Heat exchange in a fast firing kiln for glost firing of po · PDF fileHeat exchange in a fast firing kiln for glost ... After the calculation of the flow field the applicability of

Heat exchange in a fast firing kiln for glost firing of porcelain

Authors: Dipl.-Ing. Friedherz H. Becker, Dipl.-Ing. Lars Lorenz, Prof. Dr.-Ing. habil. Dr. h.c. Gerd Walter August 2006

Heat exchange in a fast firing kiln for glost firing of porcelain Authors: Dipl.-Ing. Friedherz H. Becker1

tel.: 0911/ 52 18 234 email: [email protected] Dipl.-Ing. Lars Lorenz2

tel.: 0 37 31/39 20 13 email: [email protected] Prof. Dr.-Ing. habil. Dr. h.c. Gerd Walter2

tel.: 0 37 31/39 34 94 email: [email protected] 1 Riedhammer GmbH, Nürnberg 2 Institute of heat enginnering and thermodynamics, chair of heat treatment systems, TU Bergakademie Freiberg motivation The knowledge of heat exchange inside the kiln is necessary for the exact calculation of the energy household. While heat conduction and radiation is only ruled by the physical properties of the fired goods, the temperatures and geometries of the kiln and the species concentrations, for the convective heat transfer additionally the flow regime inside the tunnel kiln is needed. In case of a tunnel kiln two facts increase complexity, firstly the superposition of two flows, longitudinal and cross flow, and secondly the complex geometry of stacks for the fired goods. Investigations up to now concentrate on simple configurations like those used in brick firing [1,2,3] respectively only one dimensional flow regimes (longitudinal or cross flow) [4,5]. determination of the convective heat transfer coefficients in the flow model The flow model for the determination of convective heat transfer coefficients is a 1:5 cold model of a real existing tunnel kiln for firing flat tableware (see figure 1).



figure 1: flow model for experimental determination of convective heat transfer coefficients This cold model was already used for investigations with pure longitudinal flow. Therefore it was only needed to be reconstructed for the experiments with cross flows. The schematic of the superposed flows is shown in

figure 2.

longitudinal Cross flow



figure 2: schematic of the superposing longitudinal and cross flow Because of the differences between model und real kiln, e.g. the model ratio of 1:5, the dimensional analysis need to be done to transfer the result into reality. In this case Reynolds similarity is applied. The similarity indicators are defined as follows:

Reynolds-number longitudinal flow: ν

TLL

dw ⋅=Re (1)

Reynolds-number cross flow: ν

TQQ

dw ⋅=Re (2)

Nußelt-number: λ

α TdNu

⋅= (3)

with: Lw velocity inside the empty canal Qw velocity of the injection Td diameter plate ν cinematic viscosity α heat transfer coefficient λ thermal conductivity of the gas The heat transfer coefficient α is calculated by the elementary equation of the heat exchange:

( )LTAQ

ϑϑα

−⋅=

•

(4)

with: •

Q heat flux A heat exchange area of the plate Tϑ surface temperature of the plate

Lϑ temperature of the air The heat flux is realised by a heated model plate (indicated in



figure 2 with “1”, “2” and “3”).The heated model plates are placed at different positions inside the stack to determine the heat transfer coefficients. The results of the measurements are represented in 3D-diagrams, as shown in figure 3. In this way the functional connection between the two Reynolds-numbers of the longitudinal und cross flow and the heat transfer (Nußelt-number) can be easily seen. Functions of the type Nu = f(ReL, ReQ) are calculated by regression of the measurements. These functions lead to a good implementation in computational codes, like the following.

050000

100000

150000

200000

Re quer

0250050

007500

10000

12500

Re längs

0

025

2550

5075

75100

100125

125150

150175

175

Nu

Nu

3d_1.xls : (8)Teller2 un, Re quer , Re längs, NuRank 3 Eqn 310 z=a+bx+cy+dx 2+ey^2+fxy+gx 3+hy 3+ixy 2+jx^2y

r^2=0.98145026 DF Adj r^2=0.95495063 FitStdErr=6.9090751 Fstat=47.030322a=24.743548 b=0.00052360574 c=0.017174281 d=-1.0578251e-09 e=-1.1687861e-06

f=-4.3311297e-08 g=1.6325993e-15 h=2.6697418e-11 i=2.0895474e-12 j=1.0225816e-14

figure 3: 3D-view of the function Nu = f(ReL, ReQ) for the middle plate in the lowest level of the stack (position 0) Determination of the convective heat transfer coefficients by computational simulation an comparison measurements - simulation In the recent years computational simulation has become more and more important for calculations of flow fields and the connected heat transfer. The commercial CFD-code FLUENT was used to do the simulation in our case. Because of the complexity of the geometry and the resulting huge number of cells for the simulation, not all three segments of the flow model were simulated at once.



The calculation was done segment by segment, that means firstly the front segment is calculated to achieve a realistic flow field in front of the real measuring segment. Afterwards the “outflow”-profile of the front segment was used as input condition for the real measurement segment. After the calculation of the flow field the applicability of simple formulas from e.g. VDI- Wärmeatlas should be proofed. In this case of tableware plates the comparison with the simple case “even plate” seems to be suitable. The formulas for the average-curve of the heat transfer can be written as follows [6]:

36640 PrRe,Nulam ⋅⋅= (5)

( )1443210370

3210

80

−⋅⋅+⋅⋅

= − PrRe,PrRe,Nu ,

,

turb (6)

22turblam NuNuNu += (7)

The curves shown in figure 4 can be achieved comparing the results of the measured heat transfer coefficients with the results provided by simulation (local Reynolds-numbers) combined with the Nußelt-equations (5)-(7). They show a very acceptable conformity between simulation and measurement.

figure 4: comparison simulation and measurement of the heat transfer coefficient (position 0)



This conformity is not only achievable at the upper level, as it is shown here. It is also valid for other plates inside the stack, other positions or other velocities of the longitudinal flow.



Solution proposals for calculations of the complete kiln With the knowledge found here principally two ways of calculation can be thought of. The first way is the complete simulation of a fast firing tunnel kiln for glost firing. This way is not applicable with available computing power because of the high complexity of the good stacks and therefore the needed fine meshing of those regions. The second way is the calculation of such a tunnel kiln with the help of balance model, whereby the results of the convective heat transfer, achieved here, are used. This should be demonstrated in the following at the example of a kiln for the glost firing of porcelain. Description of fast firing tunnel kiln for glost firing of porcelain Flow model and computer simulation each are based on a fast firing tunnel kiln for the glost firing of porcelain (figure 5), which was built by Riedhammer GmbH in the 1990ties in the German porcelain industry. The technical data of this kiln plant are listed in Table 1:

Table 1: Kiln data

The 68 high speed burners are combined to 10 control groups of 4 to 8 burners each, installed in the side walls in a staggered arrangement, firing above and below the charge into the kiln room (figure 6). The exit speed of combustion gases from the infinitely variable burners can reach up to 150 m/s. That cross flow is superimposed by the longitudinal flow in the kiln room as shown in figure 2. From the thermotechnical point of view, a tunnel kiln comprises two recuperators connected in series, in which on the one hand hot flue gases flow in counter flow direction to the firing goods from the firing zone through the preheating zone to the kiln entrance while transferring their enthalpy on to the firing goods and cooling off themselves. On the other hand, cooling air flows into the cooling zone, which is blown into the kiln at the kiln exit also in counter flow direction to the goods towards the firing zone, cooling the firing goods while heating up itself. In porcelain glost firing, heated up cooling air is extracted by so-called intermediate exhaust shortly after the firing zone (viewing in travel direction), and is supplied to further utilization, for instance for drying or preheated air. This ensures that the reduction is not influenced by cooling air, which is achieved in the burner zone between ca. 1150 °C and maximum temperature by adjusting the gas / combustion air ratio of the burners. In all sections – heating up, firing, and cooling zone – convective heat transmission is playing an important role while radiation however preponderates in the temperature range above

Kiln type TSR 70/140/50-G Firing temperature 1400 °C Kiln length 70 m Useful width 1.4 m Useful height 0.5 m Cycle time 5.5 h Capacity 16 t / 24 h Number of burners 68 Number of control groups 10



700 °C. Radiation in the heated zone is composed of solid state and gas radiation while in the cooling zone, it may basically be reduced to solid state radiation. Objective of the calculation Since energy input by fuel only takes place in the heated preheating and firing zone, only that section is treated in the following, which altogether involves 52 % of the total kiln length. The objective is to determine the kiln room temperature for a specified temperature of the firing goods, and to draw up a complete energy balance and so a statement on the energy supply for each kiln segment of any arbitrary length in the preheating and firing zones. Preconditions for calculation In order to receive detailed results, the lengths of kiln segments should be as short as possible, for example 1 m. With the exception of the preheating zone, the power input of 2 burners is so covered in each case. For calculating the heat transmission coefficient, it is necessary to determine the geometries provided here by the practical case of application. In addition to the kiln room dimensions, also the setting of ware needs to be defined, for which the crank clearance usually amounts to 65 mm. Solid state radiation in the charge requires the calculation of the irradiation coefficient ϕ, which for the case of parallel plain elements can be taken from [7], and a determination of the emission conditions ε of SiC materials at the bottom side as well as of the engobe at the top side respectively of the porcelain plate standing on it. These ε values are temperature-dependent and can be read off from publication [8]. Calculation of gas radiation requires the knowledge of the partial pressures of radiating gases – in the present case only CO2 and H2O. The concentrations result from the combustion calculation under the assumption of a constant gas/air ratio throughout the entire range of performance of each burner. Gas radiation is furthermore dependent on the layer thickness of radiating gases sgl and on the gas temperature that needs to be calculated as part of the solution while the setpoint temperature curve for the firing goods is assumed to be known. In the present case, temperature was measured by thermocouples at the surface of plates in different locations throughout the setting for the sake of a more accurate calculation (figure 7). The transient temperature field of kiln cars was determined from a prior computer simulation [9] and thereof their heat absorption. Finally, for a complete heat balance of each kiln segment, the wall losses need to be considered that were also acquired during measurements made on location. Calculation Calculation commences with the last segment of the firing zone because at that point, no unknown enthalpy flow of gases from the subsequent segment flows in counter flow direction to the goods. As unknown appear in the energy balance the energy input QE and the enthalpy flow of the flue gases QG2 when exiting from that segment. Both parameters interrelate to the heat transfer conditions in the kiln room, which means, the heat transmission coefficients for radiation and for convection need to be determined. The calculation is approached by anticipating a stirrer vessel flow in each segment, and a mean gas temperature ∂Gm being constant over the entire cross section is defined, which is confirmed as target value in the last step of calculation.



For calculating the emission conditions of flue gases εg, knowledge of the equivalent layer thickness sgl is needed, which is determined according to the following correlation:

GGgl OVs /4 •= (8)

With VG = volume of gas resulting from the total gas volume in the segment divided by the number of crank setups, and OG = surface of such a "gas solid" Between the radiating media gas, crank, and porcelain, a radiation exchange interrelation exists in the kiln room, the so-called three-way exchange, for which an effective radiation exchange ratio is calculated according to the following equation:

gesε = ⋅+ SWGS ϕε

⋅⋅+

⋅

WSW

SSWGW

WSGW

AA

εϕε

εε (9)

The indices G, W, and S relate to the media gas G, crank W, and firing goods S respectively. The individual exchange ratios between gas and firing goods εGS, between gas and crank εGW, and between crank and porcelain εWS are derived from series expansion with the respective emission coefficients εG, εW, and εS and the area ratio AS/AW and the irradiation coefficient ϕ described in the VDI Heat Atlas. It is practical to convert the radiation exchange ratio εges into a heat transmission coefficient for radiation αε, as the total heat transmission coefficient αges is composed of the one for radiation and another for convection:

εααα += kges (10 )

After several transpositions for αε and with σ = Stefan Boltzmann constant, the following equation is received:

+

++⋅⋅=

323 1

Gm

Sm

Gm

Sm

Gm

SmGmges T

TTT

TT

Tσεα ε ( 11)

A calculation of the convective heat transmission coefficient kα can be derived as a result of model analysis from the diagram figure 3 and respective transposition. Provided that in the segments heated by burners, the mean gas temperature TGm according to the stirrer vessel assumption is the same as the gas outlet temperature of the segment TG2, and the average temperature of the firing goods TSm is measured, the enthalpy increase of the firing goods must be equivalent to the quantity of heat transmitted, that is:

( )SmGmSgesSS TTAQQ −⋅⋅=− α12&& (12)



The enthalpy increase of firing goods is calculated from the known correlation:

( )112 −−⋅⋅=− SmnSmnpSSSS TTcmQQ &&& ( 13 )

The segment considered is identified by the index n, mS is the mass flow of the firing goods, and cpS is the goods' specific heat capacity. Taking the mass flow balance into consideration, the fuel quantity to be supplied by the burners to that segment is calculated according to the following formula:

( )

⋅⋅+⋅∂⋅−⋅∂⋅⋅⋅+∂⋅+

−+⋅∂⋅++−−−+−+=

B

LLGpG

B

LLpLLBpB

B

U

AbsDGGpGAbsGDSSWWVWB

LccLcHmmmcQQQQQQQQ

m

ρρλρ

ρλρ min22min

12211212

1

&&&&&&&&&&&& (14)

Whereas: teWandverlusQVW =& Wall losses nthalpieOfenwageneQW =& Kiln car enthalpy thalpieBrenngutenQS =& Enthalpy of firing goods enthalpieEindüsungsQD =& Injection enthalpy lpiegsgasenthaVerbrennunQG =& Flue gas enthalpy enthalpieAbsaugungsQAbs =& Exhaustion enthalpy uhrEnergiezufQE =& Energy input itätWärmekapazc p = Heat capacity mMassenstrom =& Mass flow heizwertBrennstoffHu = Caloric value of fuel Temperatur=∂ Temperature LuftfaktorL =λ Air factor bedarfMindestlufL =min Minimum air requirement Dichte=ρ Density The indices identify: 1 = Intake / entrance 2 = Outlet / exit Abs = Exhaustion

B = Fuel D = Injection air G = Flue gas L = Air



Finally, the gas temperature is received from the equation:

SmSges

AbsVWGGWWDEGm T

AQQQQQQQQ

T +⋅

−−−+−++=

α

&&&&&&&&2121 (15)

A comparison with the initially assumed mean gas temperature GmT calls for repetitions of the aforesaid calculation steps as often as it is necessary to achieve tolerable accordance. Thus the flue gas temperature GmT calculated last is that exhaust gas temperature sought after, by means of which the exhaust gas enthalpy AbgQ& of the kiln is calculated. For this reason, the exact and complete predetermination of a heat balance for the preheating and the firing zone of a fuel-heated tunnel kiln is feasible. Discussion of results The equation system in figure 3 allows to determine the Nusselt number for each kiln segment in function of the longitudinal and cross flows and through the correlation

T

k dNu λ

α•

= (16)

to determine the convective heat transmission coefficient αk. Developing of the heat transmission coefficient in the charge in the heated zones of the kiln can be recognized in figure 8. The convective portion with ca.15 W/m²K to 30 W/m²K is significantly below the values of the heat transmission coefficient for radiation. The least value of convection is noticeable at the end of the firing zone as expected. Because of the counter flow principle, the flow rate of combustion gases from the burners is here at a minimum, however it is adding up due to the other burners in the front part of the kiln. The high flow resistance of the charge results in a strong displacement of flow towards ceiling, border areas and bottom, so that only ca. 20 % to 30 % of the volume flow is streaming through the charge, keeping the convective heat transmission coefficient αk very low. The unsteady, fever curve-like course of the convective heat transmission coefficient αk is caused by the varying burner loads respectively exit speeds of their flue gases. The heat transmission coefficient for radiation αε reaches values of more than 75 W/m²K in the high temperature area at 1365 °C. Such expected higher values in comparison to convection can be explained by the potential influence of temperature. The sum of both heat transmission coefficients, which increases from 40 W/m²K at the beginning of heating up to ca. 110 W/m²K in the high temperature zone, is used for the calculation of flue gas temperature and energy input. For the total energy consumption of the entire kiln, an actuarial deviation of 4 % from the measured value results. The energy demand in individual kiln sections shows deviations of up to 20 % from the values measured. This however can be explained by the charging of kiln cars varying in mass. Summary Investigations were undertaken on the convective heat transfer under superimposed longitudinal and cross flow conditions in a fast firing tunnel kiln for glost firing of porcelain. Available was a cold flow model, from which conclusions were made on the conditions in the real kiln by similarity examinations. Result of these examinations is a determination of the



convective heat transmission in function of longitudinal and cross flows in the kiln room in dimensionless representation. For the application of that conjunction, it is suggestive to subdivide the heated zones of the kiln into small sections and to draw up balance models for each of those, in which – viewing in flow direction – the outlet conditions of one section correspond to the input conditions of the next one. The calculation of combustion gas temperatures is made assuming stirrer vessel flow conditions. By knowledge of the convective heat transmission coefficient, the total heat transmission coefficient under inclusion of the radiant heat portion can be calculated, and eventually the kiln room temperature and energy demand of the tunnel kiln plant. A comparison with total energy demand measurements made on a real kiln plant demonstrated satisfactory compliance. literature: [1] Riedel R.: Der Durchströmungsbrand von baukeramischen Produkten (Pass-through

firing of architectural ceramics products). Keramische Zeitschrift 53 (2001) 4, pages 288-295

[2] Weineck, S.; Jeschar, R.; Pötke, W.: Strömungstechnische Untersuchungen zur Temperaturvergleichmäßigung in Tunnelöfen (Fluidic examinations for comparative temperature measurement in tunnel kilns). Gaswärme International 44 (1995) 5, pages 211-220.

[3] Abou-Ziyan, H.Z.: Convective heat transfer from different brick arrangements in tunnel kilns. Applied Thermal Engineering 24 (2004) 2-3, pages 171-191

[4] Balek, S. and Kazmirowski, C.: Modellversuche zur Bestimmung konvektiver Wärmeübergangskoeffizienten (Model tests for the determination of convective heat transmission coefficients). Neue Hütte 36 (1991) 6, pages 229-231

[5] Kraft, T.; Riedel, H.; Raether, F.; Becker, F.: Simulation des Brennprozesses bei der Herstellung von Gebrauchskeramiken (Simulation of the firing process in the production of functional ceramics). Keramische Zeitschrift 54 (2002) 5, pages 374-381

[6] Gnielinski, V.: Wärmeübertragung bei Strömung längs einer ebenen Wand (Heat transfer at flow along a planar wall) . VDI- Wärme- Atlas (2006) Gd1

[7] Vortmeyer,D.:Eintrahlzahlen (irradiation coefficient ) VDI- Wärme Atlas (2006) Kb [8] Bauer, W.; Becker, F.; Moldenhauer, A.: Spectral emissivities of SiC kiln furniture

and porcelain. I’industrie céramique et verrière 985 (2003), pages 50-54 [9] Becker, F. : Berechnungsfortschritte für keramische Brennprozesse (Calculation

progresses for ceramics firing processes). Gaswärme-International 50 (2001) pages 360-365

This paper is based on a research task which was proposed, supported and financed by the research association of industrial kiln building (FOGI) via “Forschungskuratorium Maschinenbau e.V.”.and the “Arbeitsgemeinschaft industrieller Forschungsvereinigungen e.V”. The authors thank for their support.

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Heat exchange in a fast firing kiln for glost firing of po · PDF fileHeat exchange in a fast firing kiln for glost ... After the calculation of the flow field the applicability of