A new CO conversion quench reactor design

Fuel Processing Technology 148 (2016) 198–208

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Research article

A new CO conversion quench reactor design Konrad Uebel a, Philip Rößger b, Uwe Prüfert c, Andreas Richter b,⁎, Bernd Meyer a a b c

Chair of Energy Process Engineering and Thermal Waste Treatment, Technische Universität Bergakademie Freiberg, Fuchsmühlenweg 9, 09599 Freiberg, Germany CIC Virtuhcon, Technische Universität Bergakademie Freiberg, Fuchsmühlenweg 9, 09599 Freiberg, Germany Institute of Numerical Mathematics and Optimization, Technische Universität Bergakademie Freiberg, Akademiestraße 6, 09599 Freiberg, Germany

a r t i c l e

i n f o

Article history: Received 18 December 2015 Received in revised form 15 February 2016 Accepted 16 February 2016 Available online xxxx Keywords: Design development Syngas cooling Quench Water-gas shift reaction Numerical simulation CFD Coal gasification

a b s t r a c t Syntheses downstream of coal gasifiers face specific requirements regarding syngas quality. High syngas H2/CO ratios are a typical requirement. A new quench concept for entrained-flow gasifiers has been developed in order to increase the H2/CO ratio within the quench vessel and thereby minimize efforts in the subsequent catalytic water–gas shift reaction (WGSR). To achieve this target, steam instead of water will be injected into the quench chamber to realize CO conversion via homogeneous WGSR. The concept allows a flexible mode of operation, such as full or partial water quenching with subsequent waste heat utilization for high and mediumpressure steam production. This allows self-sufficiency in terms of steam and additionally offers the potential to increase the overall plant efficiency. To test the concept, three different syngases from various feedstock and two inlet design configurations will be investigated using CFD and compared to one other. The numerical setup has been validated against gasification reactor measurements and simulation, respectively. A sensitivity analysis shows the performance and flexibility of the proposed concept. © 2016 Elsevier B.V. All rights reserved.

Nomenclature Latin symbols d h ṁ q_ Q_

m m kg/s W/m2 W

Diameter Height Mass flow Heat flux Heat

A Ap E k L Nu R S T u r y+

m2 (cm3/mol)0.5/s J/kmol W/m2 K m −J/kmol K kg/m2 s K m/s (cm3/mol)0.5/s −-

Surface Pre-exponential factor Activation energy Heat transfer coefficient Characteristic length Nusselt number Universal gas constant Source term Temperature Gas velocity Reaction rate Wall distance

Abbreviations: BYU, Brigham Young University; DO, Discrete Ordinates; EDC, Eddy Dissipation Concept; IEC, Institute of Energy Process Engineering and Chemical Engineering; ISAT, In-Situ Adaptive Tabulation; SST, Shear Stress Transport; UDF, User Defined Function; WGSR, water-gas shift reaction; WSGG, Weighted-sum-of-gray-gas. ⁎ Corresponding author. E-mail address: [email protected] (A. Richter).

http://dx.doi.org/10.1016/j.fuproc.2016.02.022 0378-3820/© 2016 Elsevier B.V. All rights reserved.

Greek symbols ΔR H∘ Δhv λ Indexes 0 BL conv eq evap f inl LS out rad S SG SL surf tot W WB

kJ/mol kJ/mol W/m K

Standard reaction enthalpy Enthalpy of vaporization Thermal conductivity

Initial Boundary layer Conversion Chemical equilibrium state Evaporation Forward Inlet Lateral Surface Outlet Radiation Steam Syngas Slag Surface otal Wall Waterbath

1. Introduction Coal-based syngas consists of carbon monoxide and hydrogen together with other gas and solid components, such as CO2, ash and/or slag [17]. For further processing, subsequent gas cooling or quenching

K. Uebel et al. / Fuel Processing Technology 148 (2016) 198–208

and gas treatment are necessary to improve syngas quality to the level considered sufficient for syntheses or combustion [3,14]. Here, water or chemical quenching and radiative or convective waste heat recovery are common methods. Often, a combination of those methods is utilized [17]. The choice of an appropriate syngas cooling concept mainly depends on the gas exit temperature (slagging/non-slagging operation) and the gas composition with a focus on gas impurities, such as alkalis. Partial or full quenching methods are usually used to pass the sticky ash/ slag temperature zone in entrained-flow gasifiers [14]. The syngas has to be shifted to the product requirements, characterized by the H2/CO or stoichiometric ratio, often referred to as the syngas module M or stoichiometric number SN = (XH2 − XCO2)/(XCO − XCO2) [14]. Here, X denotes the molar fraction of the specific components in the syngas. The required syngas quality for different syntheses and products varies between 1 and 2.05 for SN or 0.6 to maximum attainable H2/CO ratios for H2-based syntheses [3,14,17]. Here, the water–gas shift reaction (WGSR) is used to increase hydrogen yield and/or adjust the H2/CO ratio for syngas applications [3]. Normally, two-stage catalytic water–gas shift reactors in the gas scrubbing process, operating at low temperatures of 473–723 K are used, as the temperature resistance of typically applied iron–chromium and copper–zinc catalysts is the limiting factor [14,17]. One approach to improve syngas quality in terms of the H2/CO ratio during the quench process could be the utilization of the equilibriumlimited and moderately exothermic homogeneous water–gas shift reaction (WGSR, see Eq. (1)) at high temperatures: CO þ H 2 O ⇌ CO2 þ H 2

ΔR H ∘ ¼ −41:1 kJ=mol:

ð1Þ

The high-temperature kinetics (1063–1198 K) of the homogeneous, forward and reverse WGSR at both low (0.1 MPa) and high pressure conditions (1.6 MPa) have been measured in the empty quartz reactor by Bustamante et al. [9,10]. The gas-phase mechanism devised by Bradford [5] was identified to give adequate results in numerical simulation of different reactors investigated. The GRI3.0 mechanism database [30] was also applied to calculate the individual reaction rate constants for their numerical simulations of forward WGSR. The effect of high-pressure conditions on the reaction rate was negligible. Significant CO conversion velocities without a catalyst can only be achieved at gasifier exit temperatures above 1100 K. Thus, the potential for non-catalytic water–gas shift conversion in water-based quench concepts and devices is restricted to hightemperature processes only. Kiso and Matsuo [21] investigated the potential of increasing the H2/ CO ratio, applying a 1D single/multi-stage CO conversion reactor subsequent to the outlet of the entrained-flow EAGLE gasifier. To achieve this, they used Arrhenius reaction rate expressions and homogeneous WGSR kinetics developed by Bustamante et al. [10] for their investigation. The results show that CO conversion and therefore the achievement of equilibrium suffer from liquid water injection in a single-stage reactor and rapid temperature drops should be prevented in order to enhance CO conversion. Uebel et al. [32] evaluated different cooling concepts in detail for a Siemens entrained-flow gasifier applied in an IGCC plant. Gasifiers with exit temperatures below the sticky ash zone usually employ partial quenching and convective cooling, as the syngas is no longer contaminated with tars and particulate matter [14]. However, Siemens and GE showed that radiant waste heat recovery is a feasible option even for entrained-flow gasifiers, if contact of sticky mineral matter on cold surfaces, which causes plugging, can be prevented [24,32]. In this work, a high-pressure partial/full quench conversion concept for syngas treatment and waste heat recovery will be introduced. The main aim is to increase the H2/CO ratio inside the quench, while moderate syngas cooling takes place for subsequent waste heat recovery. In this context, the proof of concept for three different syngases, using CFD models, is another major task.

199

Combining different feedstocks and gasification technologies provides a wide range of possible process parameters, such as the pressure, gasification temperature or syngas composition. Based on the temperature level and H2 /CO ratio at the gasifier exit, entrained-flow gasifiers should have the highest CO conversion potential, whereas dry-feeding compared to slurry feeding systems show the lowest overall H2/CO ratio (see Table 1). Additionally, entrained-flow gasifiers also show a high waste heat recovery potential for the above mentioned reasons and thus for the increase in the net efficiency of the process [32]. 2. Quench conversion reactor concept A new water-based syngas cooling concept with integrated CO conversion, the so called quench conversion, has been developed at the IEC [4,25]. The main task is to increase the H2 /CO ratio inside the quench chamber. Depending on the mode of operation, the quench concept allows partial to full quenching down to a syngas exit temperature of 500 K. However, higher exit temperatures of up to 1375 K for waste heat recovery and high- and medium-pressure steam production are preferred. As the WGSR reactant, steam is injected into a highly controlled flow regime rather than the commonly used water. As part of this concept, two designs, either a lateral or a central steam inlet, are conceivable, thus both are investigated. The CO conversion COconv in this work is defined as follows: COconv ¼

_ out  Y CO;0  m _ SG Y CO;out  m ; _ SG Y CO;0  m

ð2Þ

where YCO,0 and YCO,out represent the mass fraction of CO in the unconverted syngas from the gasifier and quench outlet, respectively. 2.1. Process principle The reactor consists of a central syngas inlet, cooled reactor walls and a waterbath at a constant level in the bottom area. Depending on the design, steam is injected into or around the syngas in the quench inlet area. Water nozzles are arranged in a ring just below the surface of the waterbath, pointing in the direction of the reactor walls. The quench outlet or outlets are situated in the upper third of the reactor. The concept does not provide any fixtures, hence slag deposition can not be expected, which should make the quench design extremely robust and reliable. The effects of heat transfer between slag and gas and the slag flow behavior were investigated in prior works and were found to be negligible due to the short slag residence time [32]. A general scheme of the quench conversion reactor is illustrated in Fig. 1. Hot unconverted syngas from the gasifier flows through the inlet area, where steam with 2 bars of excess pressure is added into the quench chamber, forming a central downflow towards the waterbath surface. The use of steam instead of water prevents a sudden cooling effect due to evaporation. The evolving flow and mixture profile, as well as high mixture temperatures, lead to the formation of a distinct

Table 1 Ranking of theoretical CO conversion potential for the gasification technologies considered [14,17,29]. Pressure range

H2/CO ratio

K

bar

mol/mol

1573–1973 1400–1723 N1223

1–86 30–70 35

0.36–0.97 1.6–1.9 1.8–6

Technology

Temperature range

– Entrained-flow Gas-POX Lurgi MPG

200

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quench inlet close to it and pointing in the direction of the syngas downflow. As a result, steam is sucked in the central flow, forming a conical high-temperature, highly-reactive mixture and reaction zone. Due to incremental mixing, a gradual syngas cooling can be achieved. The layout of the design configuration is shown in Fig. 2(a).

2.3. Central steam inlet design This configuration represents the inverse arrangement of the steam and syngas compared to the design presented in Section 2.2. Steam is injected through a steam pipe as an internal flow into the hot syngas in the area of the quench inlet (see Fig. 2(b)). Thus, the hot syngas encloses the steam flow completely, which prevents steam from entering the quench chamber without coming into contact with the syngas. Therefore, this configuration could imply a better mixing behavior, but in return could lead to faster syngas cooling and reduced WGSR rates in the reaction zone. In general, a combination of both steam inlet designs is possible as well as a combination of steam and water injection.

3. Quench model setup 3.1. Process parameters

Fig. 1. Scheme of quench conversion reactor concept: 1 quench (syngas) inlet, 2 possible steam injection, 3 reaction zone, 4 waterbath surface, 5 waterbath, 6 recirculation zone, 7 cooled quench wall, 8 quench outlet(s), 9 syngas streamline, 10 underwater nozzles, and 11 quench wall cooling water inlet/outlet.

reaction zone. This configuration also ensures a gradual temperature decrease in the central downflow zone, which is favorable for the homogeneous WGSR reaction rate. Above the waterbath surface, the hot flow breaks up and is inevitably redirected. Additional steam from the waterbath is generated due to surface evaporation. In this connection, a higher partial pressure of H2O for additional CO conversion may be achieved and the mixture temperature level can be reduced automatically. The hot slag drips into the waterbath and cools down, which in return heats up the waterbath before the slag is discharged. After redirection, the converted syngas flows towards the quench exit and parts of the upstream flow are recirculated into the central downflow. The converted syngas is then removed through the quench outlet(s). The quench wall has to be cooled to prevent fouling and to achieve further quenching of the syngas. To realize low syngas exit temperatures or full quenching, the necessary additional amount of liquid water can be added through the underwater nozzles. In partial quench operation, exit temperatures above 1173 K can be utilized to recover the waste heat for high- and medium-pressure steam production. In the proposed concept, the quench conversion reactor can theoretically supply its own by means of subsequent medium-pressure steam production. Additional high-pressure steam which is produced can supply other parts of the process, offering the possibility of raising the net efficiency. However, as in radiant cooler systems, volatile alkali matter must be handled properly. 2.2. Radial steam inlet design In this configuration, ring nozzles for steam injection are intended to coat the hot syngas. The nozzles are therefore positioned around the

The three different syngases (A–C) are based on a 500 MW Siemens type dry coal-feed entrained-flow gasifier with a cooling screen [6,16]. A validated [13] ASPEN PLUS® model has been applied to fit the syngas composition to real measured data using the minimization of the Gibbs free energy at corresponding gasification operating pressures and temperatures together with the temperature approach method. The syngas composition may therefore deviate from thermodynamic equilibrium. The results are outlined in Table 2. Lignite has been utilized as a feedstock for gas A and bituminous coal for gas B, respectively. Gas B showed the highest exit temperature of 1923 K and likewise the lowest H2O/syngas(dry) ratio. Gas C was based on biomass (wood) and therefore contained the lowest amount of slag as well as the lowest gasifier exit temperature (Tout). The mass flows of syngas from the gasifier in the range of 38.41 to 45.47 kg/s were comparable with one another. The H2/CO ratio at the gasifier exit varied from barely 0.36 for gas A to 0.57 for gas C.

3.2. Numerical setup A 2D axis-symmetric setup using a block-structured grid with 32,000 cells was selected. The k-ω Shear Stress Transport (SST) turbulence model and Eddy Dissipation Concept (EDC) combustion model [23] together with In-Situ Adaptive Tabulation (ISAT) were chosen and solved with ANSYS Fluent® [12]. In addition to this, the P-1 radiation model was selected in combination with the cell-based version of weighted-sum-of-gray-gases (WSGG) model to calculate the absorption coefficient. The flow and thermodynamic gas properties were calculated depending on the pressure and temperature. Second order discretization was applied for all parameters. The convergence for all simulations was checked automatically by the solver to obtain comparable solutions. Here, the difference between two iteration residuals of four flow variables permanently had to come below a predefined criteria. The numerical model assumes that the quench waterbath surface features a fixed wall instead of a free surface. However, mass and heat transfer occur between the syngas and liquid water and therefore have to be taken into consideration. A User Defined Function (UDF) was written to calculate cell-temperature-dependent, specific evaporation rates based on film evaporation theory as well as the enthalpy of vaporization of a thin predefined bulk zone hBL (evaporation zone)

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Fig. 2. Schemes of quench concept steam inlet configurations: 1 — gasifier outlet, 2 — quench inlet, 3 — left: ring nozzle (right: steam pipe), 4 — steam flow, 5 — reaction zone, 6 — solid slag layer, 7 — liquid slag layer, 8 — wall cooling, and 9 — central downflow; gasifier outlet dimensions according to Siemens reactor standard design [16], steam injection dimension depends on constructional realization (uS ≫ uSG).

laminar or turbulent Nusselt number Nu for cross-flow over a flat plate is calculated [2]:

near the surface (see Eq. (3)) [2]:  
cp k T SG  T WB;surf  ; SH2O ¼ ln 1 þ Δhv cp  hBL

ð3Þ

where SH2O denotes the area-specific evaporating vapor mass flow from the free surface in kg/m2·s. The values for specific heat cp and cell temperatures TSG in the bulk zone are directly obtained from the numerical solution. The enthalpy of vaporization Δhv is pre-determined based on the corresponding gasification pressure. The temperature of the nearsurface water TWB,surf is set at a fixed value as the UDF assumes that there is a near-surface water temperature close to the evaporation temperature, due to heat input from the hot slag. The heat amount transferred from the slag to the gas phase is negligible [32]. A slag model therefore is not considered in the numerical setup. The heat transfer coefficient k is also calculated in the UDF. To do so, the Reynolds and Prandtl numbers along the waterbath surface are determined beforehand. Depending on the flow characteristics, either the

Table 2 Syngas characteristics at gasifier outlet [15,22,36].

  Nulam ¼ 0:664  Re1=2  Pr 1=3 Re b 5  105   Pr  2  Re ≥ 5  105 : Nuturb ¼ 0:037  Re0:8   1 þ 2:443  Re0:1 Pr 3  1

–

A

B

C

Feedstock Exit temperature Pressure _ SG m _ SL m

– K bar kg/s kg/s

Lignite 1723 40 45.47 2.47

Bit. coal 1923 42 39.15 1.79

Biomass 1773 30 38.41 0.2

Components CO H2 H2O CO2 N2 H2O/CO ratio H2O/syngas(dry) ratio H2/CO ratio

Vol.% Vol.% Vol.% Vol.% Vol.% mol/mol mol/mol mol/mol

54.11 19.44 14.18 10.89 1.37 0.262 0.165 0.36

58.79 23.9 6.47 3.5 7.34 0.11 0.069 0.41

49.21 27.95 15.49 7.14 0.16 0.316 0.183 0.57

ð5Þ

Finally, k can be determined using the following Eq. (6): k¼

λ ; L  Nu

ð6Þ

where λ denotes the thermal conductivity of the syngas and L the characteristic length, which in this case is represented by the reactor diameter. The cell-specific source of evaporation Sevap (W/m3) can be calculated using Eq. (7): _ evap  Sevap ¼ m

Gas

ð4Þ

Δhv ; hBL

ð7Þ

_ evap denotes the evaporating mass from the waterbath surface. where m The precise description of the high-temperature, high-pressure kinetics of homogeneous WGSR is crucial for the proposed concept. Most kinetics found in the literature are measured or created for low-temperature and low-pressure catalytic WGSR conditions (cf. Section 1). Global mechanisms, as published by Jones and Lindstedt [19] or Bradford [5], have advantages in terms of speed over detailed mechanisms, such as GRI3.0 [31], but may be lacking in terms of precision and reaction conditions, whereas detailed mechanisms are far more precise and use appropriate thermo and transport data [8]. For that reason, the detailed GRI3.0 reaction mechanism [30] was selected to perform the numerical simulations.

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3.3. Model validation Due to the black box character of commercial software, an intensive validation procedure is necessary. Since no experimental data are available for the quench conversion concept considered in the present work, the validation will be done using data on similar systems in the literature. Ni et al. [26] compared four different gas cooling systems using ANSYS Fluent and validated their numerical results against a radiative cooling system. While the system pressure and gas temperatures were comparable, their work neglects the impact of the water–gas shift reaction. Wang et al. [35] carried out a numerical study on a full water quench for an experimental, slurry-fed entrained-flow gasifier. Due to the large amount of water resulting from the slurry feed, the influence of the water–gas shift reaction is low compared to the present quench concept. Furthermore the operating conditions (3 bar, 950 K) are different from those in the quench concept considered here. In addition, [11], [18] and [31] investigated water–gas shift reactors. In these cases, the water–gas shift process was accelerated utilizing a catalytic bed, which results in faster reaction rates and significantly lower-temperature operating conditions. It can be concluded that no appropriate validation data are available in the literature with a dedicated focus on syngas cooling systems. As an alternative, it is possible to validate the numerical method against data in the literature for different reactive systems. The first validation step was carried out for the BYU gasifier, an entrained-flow coal gasifier operated at Brigham Young University [7]. The gasifier was tested with four different coals, ranking from lignite to bituminous coals. The reactor was operated at atmospheric pressure only, but it allowed the gas composition to be measured at several positions across the reactor. This is why the BYU gasifier has become a standard benchmark for modeling entrained-flow gasifiers [1,33]. The gas composition and temperature distribution within the reactor are based on the conversion behavior of the carbonaceous material and on the water–gas shift processes, especially in the gasification zone, which legitimates these data as a reference system for the desired quench reactor. In this section, the work focuses on Utah bituminous coal only, but different coals were assumed to behave similary. The numerical setup was similar to that described above, applying a reduced GRI3.0 [20] mechanism for the homogeneous reactions. For reasons of stability, the P-1 radiation model was used. Heterogeneous reaction kinetics and the particle size distribution were taken from [7]. Fig. 3 illustrates the temperature distribution in the reactor. The flame zone in the top part of the reactor and the gasification zone in the bottom part of the reactor can be seen clearly. Three positions along the reactor were selected for data analysis, starting with the flame zone, followed by the post-flame zone and subsequently the reforming zone. Fig. 4 shows the corresponding species distributions at these three positions. A comparison of the numerical results with the experimental data reveals a good accordance between experiments and numerics at all positions. The species distribution in the flame zone with the high gradients according to the oxidation reactions are correctly reproduced and the gas composition in the zone past the flame, which is significantly influenced by the homogeneous reforming reactions, shows a very good agreement.

As a second validation step, the high-pressure, high-temperature partial oxidation of natural gas was considered. In Richter et al. [28], experimental data are given (known as the Virtuhcon benchmark) for the non-catalytic reforming of natural gas based on the Freiberg test plant HP POX. The measurement data and the reactor and burner geometry are provided in such a way that they can be easily integrated in CFD models. Four different data points were selected (50, 60, and 70 bar at 1473 K, and 50 bar at a reactor temperature of 1673 K). The data points reflect conditions that intensively stress the CFD model. For instance, at 1473 K there are distinct non-equilibrium conditions at the reactor outlet, which is a challenging benchmark for the chemical mechanism used. The experimental data comprise gas compositions at inflow and outlet of the reactor, the different mass flow rates and temperature measurements along the reactor wall. In addition, the Optisos® device, which is an optical measurement system, was applied for the in-situ measurement of the flame shape, length, and stability in the reactor. Voloshchuk et al. [34] presented initial numerical results based on the Virtuhcon benchmark. The numerical model was comparable to the model used in this work, but with a reduced GRI3.0 [27] mechanism for the homogeneous reactions and the application of the P-1 radiation model. Fig. 5 illustrates the temperature distribution inside the reactor. The small hot region marks the natural gas flame zone, while slow, endothermic reformation reactions take place in the rest of the reactor. In addition, the temperature distribution based on the Optisos® measurements is shown in Fig. 5. Details of the optical measurements can be found in [28]. Table 3 compares the measured and calculated species compositions and flame characteristics for two selected cases. The experimental flame characteristics are measured for a fixed reference temperature [28] and the numerical results are calculated based on temperature and species gradients along the axis. It can be seen that there is a very good agreement between experiments and numerics for all measured characteristics, with a maximum difference in the species concentration of 3%, 4% in the flame length, and 3% in the flame width [34]. The CFD model used in this work is similar to the model for the BYU gasifier (without the particle model) and to the model used for the Virtuhcon benchmark (non-catalytic reforming of natural gas). Based on the validation steps, to sum up it can be said that the CFD model is suitable to reproduce the conversion processes in different high-temperature conversion systems. For the high-pressure partial oxidation, especially the conditions in the reforming zone are reasonably comparable to that for the quench system, and the influence of pressure on homogeneous WGSR kinetics can be neglected (see Section 1). For this reason it can be concluded that the CFD model is capable of reliably predicting the conversion process in the endothermic quench reactor discussed in this work. 4. Results The main objective is to provide a proof of concept for the proposed quench conversion reactor in terms of CO conversion and the syngas exit temperature. The comparison of investigated syngases and a sensitivity analysis will therefore be carried out in Section 4.1, applying the radial design. Furthermore, a comparison of the steam inlet designs will determine the best configuration (see Section 4.2). 4.1. Proof of concept

Fig. 3. Temperature distribution in the BYU gasifier in K.

_ S is varied in steps of 2 kg/s from 0 at The mass of injected steam m most 24 kg/s, which represents more than 50% of the corresponding syngas mass flows. In the quench concept, TS can take values between 550 and 850 K, as the evaporation temperature of water at a maximum pressure of 44 bars (gas B conditions, 2 bars excess, see Table 2) is 529 K and should not be undershot. The mean value of TS =700 K is used as the standard steam temperature for simulations. The specific wall cooling heat flux q_ W can take values from 10 to 100 kW/m2. The material

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203

Fig. 4. Validation of the numerical setup against data taken from [7]. Species distributions along reactor axis. Solid lines: CFD; symbols: experimental data. Utah bituminous coal.

properties of water evaporating from the waterbath surface are set to the syngas-specific process conditions. 4.1.1. CO conversion The main results for syngas variation are shown in Fig. 6 and Table 4. The plots outline the results for CO conversion and the H2/CO ratio in terms of injected steam as well as the H2Otot/CO ratio (kg/kg). To obtain this ratio, the sum of the initial syngas H2O content, injected steam and waterbath evaporation rate in relation to the initial syngas CO content was determined. The CO conversion increases with additional amounts of steam and shows individual peak values for the different syngases (see Fig. 6(a)).

_ S = 10–11 kg/s. However, the slope of COconv is nearly similar up to m The CO conversion curves of gases A and C are comparable because of the corresponding syngas temperatures and H2O/syngas(dry) ratios, whereas gas B exceeds the other gases and has values above 25%. The CO conversion potential therefore is influenced by the syngas temperature and composition as described in detail in the following. CO conversion by means of the H2Otot/CO ratio is illustrated in Fig. 6(b). The results indicate that the syngas properties have an impact on the quench conversion process. In general, the reaction rate of WGSR in the gasifier is high and the CO conversion strongly depends on the syngas composition and residence time. Dry and hot syngases, such as gas B, thus will show constantly higher CO conversion rates compared to syngases with higher initial partial pressure of water. Nevertheless, all the investigated syngases show peak values in a relatively small H2Otot/CO range of 0.67 to 0.78 (see also Fig. 6(d)). The maximum achievable H2/CO ratio level depends on the initial syngas composition as illustrated in Fig. 6(c). Gas B shows the maximum increase (127.7%) and gas A the lowest (71.9%) compared to the initial syngas characteristics. As already outlined, gas B also has the highest CO conversion potential, but coincidentally shows a relatively

Table 3 Comparison of measured and calculated reactor outlet compositions and flame characteristics for two different data points [34].

Fig. 5. Detailed view of the flame zone in the HP POX (operating pressure 50 bar(g), temperature 1373 K). Fixed temperature field in CFD and corresponding results from optical measurements (small window). Pictures taken from [34].

Quantity

–

Experiment

Numerics

Experiment

Numerics

Temperature Pressure H2 CO CO2 H2O Flame length Flame width

K bar(g) Vol.% Vol.% Vol.% Vol.% mm mm

1473 50 48.27 23.79 4.19 19.33 265 41

1473 50 45.80 23.96 3.62 21.36 267 38

1673 50 48.06 25.61 3.89 21.71 – –

1673 50 47.74 26.27 3.19 22.25 311 30

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K. Uebel et al. / Fuel Processing Technology 148 (2016) 198–208

Fig. 6. Results for CO conversion and H2/CO ratio in terms of injected steam and total H2O/CO ratio.

low initial H2/CO ratio of about 0.41 compared to 0.57 for gas C. However, the high CO conversion rates of gas B cannot overcome the low initial H2/CO ratio and therefore gas C has the maximum H2/CO ratio for all three syngases (cf. Table 4). In addition to this, the peak values for gas A and C can be achieved with the same amount of injected steam, as the H2O/syngas(dry) ratio of these gases are nearly the same. Fig. 6(d) again outlines the conclusions found before. The syngases have different initial H2/CO and H2Otot/CO ratios and therefore show varying maximum H2/CO ratios in a very small H2Otot/CO range as well as comparable curve characteristics. The syngas (exit) temperature plays an important role in the quench conversion concept. Fig. 7 sheds light on this statement. In general, the syngas exit temperature decreases with as the mass flow of injected steam increases. Regarding the intersection between syngas exit temperatures at maximum H2/CO ratios, only a small temperature difference of 42 K can be obtained in the range of 1423 to 1465 K. In conclusion, the highest H2/CO ratios can be achieved in this particular small exit temperature range. The relation between the increase of reaction rate with temperature and the decrease of H2/CO ratio due to chemical equilibrium results in maximum CO conversion within these limits for the investigated geometry.

Table 4 Major results for syngas variation. Syngas

A

B

C

Intrinsic CO conversion (%) Max. CO conversion (%) Max. H2/CO ratio (mol/mol) Max. H2 content (Vol.%) Min./max. SN (mol/mol) Max. H2O utilization (%)

2.3 18.6 0.69 22.15 0.2/0.39 21.05

2.03 27.0 0.93 25.58 0.37/0.87 25.61

2.27 20.8 0.98 28.28 0.49/0.96 18.4

4.1.2. Syngas composition at quench exit In many applications, such as pressure swing absorption or ammonia production, only the maximization of H2 content in the syngas is favorable. The optimal H2Otot/CO range for the maximum H2 content (0.47–0.58) on the one hand and H2/CO ratio (0.67–0.78) on the other hand take different value ranges. Here, high H2 values can be obtained using less steam than the required for maximum CO conversion or H2/CO ratio, respectively. As the initial H2 content of gas C is already high, the amount at the quench exit does not change noticeably for a wide range of additional injected steam values (0–8 kg/s or H2Otot/ CO = 0.24–0.55) and subsequently even decreases. However, the previously defined H2O tot /CO range during which the other two gases obtain high H2 values is also valid for gas C. The maximum attainable H2 content is outlined in Table 4. 4.1.3. Waterbath evaporation The use of the waterbath to deflect the flow and provide additional steam as well as a syngas cooling source is an important feature of the quench concept. The evaporation rates obtained for gases A to C generally take values between 0.93 and 1 kg/s over a wide range of the total exit mass flow, as illustrated in Fig. 8. It can be concluded that for the investigated quench dimensions, the influence of the syngas mass flow and syngas mixture temperature on the waterbath evaporation rate is low and dominated by heat transfer. The evaporating steam constantly contributes to CO conversion in the lower part of the quench chamber. 4.1.4. Steam utilization In order to evaluate the quench concept performance, the steam utilization rate for gases A–C has to be determined. The results are outlined in Fig. 9. The syngas composition received from the gasifier may be not in thermodynamic equilibrium, as mentioned before. When no additional steam is injected, only the syngas H2O content and the waterbath evaporation rate are taken into account and are compared against the

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Fig. 7. Coherence between H2/CO ratio and syngas exit temperature.

final syngas H2O content at the quench exit. The values corresponding _ S = 0 kg/s in Fig. 9 represent the intrinsic H2O utilization rates. As to m expected, dry syngases, such as gas B, show high starting values (compare H2O/syngas(dry) ratio in Table 2). With additionally added steam, the H2O utilization rates for all three syngases rise to maximum values between 18.4 and 25.6%. Comparing the results for CO conversion, gas B (with the lowest initial water content) shows the best utilization rate, followed by gases A and C, respectively. In addition to this, the curve of gas B forms a homogeneous profile over a wide range with values close to a 25% steam utilization rate. The other gases perform comparably at a lower level, but show an earlier decrease for higher amounts of steam, again driven by the initial H2O/syngas(dry) ratio. In Table 4 the intrinsic CO conversion represents the conversion rate without steam injection in the upper part of the quench. Thus, these values only contain the waterbath evaporation rate of about 1 kg/s and syngas H2O content, which contribute to WGSR in this context. The stoichiometric number SN has the same curve shape as the H2/CO ratio, and a value of nearly 1 was obtained for gas C with the chosen setup.

cooling heat flux q_ W and the temperature of injected steam TS are two additional process parameters. The influence of these parameters has been investigated using the setup with the maximum obtained CO con_ S = 14 kg/s). The results are shown in Fig. 10. version for syngas A (m Altering TS directly influences the mixture temperature and therefore the reaction rate of WGSR or CO conversion, especially in the reaction zone. Higher temperatures for steam obviously increase the syngas exit temperatures. Fig. 11 illustrates the impact of the steam temperature on the H2/CO ratio in the quench chamber. In Fig. 10, the H2/CO ratio grows logarithmically as the steam temperatures rise, whereas Tout advances almost linear and only moderately. Thus, higher steam temperatures technically should be preferred. Moreover, there is a greater chance of producing steam with subsequent heat recovery steam generation methods for self-supply, if technical issues such as fouling due to alkalis can be prevented. As expected, the quench exit temperatures decrease with increasing q_ W . Higher wall cooling supports the syngas quenching process, but in contrast leads to a drop in CO conversion, although the gradient of steam temperature variation is significantly higher. The influence of

4.1.5. Analysis of process parameters sensitivity The results for H2/CO ratios are highly sensitive to the injected amount of steam and H2O/CO ratio, respectively. The specific wall

Fig. 8. Waterbath evaporation for investigated syngases.

Fig. 9. H2O utilization rates for gases A–C linked to steam mass flow injection.

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Fig. 10. Sensitivity analysis of radial design and syngas A using maximum CO _ S = 14 kg/s, H 2 /CO ratio = 0.69) and exit temperature T out . The conversion ( m Figure shows the influence of process parameter variation in terms of injected steam temperature TS and specific wall cooling heat flux q_ W:

the parameter variation on the waterbath evaporation rate is relatively narrow. It must therefore be concluded that the two analyzed parameters have a negative influence on one another. This leads to a trade-off situation, as a useful combination has to be found to achieve sufficient syngas cooling as well as noticeable CO conversion at the same time.

4.2. Comparison of steam inlet designs The two proposed steam inlet designs are checked against each other using gas A and the identical numerical setup. In general, both designs show comparable flow and CO conversion behavior, although the central design reaction zone is inverse and relative gas velocities of syngas to steam have changed. No major differences in the H2 /CO ratio and T out can be determined. Both designs show a _ S = 14 kg/s. However, the radial design obtains H2/CO peak ratio for m slightly higher values (max = 0.69) compared to the central design (max = 0.67). With regard to the waterbath evaporation rate, the central designs reveal continuously increasing rates up to nearly 1.35 kg/s, in contrast to the radial design (compare gas A in Fig. 8). This effect corresponds to the changing flow characteristics due to the increasing and in generally higher nozzle and central flow velocities (steam pipe configuration).

Both designs show an increase in the H2/CO ratio for high steam in_ S ≫ 20 kg/s, cf. Fig. 6). Here, the central design reveals an earjection (m lier and sharper increase. Fig. 12 compares the design-specific net _ S = 22.8 kg/s, reaction rates of H2 at different reactor heights for m which represents 50% of the syngas mass flow, whereas Fig. 13 shows the differences in the corresponding temperature profile. The reaction zone of the central steam pipe configuration is more radially extended and shows enhanced reaction rates for the water–gas shift compared to the radial design as well as configurations with lower steam injection. However, the reaction rates just above the waterbath surface are comparably low for both inlet designs. The inverse behavior can be observed in Fig. 13, which underlines the strong coherence between temperature and reaction rates and thus CO conversion. The CO conversion rates continuously decline outside the reaction zone for steam mass flows above peak values of the H2/CO ratio. In ad_ S values above 20 kg/s. In these dition to this, Tout is almost equal for m cases, the increased CO conversion therefore must to be caused only by the more turbulent flow field and mixing in the reaction zone. The overall lower temperature level in the central design in addition causes the lower CO conversion (cf. Fig. 13) compared to the radial design. Moreover, due to high amounts of steam and therefore the low syngas temperatures outside the reaction zone, quenching occurs and thus prevents further CO conversion.

5. Discussion From the technical point of view, the relatively high syngas exit temperatures and the coal ash ingredients lead us to expect alkali condensation in the subsequent heat recovery systems. Also, a heat balance of the gas treatment process chain should be conducted to identify possible efficiency advantages and to verify the ability of the quench concept to provide the system with steam. The investigation showed that the proposed concept requires significant amounts of superheated steam for maximum CO conversion. Higher steam temperatures additionally enhanced the obtained H2/CO ratios. However, for economic reasons, a TS above a certain level may not be reasonable as the steam production costs also raise. Depending on the specific steam production costs of the gasification facility, an appropriate temperature level should be selected, if it is not possible to produce enough steam. This work did not consider the impact of the underwater nozzles. However, the injection of water droplets into the quench system will have a huge cooling effect and has to be investigated in future works. Two aspects might be of interest: first, the steam substitution potential in terms of CO conversion and second, alkali removal effects from the gas phase.

Fig. 11. Contour of H2/CO ratio level and distribution for different steam temperatures.

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_ S = 22.8 kg/s) at relative quench chamber heights/diameters and different steam inlet designs. Fig. 12. Comparison of H2 net reaction rate (RR) for case (m

The two inlet designs showed comparable performance. However, the steam pipe is exposed to more difficult boundary conditions, such as liquid slag and high gas velocities. Hence, the radial design might be favored. This study was mainly focused on concept verification. Additional uninvestigated technical and economical boundary conditions may restrict the presented results. This has to be outlined in future works. As the results were only obtained for one design quench configuration, there is a high potential for further systematic investigations of system performance. In order to find the maximum attainable H2/CO ratio or other objectives, an optimization process could provide more detailed information.

6. Conclusions A new quench concept with integrated CO conversion has been investigated using CFD models. The selected numerical model was successfully validated against measurements as well as numerical data from two bench-scale gasifiers. The CO conversion potential of three syngases based on lignite, coal and biomass feedstock and gasifier exit temperatures were considered to verify the concept. In all cases, significant CO conversion rates up to 27% were achieved and therefore a minimum increase of 72% to a maximum of 128% in terms of the H2/CO ratio could be achieved. The results indicate that maximum H2/CO ratios have been attained in a close range of the H2Otot/CO ratio, which requires different

_ S = 22.8 kg/s) at relative quench chamber heights/diameters and different steam inlet designs. Fig. 13. Temperature profiles for case (m

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syngas-specific amounts of additional water. A sensitivity analysis revealed the high influence of the steam inlet temperature and wall cooling on CO conversion. The two proposed and investigated inlet designs show comparable performances. High potential for further quench conversion concept design improvements has been identified and discussed. Acknowledgments We like to thank the Federal Ministry for Economic Affairs and Energy for supporting this work in the framework of HotVeGas (project no. 0327773G). Furthermore, the authors acknowledge the financial support by the European Social Fund (ESF) and the Free State of Saxony in the framework of ProVirt (project no. 100231952). We also thank Kristin Boblenz and Yury Voloshchuk for their valuable contribution to this work. References [1] N. Abani, A.F. Ghoniem, Large eddy simulations of coal gasification in an entrained flow gasifier, Fuel 104 (2013) 664–680. [2] H.D. Baehr, K. Stephan, Heat and mass transfer, 2011. [3] D.A. Bell, B.F. Towler, M. Fan, Coal Gasification and its Applications, Elsevier, 2011. [4] K. Boblenz, K. Uebel, B. Meyer, Verfahren und Vorrichtung zur Teilkonvertierung von Rohgasen der Flugstromvergasung, DE 10 2015 219 455.5, 03.26. 2015. [5] B. Bradford, The water–gas reaction in low-pressure explosions, J. Chem. Soc. (1933) 1557–1563. [6] R.W. Breault, Gasification processes old and new: A basic review of the major technologies, Energies 3 (2010) 216–240. [7] B. Brown, L. Smoot, P. Smith, P. Hedman, Measurement and prediction of entrainedflow gasification processes, AICHE J. 34 (1988) 435–446. [8] A. Burcat, Thermochemical data for combustion calculations, Combustion Chemistry, Springer 1984, pp. 455–473. [9] F. Bustamante, R. Enick, A. Cugini, R. Killmeyer, B. Howard, K. Rothenberger, M. Ciocco, B. Morreale, S. Chattopadhyay, S. Shi, High-temperature kinetics of the homogeneous reverse water–gas shift reaction, AICHE J. 50 (2004) 1028–1041. [10] F. Bustamante, R. Enick, R. Killmeyer, B. Howard, K. Rothenberger, A. Cugini, B. Morreale, M. Ciocco, Uncatalyzed and wall-catalyzed forward water–gas shift reaction kinetics, AICHE J. 51 (2005) 1440–1454. [11] R. Chein, Y. Chen, C. Yu, J. Chung, Modeling and simulation of H2S effect in hightemperature water–gas shift reaction using coal-derived syngas, Int. J. Hydrog. Energy (2015) 8051–8061. [12] Fluent, ANSYS Fluent V15.0 – Commercially Available CFD Software Package Based On the Finite Volume Method, Ansys Inc., 2014 (URL www.ansys.com). [13] M. Gräbner, Modeling-Based Evaluation of Coal Gasification Processes for High-Ash Coal(Ph.D. thesis) Technische Universität Bergakademie Freiberg, Freiberg, 2012. [14] M. Gräbner, Industrial Coal Gasification Technologies Covering Baseline and HighAsh Coal, Wiley-VCH, Weinheim, 2014. [15] U. Günther, TEIMAB: SFG-Vergasungsreaktor mit Teil-Quench und Abhitzedampferzeuger (Project number: 0327797A), 2013.

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