Operating characteristics of a transpiring-wall SCWO reactor with a hydrothermal flame as internal heat source

J. of Supercritical Fluids 34 (2005) 35–50

Operating characteristics of a transpiring-wall SCWO reactor with a hydrothermal flame as internal heat source B. Wellig1 , K. Lieball2 , Ph. Rudolf von Rohr∗ Institute of Process Engineering, Swiss Federal Institute of Technology (ETH), CH-8092 Zurich, Switzerland Received in revised form 18 June 2004; accepted 21 July 2004

Abstract A novel supercritical water oxidation (SCWO) system with a transpiring-wall reactor (TWR) containing a hydrothermal flame as internal heat source has been investigated. This configuration may overcome the problems of corrosion and plugging of reaction vessels and other components. In the TWR, the wall contact of supercritical reaction media is prevented by fluid dynamical means. Subcritical water permanently flows through porous, non-load-bearing transpiring-wall elements and forms a “film” on the inner wall. Precipitated salt particles do not reach the wall because they are either re-dissolved in this film or are swept away by the transpiring water. The waste water reaches reaction temperature without any wall contact, i.e., by mixing with the combustion products of the hydrothermal flame. In the present work, the apparatus design, processing features, and experimental results are presented. Tests with salt-free water–methanol mixtures have been conducted at 250 bar. The influence of the intensity (i.e., mass flux) and temperature of transpiration on the reactor performance has been examined. The temperatures near the transpiring wall remain subcritical which is a quantitative measure for the wall protection from salt adherence. Methanol conversions of higher than 99% were obtained even at comparatively low transpiring water temperatures (125–250 ◦ C). At typical operating conditions, the mass flow ratio between the total transpiration flow and the core flow is around 65%. Further, it was found that natural convection effects are not negligible in a transpiring-wall reactor for SCWO. © 2004 Elsevier B.V. All rights reserved. Keywords: Supercritical water oxidation; Hydrothermal oxidation; Hydrothermal flame; Transpiring-wall reactor

1. Introduction Supercritical fluids (SCFs) possess unique physicochemical properties, which make them attractive as media for chemical reactions and separation processes. In the last decades, the number of applications has increased continuously [1,2]. An emerging technology in the environmental field is the oxidative destruction of organic wastes in supercritical water, which is known as supercritical water oxidation (SCWO) or hydrothermal oxidation (HTO).

∗

Corresponding author. Tel.: +41 1 632 24 88; fax: +41 1 632 13 25. E-mail addresses: [email protected] (Ph. Rudolf von Rohr); [email protected] (B. Wellig); [email protected] (K. Lieball) 1 Present address: Ernst Basler + Partner AG, Zollikerstr. 65, CH-8702 Zollikon, Switzerland. 2 Present address: Von Roll Umwelttechnik AG, Hard-turm-str. 133, CH8037 Zurich, Switzerland. 0896-8446/$ – see front matter © 2004 Elsevier B.V. All rights reserved. doi:10.1016/j.supflu.2004.07.003

Near and above the critical point of pure water (pc = 221 bar, Tc = 374 ◦ C), the thermo-physical properties of water are quite different from those at ambient conditions and the solvent and reaction characteristics drastically change [3,4]. SCW is much less polar than liquid water [5] and therefore becomes a good solvent for non-polar organic compounds and gases such as oxygen, nitrogen or carbon dioxide, e.g. [6–8]. Contrariwise, the solubility of inorganic salts in SCW is extremely low [9–11]. In the single-phase mixture of organic compounds and oxygen in SCW (at least under salt-free conditions), interfacial mass transfer resistances are absent. The single-phase mixture and the considerable high reaction temperatures enable short residence times and small reactor volumes (i.e., high space–time yields). SCWO is an alternative technology for the destruction of a variety of hazardous waste streams, in which the concentration of organics is generally below 20 wt.%. Typical operating conditions in the reactor types studied up to now are

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around 250 bar and 450–650 ◦ C. Under these conditions, the total oxidation of organic compounds takes place rapidly and the conversions are close to unity at residence times of approximately 1 min or less. Extensive reviews of fundamental and engineering aspects of SCWO have been published by several authors, e.g. [12–16]. SCWO, however, has two technical problems, which appear to be inevitable. First, the presence of reactive ions such as Cl− , F− , H3 O+ , in combination with SCW and oxygen can lead to severe corrosion of the high-pressure reactor vessel and process equipment [17–21]. Secondly, plugging of the reactor, preheat and cooling section and other components can occur due to precipitation of sticky salts and solids [22–25]. These two problems have inhibited the commercial breakthrough of the SCWO process up to date. The development must lead to reaction systems, which enable processing aggressive waste streams, which contain salts and solids or in which salts are formed during neutralization. A two-fold approach can be taken to overcome these difficulties. First, reactor designs with wall boundary layer control can be developed. The leading idea of such concepts is to use fluid boundaries to keep the hot and corrosive reaction mixture away from the pressure-bearing reactor walls. Secondly, extensive material tests can identify construction materials capable of withstanding the SCWO environment (e.g., for construction double-tube reactors [26– 28]). However, it must be emphasized that plugging cannot be prevented by better materials which would withstand firm corrosion. Therefore, solutions must be found using fluid dynamical and process engineering means. Beside other approaches, the concept of the transpiring-wall reactor (TWR) is a promising solution. The aim of this work is to understand the transport phenomena occurring in such a reactor.

pressure-bearing walls from the hot aggressive mixture and simultaneously inhibits the deposition of precipitated salts. Both reactors allow photographic pictures to be taken of the hydrothermal flame at a working pressure of 250 bar. First pictures of continuous methane and methanol flames were published in 1995 [35] (see also Refs. [37,41–43]). In the second-generation reactor, optical access to the full length of the flame was provided including the stabilization zone in the immediate vicinity of the core tube tip of the burner. In both reactors, a variety of burner designs were tested concerning flame stability (i.e., extinction limits), destruction efficiency (i.e., conversion ratios in the flame), as well as wall compatibility (i.e., corrosion and plugging). Pure oxygen was used as oxidizer and ignition occurred by heating-up of the reactants to auto-ignition temperature. Stable burning of the hydrothermal flame was achieved even with reactor inlet temperatures of the feed stream well into the subcritical range. They could be lowered to about 250 ◦ C with methane–water and below 100 ◦ C with methanol–water mixtures, depending on the concentration of the organic compound, burner geometry and other operating parameters. No deposition was found in the experiments with salt containing feed streams (Na2 SO2 , NaCl), but some non-pressurebearing parts were damaged by corrosion after a certain operating time. These investigations led to the conclusion that a reactor containing a hydrothermal flame as internal heat source, multiple waste water and oxidizer inlet and transpiring-wall cooling will be the main components of a future SCWO facility [42]. Therefore, a novel tubular reactor type with a hydrothermal burner and transpiring-wall technology has been developed and investigated during the last few years [43–48].

2. Previous work in the field of SCWO at ETH Zurich

3. Fundamentals of the novel transpiring-wall reactor

ETH Zurich has a long tradition of research in the field of supercritical water oxidation. In 1992, the first project was launched and new ways of processing and apparatus design were evaluated. Finally, a decision was reached to pursue SCWO in the thermal regime of a flame. At that time, the possibility of generation of so-called “hydrothermal flames” in semi-batch reactors was a well-known fact. Franck and coworkers [29–32] and a research group from Sandia Nat. Labs. [33,34] built reaction autoclaves with optical access for investigation of such flames. However, the goal of the ETH project was to examine continuously operated hydrothermal diffusion flames with liquid feed injection. A corresponding pilot plant with a feed rate of about 10 kg/h was designed and constructed. In this facility, two different high-pressure reactors containing a wall-cooled hydrothermal burner (WCHB) were examined [35–42]. The leading idea of this burner type is to confine the reaction zone (i.e., the flame) in a layer of cooling water, which protects the

The novel transpiring-wall reactor combines the advantages of the hydrothermal flame with a well-known wall protection concept. A schematic drawing of the TWR is depicted in Fig. 1. Its characteristic feature is the prevention of any wall contact of the hot, corrosive and/or particle containing reaction mixture by fluid dynamical means. For this purpose, porous, non-load-bearing cylindrical transpiring-wall elements are inserted to impose an inward-directed radial velocity component on the established flow field near the inner reactor wall. This velocity component provides convective mass transport to oppose the diffusion mechanisms that typically lead to the undesired deposition of salt particles precipitated from the reaction mixture inside the reactor. The particles are supposed not to reach the wall. They are either redissolved in the “cool” film (compared to the temperatures in the reaction zone) or swept away from the wall by the momentum of the transpiring water. The transpiring water flow simultaneously protects the pressure-bearing

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Fig. 1. Concept of the transpiring-wall reactor (adapted from Refs. [45,47]): (a) hydrothermal flame used as internal heat source to reach reaction temperature; (b) mixing zone of waste water and oxygen (omitted in this study); (c) reaction zone in which the waste water is oxidized; (d) protecting “film” formed by the cold and dense transpiring water; (e) quenched reacted mixture.

reactor wall from the corrosive reaction mixture inside the reactor. The core stream, which is fed axially, mixes continuously with the radially incoming transpiring water flows and hence increases over the reactor length. Thus, the temperature and the concentrations of each species in the core stream are changing in function of the mass flux ratio of the transpiration stream to the core stream and their composition and temperatures. The flow direction from top to bottom shall inhibit natural convection emerging from the density difference between the cold outlet and the hot reaction zone. The temperature in the reaction zone is aimed to be kept at a value high enough for complete conversion. The hydrothermal flame allows that the reactants reach reaction temperature without any wall contact (mixing of hot combustion products with cold waste water and oxygen) and hence low inlet temperatures of the waste water are possible (in fact the waste-water flow is not preheated, see Figs. 2 and 5). The wall protection concept is similar to concepts from other labs dealing with salt containing waste waters (e.g., U.S. patents [49–54], Ahluwahlia et al. [55], McGuinness [56], Sandia Nat. Lab. Reports [57–59], Crooker et al. [60] (and literature cited in Ref. [60]), Abeln et al. [61,62], Fauvel et al. [63]), yet the use of a separate hydrothermal flame as an internal heat source is novel in SCWO reactors.

Fig. 2. Sectional drawing of the transpiring-wall reactor and pilot plant configuration used in this work. Descriptions of mass flow rate and reactor inlet temperature measurements are provided in Table 1. Further measuring points: Tbn = fuel temperature in the burner nozzle (some millimeters above the tip of the burner tube); Tan1 = Ox1 temperature in the inner annular gap; Tan2 = WW temperature in the middle annular gap; Tan3 = Ox2 temperature in the outer annular gap; Tsi = TW temperature in the annular gap of section i (i = 1–3); Tout = temperature of reactor effluent.

4. Objectives Process and chemical engineering literature in this field usually does not provide detailed specifications of geometric data and operating conditions, thus making experimental results difficult to reproduce (e.g., Refs. [60,61,63]). The present work deals precisely with this issue. The aim is to promote a basic understanding of the transport phenomena occurring in such a reactor with the help of measurements under clearly defined operating conditions and show the feasibility of the proposed reactor concept. However, the objective is not to perform extensive degradation measurements with various model substances or to elucidate chemical aspects (e.g., reaction kinetics and pathways). The experiments were carried out under a number of different flow conditions using salt-free artificial fuel and wastewater mixtures. Especially, the influence of the temperature and intensity of the transpiration on the reactor behavior was

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examined. Also, the use of a separate hydrothermal flame as internal heat source was tested for the first time. It is shown that high conversions can be achieved without preheating of the waste-water stream and within short residence times (in the order of seconds). Furthermore, the proposed wall protection concept has to be proved by measurement of the temperature field in the reactor. The temperature of all load-bearing reactor walls and at the inner surface of the transpiring-wall tube must have sub-critical values. Therefore, temperature profiles were measured using special intermediate rings between the individual transpiring-wall elements. Finally, these experiments provided data for the validation of turbulent reactive computational fluid dynamic (CFD) models [44,46,48].

5. Experimental set-up and procedure 5.1. Reactor design A sectional drawing of the transpiring-wall reactor including all flows entering and leaving the reactor is represented in Fig. 2. The abbreviations of all flows and the measuring points shown in this figure are summarized in Table 1. The reactor vessel is made of Alloy 625 (Ernst Haage Apparatebau, Mühlheim a. d. Ruhr, Germany). It has an inner diameter of 34 mm and the length of the transpiration zone is 375 mm. The reactor can be operated up to a pressure of 420 bar. The burst pressure is 600 bar and the maximum permissible wall temperature is 600 ◦ C. The experiments made use of a burner configuration similar to the hydrothermal flame experiments [47]. A sectional drawing of the utilized coaxial burner is represented in Fig. 3. The air-gap insulated burner tube (nozzle i.d./o.d. 2.1/7 mm) is introduced into the top flange from the very top. The combustion chamber tube (i.d./o.d. 9/12 mm) is screwed into the top flange from below. The length of the combustion chamber is 50 mm. The outer insert (i.d/o.d. 15/18 mm), which separates the waste water and oxygen 2 stream, is screwed into the main body from above. The mixing zone of waste water and oxygen 2 (zone b in Fig. 1) was omitted in this first series of experiments. All nonload-bearing inserts are made of nickel base alloys. Detailed descriptions of the design of the high-pressure vessel and the hydrothermal burner equipment can be found elsewhere [47]. The transpiring-wall tube is made out of high-porous sintered Alloy 625 (GKN Sinter Metals Filters, Radevormwald, Germany). It is divided into four sections. Transpiring water is fed in the first three sections (TW1, TW2, TW3), while cooling water is fed in the lowest section (CW). The sinter tubes of the transpiration sections have a porosity of 17% (equivalent laminar diameter of 3 ␮m), while the lowest tube has a porosity of 21% (equivalent laminar diameter of 5 ␮m). Further, the lowest tube is furnished with six additional holes of 1 mm diameter at its circumference for injecting cooling water. Fig. 4 exemplifies a detailed sectional drawing of one transpiring-wall section (TW3). The selection of a split de-

Fig. 3. Sectional drawing of the coaxial hydrothermal burner (for detailed information see Ref. [47]).

sign provides notable advantages: (1) The transpiring water mass flow rate and temperature in each section can be controlled separately. Hence, the adjustment of different transpiration intensities and temperatures over the reactor length is possible. (2) The individual transpiring-wall elements are easy to replace. It is possible to examine configurations using TW elements with different porosities. (3) The three intermediate rings allow a lead-through of thermo-couples into the transpiration zone. With this design, radial temperature profiles at three different heights can be measured (measurement planes P2, P3, P4; plane P1 was not used in this study). At each measurement plane, the temperatures inside the TW tube is measured at four points over the radius: r = 0 (centerline), 3, 6, and 9 mm. The thermo-couples are inserted through the intermediate Alloy 625 rings (Pos. 3 in Fig. 4), while two metal C-ring seals (Pos. 4, Advanced Products, N. V., Boom, Belgium) are mounted on the outside in each case. This fully isolates the individual annular clearances from each other and prevents water from flowing into the transpiration zone via the measurement bores. The thin discs (Pos. 5) prevent the seals from slipping during assembly and dismantling. The intermediate rings must have a high dimensional accuracy; for example, the internal diameters of the sinter tubes and the intermediate rings must be identical to prevent disturbances in flow. The temperatures Ts1 –Ts4 (s means “section”) in the annular clearances (i.e., outer side of the transpiringwall elements) are not measured on the same plane as the profiles. They are measured at the inlet level of the transpir-

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Fig. 4. Detailed sectional drawing of the transpiration section TW3: (1) main part of the high-pressure reactor vessel (Alloy 625); (2) transpiring-wall element (high-porous sintered Alloy 625 tube); (3) intermediate ring (Alloy 625); (4) metal C-ring seal; (5) disc (stainless steel). Dimensions in mm.

ing water flows (opposite side of transpiring water inlet, see Figs. 2 and 4). The disadvantage of this measurement set-up is the disturbance in the flow through the integrated thermo-couples in the transpiration zone. At present, this must be accepted in the apparent absence of any other reasonably straightforward measuring technique. It is desirable for this zone to allow visual inspection, although this can only be achieved through extremely complicated design measures. 5.2. Pilot plant A simplified scheme of the pilot plant is represented in Fig. 5. The artificial fuel and waste-water mixtures are supplied by plunger metering pumps and pure oxygen by a special diaphragm-type compressor. Transpiring and cooling water is pumped with a radial piston pump. The cooling water is fed into the system at the bottom of the reactor and before the pressure control valve, respectively. Pulsation dampers

are installed in all liquid lines. The fuel mixture and oxygen for the hydrothermal flame and the transpiring water flows are preheated with special electric resistance pre-heaters [40,47]. With these pre-heaters, exactly adjustable fluid outlet temperatures up to 600 ◦ C can be quickly achieved. The flow rates of all streams are adjusted manually and they are measured by means of Coriolis-type mass flow meters. Calibrations resulted in error values lower than 1% of the actual flow rate value. The reactor inlet temperatures of the preheated flows are controlled using PID controllers. Fluid temperatures in the high-pressure lines (special Alloy 625 tees) and inside the reactor are measured by means of type K thermo-couples (measurement of temperature profiles in TW section: accuracy class 1 thermo-couples: ±1.5 ◦ C for T < 375 ◦ C, ±0.4% for T > 333 ◦ C; all other thermocouples: accuracy class 2: ±2.5 ◦ C for T < 333 ◦ C, ±0.75% for T > 333 ◦ C). After quenching, the pressure is measured with a pressure transducer (accuracy better than 0.2% of the full scale value 400 bar, i.e., ±0.8 bar) and reduced in one

Table 1 Summary of all flows entering the transpiring-wall reactor (see Figs. 2 and 5) and their mass flow rates Fi and inlet temperatures Ti Flow

Index

Mass flow rate

Inlet temperature

Fuel (water–MeOH) Ox 1 = oxygen 1 WW = wastewater (water–MeOH) Ox 2 = oxygen 2 TW 1 = transpiring water 1 TW 2 = transpiring water 2 TW 3 = transpiring water 3 CW = cooling water

f ox1 ww ox2 tw1 tw2 tw3 cw

Ff = 1.5 g/s Fox1 ≈ 1.5γFf wf Fww = 1.0 g/s Fox2 = 0.2–0.25 g/s Ftw1 = κ1 Fb Atw1 /Ab Ftw2 = κ2 Fb Atw2 /Ab Ftw3 = κ3 Fb Atw3 /Ab Fcw = 60–70 g/s

Tf = 200–350 ◦ C Tox1 = 300–400 ◦ C Tww = 20 ◦ C Tox2 = 30–35 ◦ C Ttw1 = 40–250 ◦ C Ttw2 = 40–250 ◦ C Ttw3 = 40–250 ◦ C Tcw = 28–35 ◦ C

Nomenclature: γ is stoichiometric oxygen excess for hydrothermal flame, wf is the methanol mass fraction in the fuel flow, www is methanol mass fraction in the waste-water flow, κi is transpiration intensity in section i, Atwi is inner shell surface of the sinter tube of section i, Ab is inner circular area, Fb is mass flow rate of bulk stream (Fb = Ff + Fox1 + Fww + Fox2 ).

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Fig. 5. Schematic representation of the SCWO pilot plant. FI: flow indication, PI: pressure indication, TI: temperature indication, TIC: temperature indication control, PIC: pressure indication control. A detailed process and instrumentation drawing can be found in Ref. [47].

step with the electro-pneumatic control valve and the effluent is discharged. A PI controller is used for system pressure control resulting in pressure fluctuations smaller than 0.5% of the set-point (250 bar). Samples are taken directly after the gas/liquid separator and analyzed by a gas chromatograph equipped with a flame ionization detector (GC-FID, Hewlett Packard, 5890 Series II). Residence time distributions can be measured by means of special high-pressure hightemperature conductivity sensors, which are placed directly before and after the reactor [46,48]. NIDAQ Tools for Igor Pro (WaveMetrics Inc., Lake Oswego, OR) is used for both data acquisition and analysis. Data is recorded every 2 s during a run. Each of these measuring points is equal to the average of 20 measurement values obtained over the interval of 2 s. The method for determination of the uncertainty of each measured and calculated quantity is described elsewhere [47]. 5.3. Materials Desalinated water–methanol mixtures were used as artificial fuel and waste water (in this work, we term these mixtures as “fuel” and “waste water”, respectively). Methanol (CH3 OH, purity > 99.8%) was obtained by Syno-pharm Schweizerhalle, Basel, Switzerland. Methanol was chosen as model compound for the following reasons: first, it is easy

to handle (preparation of mixtures, pressurizing and conveying), secondly, in order to demonstrate basic feasibility of the proposed concept using a simple compound, and, thirdly, because it has been widely used as model compound in SCWO studies and the reaction kinetics have been investigated by numerous research groups, e.g. [64–70]. Further, technical grade oxygen (purity > 99.5%) was supplied by Pan-Gas, Dagmersellen, Switzerland. Desalinated water with an electric conductivity of 0.6–0.9 ␮S at ambient conditions was used for the fuel and waste-water mixtures and as transpiring and cooling water. 5.4. Operating conditions All experiments were carried out at an operating pressure of p = 250 bar. The adjusted mass flow rates of all flows entering the reactor are summarized in Table 1. The methanol concentration in the fuel flow (wf ) was 16 and 22 wt.%, while the artificial waste-water flow (www ) contained either 0 (wastewater flow simulated by pure desalinated water) or 6 wt.% methanol. The total oxidation of methanol by oxygen to carbon dioxide and water is given by: CH3 OH + 23 O2 → CO2 + 2H2 O

(1)

B. Wellig et al. / J. of Supercritical Fluids 34 (2005) 35–50

The mass flow rate of the oxygen 1 flow (hydrothermal flame) was calculated with a constant stoichiometric oxygen excess of γ = 1.2, which is defined as follows: γ=

Fox1 Fox1 ≈ ϕO2 Ff wf 1.5Ff wf

(2)

where ϕO2 = (νO2 MO2 )/(νCH3 OH MCH3 OH ) ≈ 1.5. νi and Mi are the stoichiometric coefficient and the molar mass of compound i, respectively. The mass flow rate of the fuel flow Ff was around 1.5 g/s and its reactor inlet temperature Tf was adjusted well above the extinction temperature of the hydrothermal flame (Tf = 200–350 ◦ C). Using Eq. (2), one obtains an oxygen 1 mass flow rate of Fox1 = 0.43 g/s for wf = 16 wt.% and Fox1 = 0.59 g/s for wf = 22 wt.%. The reactor inlet temperature of the oxygen 1 stream Tox1 was set either to 300 or 400 ◦ C. The mass flow rate of the non-pre-heated waste water was Fww = 1.0 g/s (Tww ≈ 20 ◦ C). In all experiments, the oxygen 2 mass flow rate Fox2 was adjusted between 0.2 and 0.25 g/s (which corresponds to a stoichiometric oxygen excess of more than 2 for www = 6 wt.%) as it was not possible to adjust smaller pulsation-free oxygen flow rates. The oxygen 2 flow was also not pre-heated (Tox2 ≈ 30–35 ◦ C, dependent on the running time of the compressor). The transpiration intensity κ is an expedient and suitable criterion for adjusting the mass flow rates of the transpiring water flows. It is defined as the ratio between the mass flux of each transpiring flow and the bulk mass flux at the entrance of the transpiring-wall tube (combustion products + waste water + oxygen 2): κi =

Ftwi /Atwi Ftwi /Atwi = (Ff + Fox1 + Fww + Fox2 )/Ab Fb /Ab

(3)

Ab and Atwi are inner circular area and the inner shell surface of the sinter tube, respectively. According to this definition, the transpiring water mass flow rates for a given intensity value can be calculated as follows: Atwi 4 Ltwi (4) Ftwi = κi Fb = κi Fb Ab di where di is the inner diameter of the sinter tube (22 mm) and Ltwi the active length of the transpiration elements (geometric data see Table 2). Already the preliminary tests showed that it is very difficult to perform manually intensity adjustments of less than 1% due to flow rate fluctuations. For this reason, intensities of 1, 2 and 5% were specified for the experiments. An example of typical operating conditions of the TWR is

41

given in the last column of Table 2. For a transpiration intensity of κi = 2%, the total amount of transpiring water is 2.26 g/s, which is about 65% of the bulk mass flow rate. The reactor inlet temperature of the TW streams (Ttwi ) was varied between 40 and 250 ◦ C. Finally, the cooling water mass flow rate Fcw was between 60 and 70 g/s (Tcw ≈ 28–32 ◦ C, dependent on the running time of the radial piston pump). Until now, the heat recovery cycle has been omitted in the pilot plant. In future experiments, the amount of cooling water can be reduced by heating-up of the inlet flows (fuel, Ox1, TWi) with the hot effluent from the reactor [47]. 5.5. Procedure The first step in each experiment is the ignition of the hydrothermal flame, which can be achieved within about 30 min. The plant is pressurized to 250 bar, fuel and oxygen 1 stream are heated-up, and as soon as the auto-ignition temperature is reached, ignition occurs (the detailed procedure is described in Ref. [47]). Fuel injection temperatures above 460 ◦ C typically led to ignition for 15 wt.% methanol and more. During the ignition process, the waste-water flow consists of desalinated water. After ignition, the fuel and oxygen 1 reactor inlet temperature (Tf , Tox1 ) is set to the desired values and the exact mass flow rates and temperatures of the transpiring water flows are adjusted. Then, the changeover to the well-defined waste-water mixture is performed. As soon as stationary conditions are achieved, the first measurement can be carried out and the liquid effluent is sampled in bottles. For the next measurements, the transpiration intensities and/or temperatures are changed and steady-state conditions are awaited. After the last operating condition of interest, the fuel inlet temperature is lowered to extinction and a changeover to desalinated water is performed (fuel and waste water). Then, the plant has to be cooled for about 15– 30 min before the pressure is reduced stepwise. After the run, the liquid samples are chemically analyzed by GC-FID and overall destruction efficiencies of fuel and waste water are determined. As seen above, the experiments were carried out varying (1) the operating parameters of the hydrothermal flame, (2) the methanol mass fraction in the waste-water stream, and (3) the transpiration intensity (κi ) and the reactor inlet temperatures of the transpiring water flows (Ttwi ). Special emphasis

Table 2 Geometric data of the transpiring-wall elements: inner diameter of sinter tubes di = 22 mm, active transpiration length Ltwi (see Fig. 4), inner shell surface of sinter tubes Atwi = πdi Ltwi , inner circular area Ab = π/4di2 = 380 mm2 (bulk stream) Flow

Length (mm)

Area (mm2 )

Mass flow rate

Ftwi for Fb = 3.5 g/s and κ1 = κ2 = κ3 = 2%

TW1 TW2 TW3

Ltw1 = 77 Ltw2 = 50 Ltw3 = 50

Atw1 = 5320 Atw2 = 3460 Atw3 = 3460

Ftw1 ≈ 14κ1 Fb Ftw2 ≈ 9.1κ2 Fb Ftw3 ≈ 9.1κ3 Fb

Ftw1 = 0.98 g/s Ftw2 = 0.64 g/s Ftw3 = 0.64 g/s

In the last column, the calculation of the transpiring water mass flow rates for typical operating conditions is exemplified (Eq. 4).

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has been put on the influence of the transpiration intensity and temperature on the reactor performance.

6. Results and discussion In a first step, the operating characteristics of the methanol flame in a geometrically simple coaxial burner was examined. The detailed results from these investigations can be found elsewhere [47]. An important finding was that hydrothermal flame combustion can be sustained even with fuel injection temperatures far below the critical temperature of pure water. It can be lowered to subcritical values with at least 11 wt.% and below 100 ◦ C with 27 wt.% methanol. Ignition and extinction curves and the range of stable hydrothermal combustion were determined. For the operating conditions and methanol contents used in this work, the following extinction

XA =

around 50–100 ms in the combustion chamber and subcritical inlet temperatures. The knowledge of the methanol conversion in the flame is important for the TWR operation as it is only possible to determine the overall conversion of methanol in the fuel and waste-water stream. This overall conversion XA is defined by the following equation, where FA, in and FA, out are the mass flow rates of methanol entering and leaving the reactor (methanol is referred to as “A” in the following): XA = 1 −

FA, out FA, in

(5)

The total methanol mass flow entering the reactor is FA, in = Ff wf + Fww www . The mass flow rate FA, out can be calculated with the measured residual methanol concentration in the liquid effluent wout and the total sum of all water flows leaving the reactor:

1 − [Ff (1 − wf + wf ϕH2 O ) + Fww (1 − www + www ϕH2 O ) + Ftw1 + Ftw2 + Ftw3 + Fcw ]wout Ff wf + Fww www

temperatures can be achieved: Tbn = 284 ◦ C for wf = 16 wt.% and Tbn = 224 ◦ C for wf = 22 wt.% methanol (burner nozzle: i.d./o.d. 1.5/7 mm, combustion chamber: i.d./o.d. 9/12 mm, length 50 mm, Ff = 1.5 g/s, γ = 1.2, Tox1 = 400 ◦ C). Typical methanol conversions in the flame based on analysis of the liquid effluent are higher than 99.9% for residence times

(6)

where ϕH2 O = (νH2 O MH2 O )/(νCH3 OH MCH3 OH ) ≈ 1.125. Ff wf ϕH2 O + Fww www ϕH2 O is the amount of water generated by the oxidation reaction, i.e., a complete conversion of methanol to CO2 and H2 O is assumed. The error made by this simplification is negligible compared to the uncertainties of mass flow measurements and of the chemical analysis of the effluent.

Table 3 Operating conditions and results of the experiments discussed in this work No.

Ff (g/s)

Tf (◦ C)

A1 A2 A3 B1 B2 B3 C1 C2 C3 C4 δ

1.57 1.54 1.58 1.53 1.53 1.58 1.56 1.56 1.55 1.54 0.03

251 250 251 200 200 200 350 350 350 350 3

No.

Ftw1 (g/s)

κ1 (%)

A1 A2 A3 B1 B2 B3 C1 C2 C3 C4 δ

0.89 0.88 0.94 0.62 0.99 2.52 0.82 0.86 0.86 0.85 0.03

1.8 1.8 2.0 1.3 2.1 5.2 1.80 1.91 1.95 1.88 −

Fox1 (g/s) 0.62 0.62 0.59 0.60 0.58 0.60 0.45 0.44 0.42 0.45 0.013 Ttw1 (◦ C) 75 150 225 44 40 36 50 124 199 248 3

Tox1 (◦ C)

γ

304 300 300 397 398 398 405 401 401 401 3

1.19 1.23 1.13 1.18 1.16 1.16 1.20 1.17 1.13 1.23 0.04

Ftw2 (g/s)

κ2 (%) 1.9 1.9 2.0 1.1 2.0 4.9 1.79 1.87 1.92 1.93 -

0.60 0.60 0.61 0.33 0.60 1.53 0.53 0.55 0.55 0.57 0.02

Fox2 (g/s)

Fb (g/s)

Tan1 (◦ C)

Tan2 (◦ C)

Tan3 (◦ C)

Tbn (◦ C)

0.23 0.23 0.22 0.24 0.24 0.24 0.19 0.19 0.19 0.19 0.005

3.49 3.44 3.34 3.31 3.36 3.46 3.25 3.21 3.17 3.22 0.05

227 236 239 289 295 302 277 292 297 300 3

301 309 318 294 308 306 265 333 343 344 10

276 293 306 268 265 259 275 310 323 327 3

244 247 246 197 196 196 335 337 336 339 4

Ttw2 (◦ C)

Ftw3 (g/s)

κ3 (%)

Ttw3 (◦ C)

Fcw (g/s)

Tout (◦ C)

XA (%)

76 149 223 35 35 36 51 125 200 248 3

0.62 0.61 0.61 0.32 0.60 1.52 0.54 0.56 0.55 0.55 0.01

1.9 1.9 2.0 1.1 2.0 4.8 1.83 1.90 1.92 1.88 -

76 149 225 51 47 41 49 124 200 250 3

Fww (g/s) 1.08 1.05 0.95 0.95 1.01 1.03 1.05 1.02 1.01 1.03 0.03

73.9 73.4 73.6 64.7 63.9 64.1 73.1 72.3 72.8 72.5 0.9

61 66 70 65 65 66 58 61 62 62 3

− − − − − − − 99.1 99.6 99.4 0.12

Experiment A (Fig. 6): operation with hydrothermal flame only, wf = 22 wt.%, κi ≈ 2%, Ttwi ≈ 75, 150, 225 ◦ C. Experiment B (Fig. 7): operation with hydrothermal flame only, wf = 22 wt.%, Ttwi ≈ 40 ◦ C, κi ≈ 1, 2, 5%. Experiment C (Fig. 8): destruction of an artificial waste water, www = 6 wt.%, hydrothermal flame wf = 16 wt.%, κi ≈ 2%, Ttwi ≈ 125, 200, 250 ◦ C. In the last row of both tables, typical uncertainties δ are given.

B. Wellig et al. / J. of Supercritical Fluids 34 (2005) 35–50

In the following, three experiments are exemplified (the operating conditions and results are summarized in Table 3). In experiments A and B, the pilot plant was operated with the hydrothermal flame only (wf = 22 wt.%) and the wastewater flow was simulated with pure desalinated water (www = 0). In experiment A, the transpiration intensity κi was constant and the inlet temperatures Ttwi of the transpiring water flows TW1, TW2, and TW3 were varied, while in experiment B, the inlet temperatures were kept constant and the intensity was varied. Finally, in experiment C, an artificial waste water containing www = 6 wt.% methanol was destructed (hydrothermal flame: wf = 16 wt.% methanol). The transpiration intensity was kept at a constant value of κi = 2% and the transpiring water inlet temperatures were changed. Results of further experiments can be found in Ref. [47]. With experiments A and B, the temperature field was measured without reactions in the transpiration zone. Rotational symmetry was assumed for the representation of the temperature profiles in Figs. 6–8. According to the profiles of ex-

43

periment A, supercritical temperatures only prevail on measurement plane P2. The transition most likely occurs just above P3. The temperature curves on plane P3 appear implausible at first sight. It must be noted, however, that the uncertainties produced primarily by temperature fluctuations are smaller than on planes P2 and P4. The profile is reversed on the lowest measurement plane, i.e., low and high temperatures prevail, respectively, at the center and close to the transpiring wall tube. Temperatures on the transpiring wall itself are sub-critical, while on plane P2, they are practically independent of the inlet temperature. This does not apply above P2 where combustion products mix with waste water. In the annular clearance, the water is heated to appreciable temperatures, the differences being smaller than at the outlet of the pre-heater. Similar conclusions can be drawn from the profiles measured in experiment B. For transpiration intensities of κ = 1 and 2%, the temperature on measurement plane P2 is supercritical. Again, the transition occurs between P2 and P3. At

Fig. 6. Experiment A: operation with hydrothermal flame only. Radial temperature profiles at working planes P2, P3 and P4 for a constant transpiration intensity of κi = 2% and various transpiring water inlet temperatures (Ttwi = 75, 150, 250 ◦ C). The operating conditions are listed in Table 3.

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Fig. 7. Experiment B: operation with hydrothermal flame only. Radial temperature profiles at working planes P2, P3 and P4 for a constant transpiring water inlet temperature of Ttwi = 40 ◦ C and various transpiration intensities (κi = 1, 2, 5%). The operating conditions are listed in Table 3.

an intensity of 5%, the temperature at P2 already attains the pseudo-critical temperature of pure water at 250 bar. On planes P3 and P4, the temperatures are sub-critical over the whole radius. The change of the profile over the length of the reactor is also clearly evident. The temperature maximum is no longer located at the center, but closer to the transpiring wall. This suggests the formation of an eddy due to natural convection in the middle of the reactor (i.e., below the supercritical region). The water rises in the center of the TW tube, mixes with the hot flow on attaining a certain height and flows back down through the tube periphery. This appears to be the sole plausible explanation for the fact that the temperature at the transpiring wall is more than 100 ◦ C higher than at the center. It is also evident that the temperature at r = 9 mm is higher on plane P4 than on plane P3. One possible explanation for this effect is that the descend-

ing fluid is pushed outward, thus not mixing directly with the water in the core flow. The natural convection eddy may not disturb the desired plug flow behavior in the upper part of the transpiration zone. The intensity change influences the temperature behind the transpiring-wall element of section TW1. The temperature in the annular gap here is up to 200 ◦ C higher than at the pre-heater exit (e.g., section TW1 and κ = 1%: pre-heater outlet temperature Ttw1 ≈ 40 ◦ C, temperature in the annular gap Ts1 ≈ 230 ◦ C). Consequently, a relatively large radial heat flow in the opposite direction of the transpiration flow must be available for heating the fluid (i.e., heat losses). Finally, experiment C was used to investigate the degradation of a 6 wt.% methanol–water mixture. The first temperature field (C1) was measured before changeover to the waste-water flow; the results here confirm the statements

B. Wellig et al. / J. of Supercritical Fluids 34 (2005) 35–50

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Fig. 8. Experiment C: destruction of an artificial waste water. Radial temperature profiles at working planes P2, P3 and P4 for a constant transpiration intensity of κi = 2% and various transpiring water inlet temperatures (Ttwi = 125, 200 and 250 ◦ C). The operating conditions are listed in Table 3.

made above. As soon as a reaction occurs in the transpiration zone of the reactor, the temperature field changes fundamentally: First, the temperatures rise sharply and secondly, the profiles clearly indicate the absence of a natural convection eddy. The core flow expands, pushing down the cold region and stabilizing the entire flow field. On plane P2 and P3 the temperature of the bulk flow is super-critical and the temperatures near the wall remain sub-critical. It seems that the transpiration and bulk flow does not mix strongly and that the desired sub-critical “film” at the inner surface of the transpiring wall is available. The temperatures are not as high as the preceding flame experiments [47] would lead one to expect. However, it must be borne in mind that the distance between the combustion chamber and plane P2 is 72 mm, or 3.3 times the inner diameter of the TW tube. Although the desired hot zone where the destruction reaction takes place is relatively short and the transpiration temperatures are comparatively low (Ttwi = 125–250 ◦ C), small residual methanol concentrations were detected in the samples of the liquid effluent: measurement C2: wout = 38 × 10−6 wt.%, C3: wout = 16 × 10−6 wt.%, C4: wout = 23 × 10−6 wt.%. In all samples, no by-products were found. Using Eq. 6, destruc-

Fig. 9. Comparison of the temperatures near the transpiring wall of the three experiments discussed above. Only temperatures on plane P2 and P3 at radius r = 9 mm (inside the transpiring-wall tube) and r = 16 mm (annular gap between transpiring and load-bearing wall) are considered.

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tion efficiencies of higher than 99% were obtained: C2: XA = 99.1%, C3: XA = 99.6%, C4: XA = 99.4%. The residence time in the transpiration zone is typically in the order of some seconds [46,48]. It is worth mentioning that these measurement results are highly reproducible. Furthermore, it is clear that higher methanol contents in the fuel stream (i.e., higher flame temperatures [47]) as well as higher inlet temperatures of the fuel and transpiring water streams would lead to higher overall conversions. However, the goal here was to demonstrate that even subcritical transpiration temperatures lead to relatively high destruction efficiencies. The comparison of experiment B and C shows that the intensity of transpiration influences the temperature profiles stronger near the wall and in the bulk flow (Figs. 7 and 8). Note again the heating-up of the in-flowing transpiring water by

considerable heat losses through the TW tube. The comparison of the temperatures at the radial positions r = 9 and 16 mm in the measurement planes P2 and P3 are shown in Fig. 9. Near the load-bearing wall and the transpiring wall, the temperatures are sub-critical on all measurement planes. It can be seen from the curves of experiments A and B that the temperatures increase with increasing temperature of the transpiring water Ttwi and decrease with increasing transpiration intensity κ. The influence of the transpiration temperature is stronger than that of the intensity. A surprising fact in experiment B is that the temperatures TP2 (r = 9 mm) and TP3 (r = 9 mm) remain almost constant in spite of a variable transpiration intensity. Experiment C exhibits a completely different behavior compared to A and B. The temperatures near the wall as well as the temperatures within the annular gap are clearly

Fig. 10. Radial profiles of thermo-physical properties for two selected operating conditions of experiment C (Fig. 8, κi = 2%): measurement C1: before changeover to waste water (Ttwi = 50 ◦ C), C4: destruction of waste water containing 6 wt.% methanol (Ttwi = 250 ◦ C). Following properties are shown: density ρ, specific heat capacity at constant pressure cp , dynamic viscosity η, self diffusion coefficient D, and thermal conductivity λ.

B. Wellig et al. / J. of Supercritical Fluids 34 (2005) 35–50

higher and change comparably little with increasing transpiration temperature. The temperature TP3 (r = 9 mm) is even higher than TP2 (r = 9 mm). Therefore, the temperature gradient in axial direction in the transpiration zone (P2–P3) is much smaller than in the experiments discussed above. This seems to lead to a stabilization of the flow field within the transpiration zone. To gain more insight into the transport phenomena, the radial profiles of density ρ, specific heat capacity cp , dynamic viscosity η, self diffusion coefficient D, and thermal conductivity λ were computed for two selected operating conditions of experiment C (Fig. 10; measurements C1 (before changeover to waste water) and C4). All properties are calculated for pure water using the IAPWS-IF97 formulation [71] (except for the self-diffusion coefficient [72]). Furthermore, the values for ambient conditions and for steam at ambient pressure are given for comparison. It is evident that large radial and axial gradients within the transpiration zone exist. For example, the density ρ of the bulk flow on plane P2 is 10 times smaller than near the transpiring wall (measurement C4: ρ = 65 and 670 kg/m3 ). The high values of the heat capacity cp on plane P3 indicate the transition range from the super- to the subcritical region. ρ and cp reach liquid-like values until plane P4. The dynamic viscosity η in the supercritical region is more than 30 times smaller than the liquid viscosity and is in the same order as gas-like values (measurement C4: ηbulk ≈ 30 × 10−6 kg/m s on P2 and P3, ηgas = 12 × 10−6 kg/m s). Accordingly, the self diffusion coefficient D is inversely higher. D in the bulk flow is more than 100 times larger than the value for liquid water (measurement C4: D ≈ 20–55×10−8 m2 /s on P2 and P3, Dliq = 2.0 × 10−9 m2 /s) but still several orders of magnitude below the value of steam at 1 bar (Dgas = 3.4 × 10−5 m2 /s). The single-phase system and the relatively fast diffusional mixing in SCW leads to the rapid oxidation reactions. Finally, the thermal conductivity λ is intermediate between the liquid and gas values at ambient pressure. In conclusion, the flow in the transpiring-wall tube features beside fast oxidation reactions and large variations of the thermo-physical properties. These variations have been taken into account in concurrent reactive CFD simulations using appropriate models for description of thermo-physical properties [44,46,48].

7. Conclusions and outlook In this work, the operating characteristics of a transpiringwall reactor with a hydrothermal flame as internal heat source and transpiring-wall technology are reported. As a preliminary step in the novel process study, this work presents first experimental results under salt-free conditions. Numerous phenomena could be identified and explained with the conducted experiments. The reactor has proven highly reliable and is able to withstand the pressure and temperature conditions which prevail. Its special features include the design of the wall-cooled coax-

47

ial burner and the accessibility of the transpiration zone for temperature measurements. The desalinated water stream, which permanently flows through the transpiring-wall elements, ensures that the load-bearing vessel walls are kept at “low” temperatures, i.e., they are never in contact with supercritical reaction media. Several experiments were performed to investigate the influences of the operating conditions on the reactor performance. Especially, the influence of the temperature and intensity of the transpiration was systematically examined. In the case without reactions in the transpiration zone, the measured temperature profiles indicate that a natural convection eddy is formed below the supercritical region. In contrast to that, the flow field seems to be stabilized as soon as reactions occur in the transpiration zone; the profiles indicate the absence of an eddy. In all experiments, the temperatures near the transpiring and pressure-bearing wall remained subcritical (at least at the accessible measurement planes). This fact is a quantitative measure for the wall protection from salt adherence since salt particles precipitated from the supercritical reaction mixture would presumably be redissolved in the “cold film” near the transpiring wall. The transpiration temperature exhibits the strongest influence on the temperature profiles. Furthermore, an artificial waste water containing methanol as model compound was destructed in the TWR. It was found that even relatively low transpiring water inlet temperatures (125– 250 ◦ C) lead to methanol destruction efficiencies of 99% and more. In conclusion, a TWR featuring a hydrothermal flame meets the requirement for a cold supply of the waste water. The non-preheated waste water and oxygen 2 stream reach reaction temperature by mixing with the hot combustion products, i.e., the flame fulfills its function as internal heat source. The wall contact of the hot reaction mixture is prevented by fluid dynamic means, thus avoiding the two main technical difficulties of SCWO (corrosion and plugging). At typical operating conditions, the ratio between the total transpiring water mass flow (without cooling water) and the bulk mass flow is about 65%. Note that natural convection is not negligible in transpiring-wall reactors for SCWO. The reactor has a good performance, when the operating conditions are carefully selected (i.e., stable burning hydrothermal flame, adjustment of suitable transpiration intensity and temperature). This experimental study has also from another standpoint a great importance. It is commonly accepted that in most CFD calculations experimental data for validation purposes is lacking. Our measuring data form a solid basis for the validation of such models as all geometric data and operating parameters are given. In a parallel project [44,46,48], a basis has been set to the CFD modeling of such a reactor. Reactive CFD simulations were performed including the whole complexity of the geometry, as well as several species with an overall reaction. As boundary conditions, the measured operating conditions from the experiments were used. The temperature field in the reactor and the residence time distribution were used for val-

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idation of the calculations. The comparison of the calculated with the experimentally measured values showed generally a good agreement. Both experiments and CFD calculations yielded valuable information for a better understanding of the phenomena occurring in such a reactor and for the design of such SCWO systems. Finally, future work will include the destruction of salt containing artificial and real waste water and a further optimization of the process parameter.

[13] [14] [15]

[16]

Acknowledgements This work was funded by ETH Zurich and Swiss National Science Foundation (SNF). The authors thank Profs. A. von Rotz and F. Rothenbiihler and their students from the University of Applied Sciences of Central Switzerland for performing an extensive stress analysis by means of finite element method (FEM). Technical support by Dr. W. Dörfler, B. Kramer, R. Plüss and M. Huber is greatly appreciated.

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