flame (FPF), a perfectly stirred reactor (PSR) and a plug flow reactor (PFR) has been used (Fig. 14). This decouples the mean dilution residence time (tdil ) of the cooling jet and the flame exhaust, from the exhaust gas cooling residence time (tcool ). These simulations were conducted using the Chemkin package [30], with the USC C1 − C3 mechanism [28]. As shown in Fig. 14, a mixture of C3 H8 − Air at φ = 0.85 enters the FPF, which then generates flame exhaust at TAd . The flame exhaust, with a temperature of TF PF = TAd and a mass flow rate of m˙ F PF , is then mixed in the PSR with the cooling jet, at a temperature of Tjet and a mass flow rate of m˙ jet . In the PSR, both effects of jet dilution and cooling occur, resulting in Tmix and COmix . The PSR outlet then enters the PFR with a specified temperature profile, calculated analytically using Eq. (6), with Cw tuned to give a PFR Tend equal to TEG = 430 K within 250 ms. These end-point temperature and residence times were chosen to match the experimentally measured temperature and the estimated experimental residence time of τexp = 250 ms, as per Section 2.1. The PSR simulates the mixing of two or more fluids, including their subsequent reaction, and is often used in CFD-CRN simulations of gas turbine-like geometries (e.g., [35,36]). The tdil of the PSR is determined as follows:

tdil =

ρV , m˙ in

383

Table 7 Summary of CRN cases presented. Case

tdil (ms)

m˙ ∗jet (%)

Tjet (K)

tcool (ms)

tdil Run m˙ ∗jet Run

0–17.8 1.1

2.5 2.5–20

300 300

0–250 0–250

where V is the volume of the reactor, and m˙ in is the sum of the mass flow rates into the PSR, which is equal to m˙ in = m˙ F PF + m˙ jet . As m˙ jet increased with dilution level, both m˙ in and ρ changes, the latter through the cooling experienced by the mixing of the cold m˙ jet with the hot m˙ F PF . Thus, V was changed to accommodate this and achieve the desired tdil which was varied throughout this study. The PFR simulates the evolution of perfectly radially homogeneous gas through a streamtube, and has also previously been used in CFD-CRN studies, connected to PSRs, to model the exhaust section of the combustor (e.g., [37,38]). The tcool of the PFR is determined as follows:

tcool (x ) =



x→L 0

1 dx, u (x )

(19)

where L is the PFR length, and u(x) is the local gas velocity. To control tcool , L was varied while accounting for the changing u(x) due to the decreasing temperature profile in the PFR. The individual behaviours of the PSR and PFR were tested against the DDC and EGC FRS cases, and showed temperature and CO profile differences of less than 7%, and is considered appropriate for this study. Thus, simulations were conducted varying tdil , m˙ ∗jet and tcool , for both the air and N2 diluent cases. The simulated cases are summarised in Table 7, with the fully mixed CO results presented as dry concentrations, corrected to 15% O2 . 6.1. Perfectly stirred reactor and cooling jet dilution The PSR results for variations in tdil are illustrated in Fig. 15, which shows COmix and Tmix for different tdil . It shows that as tdil is increased, Tmix increases slightly by 4 K. However despite the Tmix curves for both diluent types overlapping, the air diluent case produces an average of 159 ppm less COmix , relative to the N2 diluent case in Fig. 15. As tdil increases, the COmix for both diluent types decreases by an average of 372 ppm. The PSR results for variations in m˙ ∗jet are illustrated in Fig. 16, which shows COmix and Tmix for different m˙ ∗jet , at tdil = 1.1 ms. This tdil was chosen to represent the fast mixing timescale case. As m˙ ∗jet is increased, the Tmix decreases by 232 K. This is expected, as increasing m˙ ∗jet increases the jet contribution to the mixture enthalpy, which lowers Tmix towards Tjet . Fig. 16 also shows that

(18)

Fig. 14. Schematic of the chemical reactor network used in this study, consisting of a freely propagating flame (FPF), a perfectly stirred reactor (PSR), and a plug flow reactor (PFR).

Fig. 15. PSR-exit temperature (green) and CO (red) at different tdil , for air (dots) and N2 (solid) as the diluents. Results are shown for m˙ ∗jet = 2.5% and T jet = 300 K. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

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COmix decreases by an average of 1143 ppm as m˙ ∗jet is increased, with the air diluent case producing an average of 412 ppm less COmix , relative to the N2 diluent case. High levels of cooling jet dilution of flame exhaust promotes significant CO oxidation, even at tdil = 1.1 ms. 6.2. Plug flow reactor and exhaust gas cooling The PFR results for variations in tdil are illustrated in Fig. 17, which shows the CO concentration along the PFR. Increasing the PSR tdil from 0.6 to 17.8 ms reduces the COmix input into the PFR by 353 ppm for the N2 diluent case, and by 390 ppm for the air diluent case, as in Fig. 17 (inset). Despite the different initial CO values with tdil , by tcool = 4 ms these differences converge. However, the differences in CO between the air and N2 diluent cases do not converge until tcool ≈ 10 ms. The convergence is likely due to the convergence of the temperature profiles, which for all cases are tuned to produce 430 K at tcool = 250 ms, the approximate experimental end-gas T and tcool . These results demonstrate that sufficient exhaust cooling tcool results in CO converging towards their cold equilibrium values. Thus despite varying PSR tdil producing different COmix values, long exhaust cooling tcool reduces the impact on COend . The PFR results for variations in m˙ ∗jet are illustrated in Fig. 18, which shows CO concentration along the PFR. As expected, the large variation in COmix shown in Fig. 16 produces significant differences in the initial CO result of the PFR. As a result, the convergence of the CO − tcool profiles observed in Fig. 18 varies greatly with m˙ ∗jet . At m˙ ∗jet = 20%, both the air and N2 diluent cases converge at tcool ≈ 20 ms, with this convergence point increasing as m˙ ∗jet is decreased. High levels of m˙ ∗jet promote CO oxidation and can reduce the required tcool for complete CO oxidation.

To further investigate the sensitivity of the predicted CO to the assumed m˙ ∗jet , the proportion of CO oxidised in the PSR (%COOx,PSR ) was calculated for different PFR tcool . The results of this calculation is shown in Fig. 19. As m˙ ∗jet is increased, %COOx,PSR increases. As tcool is increased, %COOx,PSR decreases up to the value at tcool = 50 ms, after which they overlap. As tcool → 50 ms, %COOx,PSR decreases from 62.4% to 16.0% at m˙ ∗jet = 2.5%, while %COOx,PSR decreases from 79.9% to 68.4% at m˙ ∗jet = 20%. The sensitivity of %COOx,PSR to tcool at low m˙ ∗jet suggests that as tcool increases, the PFR tends to share a greater portion of CO oxidation, relative to the PSR. Thus in this case, the exhaust gas cooling model has a greater impact on the CO oxidised, reducing the CO concentration by about 30% as tcool is increased. Conversely, the insensitivity of %COOx,PSR to tcool at high m˙ ∗jet , implies that the majority of CO oxidation occurs only in the PSR. Therefore in this case, the cooling jet dilution model has a greater impact on the CO oxidised, with the PFR only oxidising 10% more as tcool is increased. This illustrates that under certain conditions, specifically at high m˙ ∗jet , greater model fidelity of the cooling jet mixing and dilution processes may be required to accurately predict exhaust CO concentration. 7. Reaction timescale analysis As shown in the previous section, CO oxidation can be limited by residence time, and vary with dilution timescale, jet mass flow fraction, and exhaust cooling rate. In this section, the dynamics of CO oxidation are explored further by considering timescales relevant to practical combustion systems. For the cooling jets, the combined effect of tdil and m˙ ∗jet variations on CO oxidation are explored in the context of timescales relevant to gas turbine combustors. Meanwhile, the sensitivity of CO oxidation to exhaust cooling

Fig. 16. PSR-exit temperature (green) and CO (red) at different m˙ ∗jet , for air (dots) and N2 (solid) as the diluents. Results are shown for tdil = 1.1 ms and T jet = 300 K. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

Fig. 18. PFR CO − tcool profiles at different PSR m˙ ∗jet , for air (dots) and N2 (solid) as the diluents. Results are shown for the case where tdil = 1.1 ms and T jet = 300 K.

Fig. 17. PFR CO − tcool profiles at different PSR tdil , for air (dots) and N2 (solid) as the diluents. Results are shown for the case where m˙ ∗jet = 2.5% and T jet = 300 K.

Fig. 19. %COOx,PSR for different m˙ ∗jet and tcool , for the air diluent case. Results are shown for tdil = 1.1 ms and T jet = 300 K. Note that lines for tcool ≥ 50 ms overlap.

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rates are explored by considering temperature drops and cooling residence times associated with the combustor and turbine sections of gas turbines. 7.1. Timescale analysis of cooling jet dilution Based on the results from the study using the CRN in Fig. 14, it was demonstrated that in the PSR, the choice of tdil and m˙ ∗jet can have a significant impact on the resulting CO concentration. To further explore the interaction between these two parameters, simulations were conducted using only a PSR, with high resolution variations in m˙ ∗jet and tdil over 0 − 50% and 0.1 − 100 ms, respectively. These simulations were conducted with T jet = 300 K, for both the air and N2 diluent cases. The dry, 15% O2 corrected CO concentrations are plotted in Fig. 20. For reference, three characteristic timescales (τ ) are also plotted. The first is the characteristic CO oxidation timescale (τ Ad ) of an adiabatic freely propagating flame. This was calculated using the method outlined in Kosaka et al. [27], where τ was measured from the time of the peak CO concentration, to when the concentration drops to within 10% of the end-point value. For a C3 H8 − Air flame at φ = 0.85, this was calculated to be τAd = 4.1 ms. Using the same method, the characteristic CO oxidation timescale for the exhaust-cooled flame (τ EGC ) in Fig. 12 was also calculated, resulting in τEGC = 7.4 ms. This increase in τ results from the lower COend value for the exhaust-cooled flame relative to the adiabatic flame, as shown in Fig. 12. The final τ presented in Fig. 20 is an indicative residence time of a gas turbine combustor, estimated to be τGT = 20 ms [39]. Fig. 20 shows that, as tdil is increased, the CO concentration decreases. This implies that given long dilution residence times between the flame exhaust and the cooling jet, CO oxidation

Fig. 20. Contour plot of CO at different m˙ ∗jet and tdil , for air (a) and N2 (b) as the diluents. Results are shown for T jet = 300 K. Also shown are the characteristic timescales of τ Ad (black dash), τ EGC (blue dash) and τ GT (red dash). (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

385

continues to occur as is dictated by COeq . However, at m˙ ∗jet < 15%, the CO concentration appears to achieve a steady-state value at tdil > τ Ad . In contrast, at m˙ ∗jet > 15%, the CO concentration does not achieve a steady-state value. For this condition, the CO concentration continues to decrease even after tdil > τ GT . Hence the relative species concentration differences due to varying m˙ ∗jet , affects the CO oxidation timescales and the sensitivity to differences in assumed tdil . 7.2. Timescale analysis of exhaust gas cooling As illustrated in Fig. 20, variations in tdil and m˙ ∗jet result in differences in temperature and CO concentration, however the exhaust cooling in the PFR was shown to also influence COend concentration. To further explore this effect in isolation, simulations were conducted using only a PFR, fed with undiluted (m˙ ∗jet = 0%) flame exhaust. Decreasing temperature profiles were imposed on the PFR from TAd to Tend , over different maximum tcool values (max(tcool )), thus inducing different cooling rate conditions. These temperature profiles were generated analytically using Eq. (6), with Tw equal to Tend , and Cw varied to achieve these Tend values over max(tcool ) values of 5 − 20 0 0 ms. Three different Tend values were used to represent three GT-relevant exhaust temperatures. A Tend value of 1770 K represents combustor exit temperatures, while 860 K represents compressor exit temperatures and 430 K represents engine-out exhaust temperatures [5]. The temperature profiles used in these PFR simulations and their corresponding CO profiles are shown in Fig. 21. The temperature profiles are shown in linear time scale to highlight their shape, while the CO profiles are shown in logarithmic time scale to highlight their COend behaviour. In Fig. 21a, the temperature only decreases by 354 K, so the profiles appear almost linear, with decreasing max(tcool ) resulting in increased temperature gradients. The corresponding CO profiles are shown in Fig. 21b. As can be seen, the COend values for shorter max(tcool ) are higher than for longer max(tcool ), despite Tend being identical for each max(tcool ) case. This demonstrates that, while COeq thermodynamically favours CO oxidation, it can be limited by the max(tcool ). Fig. 21c shows the PFR temperature profile for the Tend = 860 K case. The larger 1264 K decrease in temperature shows more non-linear temperature profiles, with decreasing max(tcool ) also increasing the temperature gradient. In response, the CO profiles are shown in Fig. 21d. As tcool → max(tcool ), the CO profile gradients “flatten”, which implies reduced CO oxidation rates. Furthermore, these occur at COend values that decrease as max(tcool ) is increased, implying increased overall CO oxidation as the total residence time is increased. These results demonstrate that CO oxidation is also dependent on the cooling rate and the total residence time, in addition to the absolute temperature. Fig. 21e and f shows the PFR temperature and CO profiles, respectively, for the Tend = 430 K case. This condition has the largest temperature decrease, and produces steep and highly non-linear temperature gradients. The CO profiles for the Tend = 430 K case also show the gradient flattening behaviour as tcool → max(tcool ), similar to the Tend = 860 K case. However in contrast, the COend values that these correspond to are higher for the Tend = 430 K case. This result suggests that the increased cooling rate of this case, due to the higher temperature drop, can result in significant inhibition of CO oxidation, leaving values of COend greater than 1500 ppm. The COend values for these different Tend and max(tcool ) cases are summarised in Fig. 22. For the Tend = 1770 K case, COend presents a decreasing trend with increased max(tcool ), converging to a value of 115.7 ppm at max(tcool ) = 20 0 0 ms. This COend value is close to the

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Fig. 21. Profiles of temperature (left column) and CO (right column) in the PFR at varying max(tcool ). Results are shown for Tend values of 1770 K (crosses), 860 K (triangles) and 430 K (circles). The CO results are shown in logarithmic time scale to highlight differences in COend , while the T results are shown in linear scale to highlight their shape. The Tend value (black) is also highlighted.

Fig. 22. Values of COend (solid) for different max(tcool ), plotted on a logarithmic time scale, for Tend values of 1770 K (crosses), 860 K (triangles) and 430 K (circles). Also shown are the COeq (dots) values for the different Tend cases.

COeq value of 113.1 ppm dictated by this Tend condition. Thus, this can be considered as reaching equilibrium and implies an approximately zero net oxidation rate. This oxidation regime can be described as “equilibrium-dominated” since the oxidation behaviour is largely dictated by the equilibrium state. For the Tend = 860 K

case, a similar decrease in COend is exhibited as max(tcool ) is increased, but converges to a lower COend value of 0.005 ppm at max(tcool ) = 20 0 0 ms. The COend value at this max(tcool ) condition is due to the COeq of 0.002 ppm, dictated by a Tend = 860 K. This difference in COend at max(tcool ) = 20 0 0 ms, highlights the strong temperature dependence of CO oxidation, not only due to the rate constant, but also due to the COeq imposed by the temperature. In contrast, the Tend = 430 K case shows a COend value of 0.12 ppm at max(tcool ) = 20 0 0 ms, which is three orders of magnitude larger than the corresponding COeq value of 0.0 0 09 ppm. Given the flattening of the CO profiles shown in Fig. 21, this is evidence of decreasing temperatures inhibiting CO oxidation despite COeq thermodynamically promoting CO oxidation. Thus, this regime of oxidation behaviour can be described as “kineticallydominated”, where oxidation is limited by the kinetic rate despite being thermodynamically favoured. Under this condition, high COend values are achieved, with 68.7 ppm produced at max(tcool ) = 100 ms. This is similar in order of magnitude to [CO]exh produced by industrial gas turbines [1], despite not accounting for changes in pressure. These results demonstrate competing effects on CO oxidation, that can be highly influential in determining [CO]exh .

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8. Conclusion This study investigated a laminar, premixed flame interacting with cold gas jets. A novel burner and wall configuration were developed to experimentally induce flame-cooling-air interaction (FCAI). Measurements of flame shape, exhaust temperature (Texh ) and exhaust CO emissions ([CO]exh ) were conducted to determine their response to different cooling jet mass flow fractions (m˙ ∗jet ) and diluent types, namely air and N2 . The conditions that the cooling jets imposed on the flame were then isolated using onedimensional (1D) fully resolved simulations (FRS) with a detailed chemical kinetic mechanism. A novel 1D cooling jet model was introduced to simulate the effects of jet cooling and dilution independently. Chemical reactor network (CRN) modelling was then conducted to further isolate the cooling jet dilution and exhaust gas cooling processes, using a perfectly stirred reactor (PSR) and a plug flow reactor (PFR), respectively. The primary findings are summarised as follows: • Decreased [CO]exh was experimentally found to correspond to decreased Texh , with the air diluent case producing lower [CO]exh relative to the N2 diluent case. The FRS demonstrated that decreasing Texh and increasing exhaust O2 concentration introduced by the cooling air independently produce decreased exhaust CO concentrations. This behaviour was illustrated by an equilibrium analysis to be due to the thermodynamic favouring of CO oxidation under these conditions. • The CRN modelling confirmed that longer exhaust gas cooling residence times promote more complete CO oxidation. Exhaust gas cooling residence times greater than 50 ms were found to reduce the sensitivity of predicted [CO]exh by 30%, to modelling assumptions regarding cooling jet dilution. However, this effect was found to be less significant for high m˙ ∗jet , reducing the sensitivity by only 10% which increases the contribution of cooling jet dilution to the overall [CO]exh oxidation. Furthermore, high m˙ ∗jet was found to increase the sensitivity of [CO]exh to the cooling jet dilution timescale, which has implications for modelling. • In the absence of cooling jet dilution, exhaust temperature reduction from adiabatic to 430 K, within 100 ms, were found to result in [CO]exh of about 70 ppmvd (corrected to 15% O2 ). Competing effects were found to influence this behaviour, with “kinetically-dominated” CO oxidation behaviour at steep cooling gradients and short residence times producing high levels of [CO]exh . These findings indicate that exhaust cooling conditions influence [CO]exh , even downstream of the combustor. This extends the literature regarding CO oxidation in gas turbines [40], and should be considered to develop more robust and predictive engine-out emissions models. This study has found that initial temperature and composition, exhaust O2 content, exhaust cooling rate and residence time, and final exhaust temperature, all compete in the determination of [CO]exh . The results presented show the influence of exhaust gas cooling and cooling jet dilution on exhaust CO emissions. These findings should be extended to explore the effect of turbulence, pressure and wall reactivity, as well as partial or imperfect premixing. This will further quantify how FCAI influences the exhaust CO emissions of practical gas turbines. Acknowledgments This research was supported by the University of Melbourne through the Research Training Program Scholarship. The research benefited from the computational resources provided through the Energy & Resources Scheme using the Pawsey Supercomputing Centre, and through the National Computational Merit Allocation

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Scheme using the Australian NCI National Facility, all supported by the Australian Government. The authors acknowledge the generous support of the European Centre for Research and Advanced Training in Scientific Computing (CERFACS), for providing NTMIXChemkin. The authors especially thank Dr. Bénédicte Cuenot for her help with NTMIX-Chemkin. The authors are grateful for the help of Dr. Arne Scholtissek, Prof. Dr.-Ing Christian Hasse and Prof. Dr. habil Andreas Dreizler of the Technische Universität Darmstadt, in applying their heat transfer coefficient formulation to this work.

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