Cold model testing of an innovative dual bubbling fluidized bed steam gasifier

Accepted Manuscript Cold model testing of an innovative dual bubbling fluidized bed steam gasifier Andrea Di Carlo, Monica Moroni, Elisa Savuto, Vanessa Pallozzi, Enrico Bocci, Patrizio Di Lillo PII: DOI: Reference:

S1385-8947(18)31549-3 https://doi.org/10.1016/j.cej.2018.08.075 CEJ 19689

To appear in:

Chemical Engineering Journal

Received Date: Revised Date: Accepted Date:

18 June 2018 7 August 2018 12 August 2018

Please cite this article as: A. Di Carlo, M. Moroni, E. Savuto, V. Pallozzi, E. Bocci, P. Di Lillo, Cold model testing of an innovative dual bubbling fluidized bed steam gasifier, Chemical Engineering Journal (2018), doi: https:// doi.org/10.1016/j.cej.2018.08.075

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Cold model testing of an innovative dual bubbling fluidized bed steam gasifier Andrea Di Carlo1*, Monica Moroni2, Elisa Savuto1,Vanessa Pallozzi3, Enrico Bocci4, Patrizio Di Lillo5 1 University of L'Aquila, Via Campo di Pile, L'Aquila, Italy; 2 Sapienza-University of Rome, Via Eudossiana 18, Rome, Italy; 3 Tuscia University, Via S. M. in Gradi 4, Viterbo, Italy; 4 Marconi University, Via Plinio 24, Rome, Italy; 5 Walter Tosto S.p.A., Via Erasmo Piaggio 62, Chieti, Italy *Corresponding author: [email protected]

Abstract

Biomass gasification by a dual fluidized bed reactor is a very promising process to produce a hydrogen rich syngas from biomass wastes. In this process, the bed material circulation must be enough to transport heat from the combustor to the steam gasifier, and at the same time siphons/loop-seals must be properly designed to avoid gas leakage between the two reactor chambers (such as N2 from the combustor to the gasifier). A cold model of an innovative pilot scale dual bubbling fluidized bed gasifier (100 kWth as biomass input) has been realized. The Hybrid Lagrangian Particle Tracking (HLPT) technique and tracer gas analysis, both applied to the cold model, have been used to evaluate the flow rate of bed material circulation and the gas leakage between the steam gasification and the combustion chambers. The results have shown that the bed material circulation is 2-3 times the minimum required to assure allothermal gasification, while gas leakage is negligible for every operating condition evaluated experimentally. Keywords Dual fluidized bed gasifier, Cold Model, Hybrid Lagrangian Particle Tracking Nomenclature

Abbreviations CM

Cold Model

DFB

Dual Fluidized Bed

fps

frames per second

GR

Gasifier Real 1

HLPT

Hybrid Lagrangian Particle Tracking

Symbols Ar

Archimedes number

Asiph

Surface area of the upper siphon, m2

De

Density number

dp

Sauter mean particle diameter, m

Fl

Flow number

L

characteristic length, m

Le

Length number

m&

mass flow rate, kg/h

nsiph

number of upper siphons

Q

volumetric flow rate, m3/h

Qr

volumetric gas flow ratio

tin

initial time, s

tfin

final time, s

U0

characteristic velocity, m/s

ubr

combustor superficial velocity, m/s

ugr

gasifier superficial velocity, m/s

uls

lower siphon superficial velocity, m/s

umf

minimum fluidization velocity, m/s

uus

upper siphon superficial velocity, m/s

uz

mean velocity component in z direction, m/s

Greek letters εmf

void fraction at minimum fluidization

µg

gas viscosity, Pa·s

ρg,p

gas(g) and particle(p) density, kg/m3

2

Τ

characteristic time, s

1. Introduction

Biomass gasification is a thermo-chemical process that allows to convert biomass into a fuel product gas. In particular, steam fluidized bed gasification can produce a syngas rich of hydrogen and carbon monoxide [1–3] suitable for high-efficient power production or for the synthesis of chemicals and biofuels. Dual fluidized bed steam gasification is a very efficient process configuration to produce a hydrogen rich syngas from biomass wastes [4–9], although several issues have to be solved in the design of such reactors. The main drawback of these technologies is to maintain the temperature of the endothermic steam gasifier reactions higher than 1073 K to guarantee a high conversion efficiency. This is ensured by the heat transport from the exothermic combustion reactor to the gasifier. The contact wall between these two reactor chambers has not usually enough surface area to transfer to the gasifier the (very high) amount of heat required there, when the temperature gap with the combustor is limited. Therefore, the heat must be mainly transferred by circulation of granular solid [10]. In practical applications, bed material circulation could be enough to transport heat from the combustor to the steam gasifier, as demonstrated by Hofbauer et al [11] with reference to a dual fluidized bed gasifier composed of two separated reactors (combustor and gasifier), thus with wallout additional heat transfer through a contact. In their system, to obtain a temperature difference below 50 K between the gasification zone (at 1073 K) and the combustor, a high circulation rate of 50 kg bed material per 1 kg dry wood chips is necessary. A different reactor geometry was proposed by Kuramoto et al. [12]: in this case, the two reactor chambers are housed in a single vessel allowing for a contact wall surface area, which could improve the heat transfer rate. In fact the heat could be transferred both by the contact wall surface and by the solid circulation, reducing the intensity of the latter one. Some gas leakage may occur between the two reactor chambers, because of the interconnecting solid loop, for this reason 3

siphons/loop-seals have to be inserted and properly designed. For instance, N2 leak from the combustor to the gasifier would dilute the product syngas, undermining the effort in the realization of the dual fluidized bed (DFB) system. A possible way to evaluate these critical aspects is to develop tests at laboratory scale with reactor models able to faithfully reproduce the hydrodynamic behavior of the real DFB gasifier. To this scope, Glicksman[13] was the first to introduce the concept of dynamic similarity as an important tool in the design of fluidized bed reactors, furnishing reliable rules for the construction and operation of “cold models” of prototype reactors, of reduced size and capacity, which may be easily operated at ambient conditions to predict the fluid dynamic behaviour of real reactors. In particular, this can be done by a set of scaling laws for fluidized beds based on corresponding dimensionless parameters. The usefulness of this approach has been experimentally validated for a large number of fluidized systems [14–17]. Scaling rules are based on the dimensionless equations of change for fluidization, which define a set of characteristic dimensionless groups [18]. Fluidized systems possessing similar values for the corresponding dimensionless quantities exhibit similar dynamic behaviour. With reference to bubbling fluidized beds, in addition to the requirement of geometric similarity, which fixes all length dimensions of the cold model on the basis of the length number Le (and should encompass also particle average size, shape and size distribution), the equality of the corresponding Archimedes numbers Ar, density numbers De, and flow numbers Fl, gives rise to dynamic similarity of the gasifier reactor (GR) and its cold model (CM) [19] The aim of this work is to realize a cold model of the innovative pilot scale dual bubbling fluidized bed gasifier-HBF2.0 [20] (100 kWth as biomass input) according to the scaling rules mentioned above, suitable to evaluate the bed circulation rate between gasifier and combustor and the gas leakage between them. In particular, the solid circulation rate investigation was carried out via a Lagrangian Particle Tracking (LPT) technique, namely Hybrid Lagrangian Particle Tracking (HLPT). The HLPT algorithm is based on the solution of the optical flow equation and selects areas 4

of each image where significant intensity gradients exist. Such areas are associated to tracer particles and are good features to track from frame to frame. Once the particles have been identified, the algorithm calculates the coordinates of the barycenter and reconstructs the trajectory of each particle, calculating their displacement in the subsequent frames. The flow field is described with a Lagrangian approach and the particles velocity is evaluated by calculating the displacement along the reconstructed trajectory and the corresponding elapsed time [21–23]. HLPT allowed the reconstruction of the velocity field of fluidized bed solid particles inside the apparatus. LPT techniques are usually less extensively used with respect to standard Particle Image Velocimetry (PIV) but, compared to PIV, allow a higher spatial resolution (being able to detect regions very close to the walls) and an increased dynamic range [24]. In addition, the gas leakage was evaluated by means of tracer gas injection. In particular, a known concentration of CO2 has been fed with the fluidization air, alternatively in each of the two CM chambers, and then measured in the exit stream from the adjacent chamber by means of on-line gas analyzer, under steady state flow conditions. The gas leakage was finally evaluated by means of mass balances. 2. Design concept

The reactor concept is shown in Figure 1. It consists of two adjacent fluidized beds inside a single vessel: 1. the gasification zone (external cylinder) fluidized by steam and 2. the combustion zone (internal cylinder) fluidized by air. The two chambers are connected with orifices at a given vertical distance to allow bed material circulation. The two fluidized beds operate at different temperature and superficial velocity (u s): (i) the fast bed (combustor) at T~1173 K and us = 5-10 umf and (ii) the slow bed (gasifier) at T ~1073 K and us = 2-3 umf.

5

Figure 1-3D sketch of the new dual fluidized bed gasifier concept: 1) gasifier chamber; 2) lower siphon region; 3) combustor; 4) upper siphons The circulation system considered here is based on the application of the physical concept originally studied experimentally by Kuramoto et al. [12] (see Figure 2).

6

Figure 2-Solid circulation between two interconnected fluidized beds (IFB). It consists of two granular beds, fluidized at different gas velocities and interconnected by means of a baffle plate with one opening at the base and one at the bed surface. The fluidized condition implies that the pressure drop across the bed balances the solid load per unit cross section, and that bed expansion increases with gas velocity, mainly as a result of an increase in bubble hold-up. The presence of the interconnecting lower orifice brings the twin bed system to behave as a corresponding liquid system, in accordance with the principle of communicating vessels. If the overall bed inventory is sufficient to allow for the flow of particles over the top of the upper orifice, a solid circulation is induced and sustained by differences in the gas superficial velocities fluidizing the two beds. This solid circulation gives rise to a pressure difference across the base opening, directly proportional to the difference in the average densities of the two fluidized beds and to their 7

heights. As illustrated in Figure 2, the more dense bed will move downwards (DFB: down-flowing bed), while the other, which contains a greater fraction of bubbles, will move upwards (UFB: upflowing bed). The light biomass particles will then be fed to the more dense bed so that they are dragged inside the bed. As a matter of fact, due to the circulation of bed material established within the device, heat is exchanged between the combustor and the gasifier: (i) sand and residual char in the slow bed (gasifier) flow into the fast bed through the lower orifice and (ii) hot sand is recycled back into the slow bed through the upper orifice. Char combustion in the fast bed supplies the heat to be transported to the gasification chamber. To avoid gas leakage, loop seals just fluidized with steam are included in the reactor design (see Figure 1). The main novelties of this design are summarized below: •

the system is compact and thus suitable for small scale applications, being both reaction chambers (gasification and combustion) integrated in one cylindrical body;

•

the heat exchange between the two chambers occurs through bed material circulation and also by conduction/convection through the wall of the internal cylinder;

•

the higher temperature chamber (combustor), operating at 1173-1223 K, is thermally insulated; this reduces the drawback of thermal losses in small scale applications;

•

longer residence time in the combustor (bubbling bed) allows complete burning of char particles.

The main issues in the design of such reactors are: •

bed material circulation must be enough to guarantee sufficient heat transfer from the combustor to the gasifier,

•

siphons/loop-seals must avoid gas leakage between the reactors (e.g. N2 from the combustor to the gasifier)

8

In an internally circulating fluidized bed reactor, the circulation rate of the bed material is a very important feature. As mentioned in the introduction, Hofbauer et al. [11] demonstrated that to get a temperature difference below 50 K between the gasification zone (at 1073 K) and the combustor a very intense solid circulation rate is necessary: for a 100 kWth gasifier (~20 kgdry,bio/h) the estimated bed circulation rate should be around 1000 kg/h. For this reason siphons must be properly designed to guarantee the required bed circulation avoiding as much as possible gas leakage. 3. Cold model testing

A cold model of the dual bubbling fluidized bed gasifier briefly illustrated above (100 kWth) was realized and operated. The scaling rules proposed by Foscolo et al. [15] were adopted. According to them, the hydrodynamics of the real gasifier can be checked with experimental tests at ambient temperature and pressure by means of a cold model of reduced size and capacity, which may be easily operated to predict the fluid dynamic behaviour of the envisaged prototype reactor. Four dimensionless numbers should assume close values in the cold model and in the reactor: 1. Density number, De =

ρg ρp

2. Archimedes number, Ar =

3. Flow number, Fl =

d p3 ρ p ( ρ p − ρ g ) g

µ2

U0 u mf

4. Length number, Le =

L (geometric similarity) dp

Once the size (L|GR), and operating conditions of the gasifier are known ( ρ g , µ g , ρ p , d p ,U 0

GR

) and

air at ambient temperature and pressure is chosen as fluidizing gas for the cold model ( ρ g , µ g the 4 above expressions allow to define the remaining 4 parameters ρ p , d p ,U 0 , L( size)

CM

CM

),

.

9

Table 1 reports the results of calculations matching the dimensionless numbers for the real gasifier and the cold model. Table 1-Choice of the cold model based on hydrodynamic similarity rules.

Gasifier Reactor (GR)**

Cold Model (CM)

T=1123 K, p=1.11·105 Pa

T=298 K p=1.01·105 Pa

Gas density, ρ g (kg/m3)

0.29

1.2

Gas Viscosity, µ g (Pa·s)

3.6·10-5

1.8·10-5

Particle density, ρ p (kg/m3)

2400 (olivine sand)

9000 (copper sand)

Particle size, d p (m)

500·10-6

~125·10-6

Minimum fluidization velocity, umf (m/s)

0.11

0.044

Linear dimension scale ratio LGR/LCM

4:1

Gas flux ratio UGR/UCM

~2:1

Volumetric flow rate ratio Qr=QGR/QCM

~32:1

Time scale ratio τ GR / τ CM

2:1

**Averaged values between Gasification and Combustion chambers

As shown in Table 1, the particle density in the cold model turns out to be much higher than in the gasifier, while the average diameter is reduced by a factor of four; this reduction ratio applies to all the cold model linear dimensions. Finally, the fluidizing velocities at which the cold model must be operated are around one half of the corresponding ones for the gasifier, this results in a volumetric flow rate reduction of 32 (i.e., 4×4×2). This substantial reduction both in size and flow rate, with 10

the obvious advantage of operating at ambient conditions, makes the cold model investigation feasible at lab scale to evaluate the hydrodynamic behavior of prototype gasifier reactors of capacity up to the order of a MWth. Copper sand was employed to satisfy the Cold Model requirements for particle density and average diameter (Table 2): Table 2- Main characteristics of copper sand chosen as bed material for the CM

Particle density, ρ p (kg/m3)

8822

Particle size, d p (m)

122·10-6

Minimum fluidization void fraction εmf

0.4

Minimum fluidization velocity umf (m/s)

0.044

To allow direct visualization of the bed dynamic behavior under fluidization conditions, the gasifier cold model was realized in Plexiglas. Figure 3 shows a drawing with the dimensions (in m) of the model key elements and the air inputs for the fluidization and recirculation of the bed material (red arrows):

11

Figure 3-Dimensions of the CM key elements 3.1 Lagrangian Particle Tracking analysis to estimate the bed circulation rate

In order to verify the solid circulation rate in the cold model, some tests were carried out using HLPT. A high-speed and high resolution CCD camera (Mikrotron EoSens) equipped with a Nikon 50 mm lens was used for this scope. The camera is able to acquire images with a speed of up to 500 fps and resolution of 1280 x 1024 pixels. A digital video recorder (Io Industries DVR Express Core) was used to store the images. The area investigated is shown in figure 4, the size of the field of view is 0.133 m x 0.166 m (H x L), however the particle tracking analysis was carried out on a smaller portion of 0.034 m x 0.039 m (H x L) corresponding to the location of the upper siphon. This region was lighted with a suitable apparatus. The light scattered by the copper sand particles was captured by the camera. The images were acquired at a rate of 250 fps. 12

HLPT allows to track the path of the solid particles and, knowing the acquisition frame rate, i.e., the time interval between two consecutive frames, to calculate the Lagrangian velocity of each single particle. The advantage of using this image analysis technique lies in its suitability to provide a large number of velocity vectors from which the material flow rate for a specific region of the reactor may be computed [21,22].

13

Figure 4-a) Sketch of the reactor where the field of view (black rectangle) and the image portion processed with the particle tracking algorithm (red rectangle) are highlighted; b) one acquired image The image analysis tests were carried out at the operating conditions described in Table 3. Table 3- operating conditions for cold model tests H static bed (m)

0.15

ubr/umf

7

ugr/umf

2

uls/umf, uus/umf

1.5

14

Figure 5 shows a time sequence where the trajectories reconstructed by the HLPT algorithm are overlapped onto the acquired images. The outgoing sand from the siphon is clearly discernible.

Figure 5-Snapshots showing copper sand “eruption” from the upper siphon With reference to the periodic “eruption” of copper sand from the siphon, it was possible to evaluate for each frame the exit velocity of the particles, selecting only the trajectory portions with uz> 0, i.e., the outgoing siphon velocity component. The frequency of copper sand “eruption” is around 10 times per second. Being the frame rate (250 fps) much higher, this ensures that no sampling distortion problems took place in monitoring particles flowing out of the upper syphon. Figure 6 shows the mean velocity component uz averaged over all tracked particles versus time for roughly 2 seconds of sampling.

15

Figure 6- Mean velocity component u z averaged over all tracked particles versus time Being the copper sand a Geldart Group B particulate solid, and the superficial velocity in the upper siphons close to u mf, it was assumed that the volume fraction of the dense bed (outgoing from the siphons) is that at minimum fluidization (1-εmf) [25]. Furthermore, knowing the dimension (Asiph) of both upper siphons (nsiph= 2), the volumetric flow rate of the sand (QCM) was evaluated as follows (the integral was computed along the entire acquisition time, i.e, 60 seconds)

QCM =

nsiph Asiph (1 − ε mf

)∫

t fin

tin

u z (t ) dt

t fin − tin

Using the volumetric flow rate scale factor (see Table 1) and the density of the sand used in the Gasifier Reactor-GR (olivine sand), the solid circulation mass flow rate in the real gasifier is estimated:

16

m&GR = QCM ⋅ Qr ⋅ ρ P = 2600

kg h

This flowrate is 2 to 3 times higher than the required one (~1000 kg/h), thus it should be enough for the allothermal gasification process. Image analysis turned out to be effective to verify the good fluidization regime and, most important, to check the efficacy of the new gasifier design in the recirculation of the bed material between the two chambers of the reactor.

3.2 Gas leakage using tracer gas

In order to avoid dilution of the producer gas with N2 contained in air or with the exhaust gases in the combustor, it is essential to avoid any gas leakage between the two reactor chambers. To check for any gas interchange between the CM chambers, a known concentration of CO2, approximately 10%, was fed with the fluidization air, alternatively to each chamber, as shown schematically in Figure 7. In experiment 1, CO2 was fed with the fluidization air in the CM combustion chamber and measured at the exit of the CM gasification chamber. Vice-versa in experiment 2, CO2 was fed with the fluidization air in the CM gasifier and measured at the exit of the CM combustor. The ABB URAS 14 on-line gas analyzer was used to continuously detect the percentage of CO2 in the exit gas streams. Successively, by means of mass balances, the amount of gas flowing from combustor to gasifier and vice-versa was evaluated.

17

Figure 7-Test rig and procedure adoptedto evaluate gas leakage between CM fluidization chambers by means of experiments 1 and 2, respectively Analyses were carried out using a superficial velocity of 2 u mf in the gasifier and varying the superficial velocity in the combustor (between 5 and 10 umf) and in the siphons (between 1.5 and 2 umf). Tables 4-5 show the test conditions adopted: Table 4- Gas flow rate and superficial velocity at the inlet of the cold model to evaluate gas leakages from combustor to gasifier (1st case) Nm3/s Air to the upper siphon (uus)

0.3·10-4 (2 umf )

Air to the lower siphon (uls)

1.1·10-4 – 1.4·10-4 (1.5-2 umf )

Air to the gasifier (ugr)

4.6·10-4 (2 umf )

Air/CO2 to the combustor (ubr)

5.6·10-4 – 9.7·10-4 / 0.8·10-4 (~5-10 umf ) 18

Table 5- Gas flow rate and superficial velocity u s at the inlet of the cold model to evaluate gas leakages from gasifier to the combustor (2nd case) Nm3/s

Air to the upper siphon (uus)

0.3·10-4 (2 umf )

Air to the lower siphon (uls)

1.1·10-4 – 1.4·10-4 (1.5-2 umf )

Air/CO2 to the gasifier (ugr)

3.75·10-4 /0.8·10-4 (2 umf )

Air to the combustor (ubr)

5.6·10-4 – 9.7·10-4 (~5-10 umf )

Each test was repeated 3 times, each lasting 30 minutes from the time a constant composition was measured by the gas analyser. Several minutes were waited before starting CO2 injection, to allow reaching steady state fluidization conditions; these were monitored by means of continuous measurement of pressure drop through the outer cylinder fluidized bed (gasifier), which reached a constant value of about 50-60 mbar. Figure 8 shows the average values of CO2 concentration measured in the outlet stream in all experimental tests.

19

Figure 8-Results of gas leakage tests: % CO2 measured in 1) 1st case and 2) 2nd case. 20

As clearly shown in Figure 8, for both conditions and at any superficial velocity in the combustor and in the upper siphon, respectively, the outlet CO2 concentration in the adjacent chamber is negligible, the measured values being close to the sensitivity of the gas analyzer (~ 0.3%). This result confirms that the upper and lower siphons, under the tested conditions, are able to guarantee high recirculation of solid material and very limited gas interchange between the two different reaction zones. Finally, a further test was carried out with the cold model to evaluate how the gas fed to the lower siphon (i.e. steam in the high temperature gasifier) distributes between the 2 reactor chambers. The experimental procedure is the same adopted for the gas leakage tests: a known amount of CO2 was added to the fluidization air fed in the lower siphon and its concentration at the outlet of both chambers was measured. Being (i) negligible the gas leakage between the 2 reactor chambers, as demonstrated above, and (ii) negligible the gas flow rate in the upper siphon (less than 10% compared to that in each chamber-see Tables 4 and 5), straightforward mass balances allowed to evaluate the distribution between gasifier and combustor of gas fed in the lower siphon. Table 6 shows the adopted test conditions: Table 6- Test conditions to evaluate the distribution between gasifier and combustor of the gas fed in the lower siphon

Air to the upper siphon (uus) Aria/CO2 to the lower siphon (uls) Air to the gasifier (ugr) Air to the combustor (ubr)

Nm3/s 0.3·10-4 (2 umf ) 0.6·10-4/0.8·10-4 (2 umf ) 4.6·10-4 (2 umf ) 5.6·10-4 – 9.7·10-4 (~5-10 umf )

Figure 9 shows that 25-30% of the gas flow rate fed in the lower siphon goes to the gasifier (whereas the complement to 100% goes to the combustor).

21

Figure 9- Percentage of gas flow rate fed in the lower siphon that is found in the gasifier The percentage of lower siphon gas flowing into the gasifier decreases as the combustor superficial velocity increases. This result is probably due to a corresponding increase of pressure drop between the gasifier and the combustor. The bed material contained in the gasifier increases at the expense of that in the combustor, with a consequent adjustment in pressure levels in the whole system, which guaranties the continuous recirculation of bed material between both reactor chambers.

4. Conclusions In this work, a cold model of a dual bubbling fluidized bed gasifier with novel design (100 kWth as biomass input) was realized and used to investigate bed material circulation and gas leakage between gasification and combustion chambers. The results obtained by means of Lagrangian Particle Tracking showed that the obtainable bed material circulation is higher by a factor of 2-3 than that assuring effective heat exchange between the two chambers, as required by an allothermal gasification process. The gas leakage between the two reactor chambers is negligible, being close to 22

zero the output concentration in the adjacent chamber of a tracer gas (CO2) injected in either the combustor or the gasifier, respectively. These results demonstrate that the designed siphons can guarantee sufficient bed material circulation avoiding at the same time contamination of syngas with exhaust flue gas. Finally, with a similar procedure (tracer gas injection) it was verified that 70 to 75 % of the gas fed in the lower siphon flows to the combustor. This percentage value increases with an increase of the superficial velocity in the combustion chamber (from 5 to 10 umf). This result was attributed to the corresponding adjustment of the pressure difference between the two chambers (higher in the gasifier and lower in the combustor), i.e. the operating principle of a dual bubbling bed system. The positive results of the experimental tests carried out with the cold model confirm that the innovative dual bubbling fluidized bed gasifier design should be effective for the production of high Calorific Value fuel gas from biomass gasification. Demonstration developments will involve the realization of a full scale stainless steel gasifier with such configuration, to perform biomass gasification test. Acknowledgement

The authors would like to thank the financial support of the Italian Ministry of the Economic Development for the Project HyBioFlex 2.0 - CCSEB00224. Bibliography

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