- Email: [email protected]
Spray footprint effect on the induced distortion by the cooling process in the aluminum extrusion process
Applied Thermal Engineering 57 (2013) 14e23
Contents lists available at SciVerse ScienceDirect
Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Spray footprint effect on the induced distortion by the cooling process in the aluminum extrusion process Saeed Bikass a, b, *, Bjørn Andersson a, Artem Pilipenko a, Hans Petter Langtangen b, c a
SINTEF Material and Chemistry, PB 124 Blindern, 0134 Oslo, Norway University of Oslo, Department of Informatics, Oslo, Norway c Center for Biomedical Computing at Simula Research Laboratory, Oslo, Norway b
h i g h l i g h t s Elliptic distribution of the cooling intensity on the surface was modeled. Input data were real and laboratorial data. The model was evaluated by experimental results. Process parameters were like a real process. The distortion mechanism was illustrated and discussed.
a r t i c l e i n f o
a b s t r a c t
Article history: Received 6 May 2012 Accepted 14 March 2013 Available online 28 March 2013
An efficient cooling following an aluminum extrusion process can eliminate the extra solution heat treatment operation prior to the age hardening and produce sound extrudates (if properly controlled) without large distortions and residual stresses. Generally, a single nozzle provides a spatially varying water flux distribution and causes a non-uniform cooling rate over the surface, which is called a “cooling pattern” or a “footprint” in this study. In quench boxes, a set of several nozzles are used in order to spray over a hot extrudate (usually initial temperature is above 500 C). The spray of the adjacent nozzles can overlap and provide an ideally uniform water flux distribution. In this study, distortions of a free plate due to cooling by two types of nozzles with different footprints were compared: the elliptic and the uniform patterns of singular nozzles and multiple overlapped nozzles, respectively. The goal was to find the benefits and the disadvantages of the two footprints by comparing the distortions due to cooling. Some lab experiments were done in order to measure the impact force distribution of the sprays and the heat transfer coefficient (HTC). These experimental results were used to simulate the cooling process. Finally, it was shown that the setting parameters (i.e., ellipse angle, adjacent nozzles center to center distance called “pitch” and ellipse major diameter) of the elliptic footprint had a significant effect on the resulting distortion and should be selected properly. Using a set of nozzles with elliptic footprint overlapping each other can be advantageous because the cooling pattern can be tailored over the section being cooled and the higher cooling rates can be applied. Ó 2013 Elsevier Ltd. All rights reserved.
Keywords: Aluminum extrusion Cooling process Distortion Cooling footprint Nozzle type
1. Introduction Cooling process in high temperature manufacturing processes like extrusion is an important issue. A controlled cooling in-line after the process can be incorporated to improve the mechanical
* Corresponding author. SINTEF Material and Chemistry, PB 124 Blindern, 0134 Oslo, Norway. Tel.: þ47 40310386. E-mail addresses: [email protected], [email protected] (S. Bikass). 1359-4311/$ e see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.applthermaleng.2013.03.036
properties without too much compromising [1]. Forced convective air cooling can be used to provide an appropriate cooling rate if the components are sufficiently thin. However, there are many alloys and components for which air cooling cannot provide a sufficiently high cooling rate [2]. For these situations, water spray could be a proper choice. Spray cooling is a technique in which droplets are generated by a nozzle. The droplets are directed toward a surface, and the heat is removed via the corresponding heating and the subsequent evaporation of the liquid [3]. Spray cooling with phase change has long
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
been recognized [3] as a powerful method to remove high heat flux (>1000 W/cm2) from surfaces [4]. Air-assisted atomization has an advantage over conventional water sprays as the variation of the air pressure and the water pressure can provide well dispersed sprays over a wide range of flow rates [2]. Then it is possible to adjust the quenching according to the necessary cooling conditions. However, problems with the geometrical distortions after quenching and variations in mechanical properties may arise due to rapid and uneven cooling. A distorted extrudate can be straightened, but it incurs extra costs and will induce internal stresses [5] of unknown consequences, whereas the varying mechanical properties can be critical to the product quality. The surface temperature is generally the most important parameter in quenching and is used to define four distinct heat transfer regimes (Fig. 1): (1) film boiling, (2) transition boiling, (3) nucleate boiling, and (4) natural convection (single-phase liquid cooling) [6]. A single hollow cone nozzle provides a bell shape water flux distribution over the surface and causes a non-uniform cooling rate over the surface, which is called a “cooling pattern” or a “footprint”. To get a relatively uniform water spray in quench boxes, the cooling can be performed with sets of nozzles. If neighboring nozzles are close enough to overlap, then it is possible to create more or less uniform water flux over the cooled surface. It is important to have the possibility to control the nozzles separately. This is especially advantageous for the sections with a varying thickness across the width, several newer developments use a high number of small nozzles for quenching [5]. This study aimed to compare the effects of two types of footprints; I) an elliptic footprint caused by using singular nozzles and II) a uniform/rectangular pattern which might be approached by overlapped nozzles. The Bertin nozzle technology was selected as an example with the rectangular footprint. This kind of sprayer has been developed in such a way that a uniform pattern of water is generated by low air pressure and constant airflow rate [1]. It was shown that the corresponding cooling was uniform as well [1]. A set of nozzles that provided elliptic footprints as proposed by Lechler [7] was selected for the study of effect of the elliptic footprint. In this study, we will refer to these two cooling patterns as the “rectangular” and the “elliptic” footprints. The finite element (FE) method was applied to simulate the cooling process. The FE method’s quality has been proven before; for example Järvstråt and Tjøtta made a FE model in ABAQUS to
15
study the thickness effect on the distortion during cooling of a section [8]. Becker et al. [9] carried out a fundamental study on distortions and stresses induced by cooling of an aluminum sample from one side and found an agreement between the residual stress predictions and experiment results. Similar work has been done by Kristoffersen later [10]; he showed that there was a good agreement between the theoretical and the empirical results, too. Kaymak [11] and Pietzsch et al. [12] also studied longitudinal distortions of a L-shape section. They showed that an enhanced cooling at the mass-lumped region combined with a reduced cooling intensity at the edges and thin parts could reduce both distortion and residual stresses. Walle investigated on symmetric non-uniform cooling across the width by 3D FE simulations [13]. He used moving cooling source and illustrated that the temperature gradients in the length direction significantly affected on the extrudate distortions. In this study, a free plate was exposed to cooling from one side using the rectangular and the elliptic footprints. The quality of the prepared FE model had been validated by the experimental results previously [14]. Some lab experiments were implemented to measure the impact force distribution and to inversely calculate the heat transfer coefficient HTC (based on the measured cooling history). The HTC distribution over the considered surface was determined by using the relationship between HTC and the water flow rate [15,16]. The main goal of this study was to evaluate these two types of footprints (rectangular and elliptic) and find the benefits and the disadvantages of them by comparing the simulated distortions caused by the non-uniform cooling. 2. Lab experiments Two sets of experiments had been done by Lechler GmbH to illustrate two relevant specifications of the nozzles: 1. Footprint: A useful experiment is “the impact force” test of the spray [17], which is measured based on a resistance strain gage [18] and [19]. The impact force addresses the water flow rate, the nozzle feed pressure and the impingement area over the cooled surface [20]. The impact force test indirectly characterizes the cooling intensity experienced directly on the surface [21] (higher flow rate and feed pressure had higher impact
Fig. 1. The water spray cooling regimes (“a” and “q” are the heat transfer coefficient and the heat flux, respectively) [3].
16
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
force on the surface [20] and consequently higher cooling intensity [20] and [22]). Therefore this experiment was used to identify the footprint. The impact force value and distribution for the selected nozzle settings are illustrated in Figs. 2 and 3. As it is shown in the figures, the impact force distributed almost evenly for the Bertin nozzle, while the other nozzles gives higher impact force at the nozzle center. 2. HTC: In another set of experiments, the temperature close to the surface of a sample was measured and the cooling histories of the cooled surface for both nozzle types were recorded. Then the HTC of every experiment was calculated by solving the heat transfer equations (by the finite difference method) in the range of 100 Ce500 C (Fig. 4) [23].
the effect of the cooling pattern (uniform and elliptic) on the distortion without interference from the HTC difference. In the second step, both the characteristic HTC and the pattern introduced by each nozzle were considered. Therefore, in this step, the HTC and the cooling pattern of the rectangular pattern were the same as in the first step (low HTC and uniform footprint), whereas the other one had a higher HTC (Fig. 4a) and an elliptic footprint. The goal of this step was to study the combined effect of the cooling pattern and the HTC on the resultant distortion. Some parameters for arrangement of the elliptic type footprints were selected (Fig. 6).
As shown in "fig2 fig3 fig4"Figs. 2e4, both the HTC and the impact force of the elliptic footprint were higher than those of the rectangular one. For the rectangular footprint, the impact force was almost uniform (Fig. 3). Therefore, for the finite element simulation, it was considered to be ideally uniform over the sprayed surface. For the other footprint, the pattern was idealized by considering it as an elliptic pattern with a Gaussian distribution that had a maximum value at the center and gradually decreased to zero toward the boundary of the ellipse. The cooling pattern over the sample surface was shown in Fig. 5. This is more clarified later in this paper.
This was the angle of the ellipse major diameter with the extrusion direction (x-axis). This parameter is responsible for adjustment of the heat transfer distribution over the surface. Its values were considered from 23 up to 80 . 3.2. Ellipse pitch
3. Investigation method and modeling case design
3.3. Major diameter
A 350 200 2 mm3 (length width thickness) free plate was to be quenched from the top surface by the moving cooling source. The initial temperature of the sample and the sprayed water were 540 C and 17 C, respectively. The cooling source was moving along the length of the sample with a 0.25 mm/s velocity. The cooling source was introduced by the HTC as a function of the surface temperature of the selected section (T) and positions on the surface (x,y):
The size of the transverse (larger) diameter of the elliptic pattern was named by “2a”. In this paper, only this diameter (2a) was changing between 200 e350 mm while the other one (2b) was considered constant. As 2a increased, the ellipse was elongated (the ellipse eccentricity approached 1).
HTC ¼ f ðT; x; yÞ The cooling pattern was either uniform or elliptic. Both HTC curves were simplified to decrease the calculation time (see the green line and the blue line in Fig. 4 in web version). As mentioned before, the HTC curves and the footprints of the nozzle examples were significantly different. The changing shape response could be caused by either the differences in HTCs or the footprints. To separate these factors, the study was done in two steps. In the first step, one of the simplified HTCs (simplified Bertin in Fig. 4b) together with different footprints were considered to study
3.1. Ellipse angle
The center-to-center distance of the adjacent ellipses (or the nozzles) across the extrudate width direction (y-axis) was called the “pitch” (Fig. 6). This parameter was studied in the range of 84 e 130 mm.
3.4. Minor diameter The conjugate diameter 2b was 52 mm for all studied cases. A “reference case” with suggested values for the selected parameters by the manufacturer was also defined (Table 1). To study the first and the third parameters (Angle and 2a), the pitch was chosen as small as possible without overlapping of two adjacent ellipses. 4. Material and process parameters The sample material was aluminum alloy 6082. The physical properties of the alloy are given in Table 2. The mechanical properties (including strength, elastic modulus and Poisson ratio) were
Fig. 2. The impact force distribution over the surface for the selected Lechler nozzle (water pressure ¼ 4 bar, air pressure ¼ 4 bar, distance between nozzle and cooled surface ¼ 200 mm).
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
17
Fig. 3. The impact force distribution over the surface for the Bertin nozzle (water pressure ¼ 2 bar, air pressure ¼ 0.1 bar, distance between nozzle and cooled surface ¼ 200 mm).
Fig. 4. Real (an example) and simplified (the implemented and claimed by the manufacturer) HTC curves of a) Lechler and b) Bertin nozzles.
5. FE model The program ABAQUS/Standard was used to simulate the cooling process and the distortions in 3D. The analysis type was “coupled temperature-displacement”, with the brick element, called C3D20T in the ABAQUS element library. This is a second order hexahedral solid element with 20 nodes. Four elements were used throughout the thickness (Fig. 8) [13,14]. The elasticeplastic material model was used in order to introduce the selected material behavior under thermal loads. The data achieved from laboratorial tests was entered as a table in the software. The software solves the following equation in order to calculate temperature distribution in the sample [24]. Fig. 5. Elliptic cooling pattern over the considered sample.
defined to be temperature and strain rate dependent. The temperature dependency of the elastic modulus and the strengths (yield and ultimate strengths) are shown in Fig. 7. The extrusion process parameters were selected close to the real process in the plants (Table 3).
VðkVTÞ ¼
rcp vT vt
Table 1 Details of the reference case and the selected parameters. Parameter
Range
Reference case
Angle ( ) Pitch (mm) 2a (mm) 2b (mm)
23e80 84e130 200e350 52
50 84 270 52
Table 2 Physical properties of the selected alloy.
Fig. 6. The selected parameters of the elliptic footprint (dimensions are in mm).
Density [kg/m3]
Thermal expansion coefficient [1/ C]
Thermal conductivity [W/m K]
Specific heat capacity [J/(kg C)]
2700
2.55e-05
200
900
18
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
80 70
250
Stress (MPa)
60 200
50
150
40 30
100
20
50
Elastic Modulus (GPa)
300
YS, SR= 0 YS, SR= 1 UTS, SR= 0 UTS, SR= 1 EM
10 YS=yield stress (plastic strain=0)
0
0
0
200 400 Temperature (°C)
600
UTS=ultimate stress (plastic strain=31%) SR= strain rate EM= elastic modulus
Fig. 7. Stress and elastic modulus curve versus temperature for different plastic strains.
Table 3 Extrusion process parameters. Run-out speed [m/s]
Initial profile temperature [ C]
Water temperature [ C]
Ambient temperature [ C]
0.25
540
17
17
where r is the density, cp is the specific heat capacity, k is the thermal conductivity. T and t are temperature and time, respectively. For coupled temperature-displacement systems ABAQUS solves a system shown as following [24]:
Kuu KTu
KuT KTT
Du DT
¼
Ru RT
where Du and DT are corrections to the incremental displacement and temperature, respectively. Kij are the sub-matrices of the coupled stiffness matrices and Ru and RT are the thermal residual vectors, respectively. These systems are solved simultaneously using the full Newton’s method. The cooling source was moving along the x-axis from the puller side (right hand side in Fig. 8) toward the die side. As previously mentioned, the studied case was a free plate, and the die and the puller sides were only about cooling source moving direction; there was no mechanical boundary condition representing the die and the puller. The FILM subroutine was used to define the moving cooling source with either the uniform or the elliptic footprint. This
subroutine was used to provide heat transfer coefficients and sink temperatures for fully coupled thermal-stress analysis. An HTC (called “h” in this subroutine) could be conditionally defined versus temperature and coordinates. Therefore an elliptic footprint was introduced by ellipse function and conditional statements (if... then) over the surface coordinates (x and y). Other conditional statements were used to define the HTC dependency on the surface temperature. The elliptic footprint was addressed by means of several scaledconcentric ellipses representing the cooling intensity boundaries. The resultant HTC and the heat flux distribution are shown in Fig. 9. In the studied cases, several similar distributions moved from the one end to the other end along the x-axis (similar to Fig. 5). 5.1. Boundary conditions The plate was fixed at the width middle points of the two edges in the y- and z-directions (Fig. 8) to represent a free plate behavior. Simultaneously, the edge at the puller side was constrained in the x-direction to make the model more sensitive to the distortion. Generally, thermal boundary condition is defined as below:
kvT ¼ hðT T0 Þ vn where n is the normal direction of the sprayed surface, h is the HTC between the surface and the cooling water, T0 is the water temperature and T is the surface temperature of the sample.
Fig. 8. The prepared meshed model for FE simulation (“Ui” is displacement in “i” direction).
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
19
Fig. 9. a) The HTC distribution b) the heat flux distribution on the surface evaluated in the FE model.
A uniform initial temperature was set at 540 C throughout the sample. 6. Results and discussion 6.1. First step e the same (and lower) HTC The goal of this step was to investigate on the effect of the footprints on the resulted distortions due to cooling. Therefore, the samples were cooled by both the rectangular and the elliptic types. As previously mentioned, the HTC was the simplified Bertin’s HTC (Fig. 4b). An example of the temperature distribution over the cooled surface was shown in Fig. 10. The main interest was the induced distortions (the flat sample’s out of flatness due to cooling) across the width. Because it was difficult to compare the total 3D distortions, we decided to focus on the distortion at the crosssection at the puller side (right side in Fig. 10). To avoid the
influence of the constraint on the right-hand side (Fig. 8), the crosssection was selected 8 mm away from the edge of the puller side (Fig. 10). This position was representative for the sample distortion and was used to compare the distortions at different cooling conditions. The corner points of the cross-section were also used to monitor the cooling history. Cooling with elliptic footprint and lower HTC (like Bertin nozzle’s) gave significantly smaller distortions compared to the uniform spray distribution (as it was shown in Figs. 11 and 12). This result can be explained by reviewing the cooling history (Fig. 13). Cooling by the rectangular pattern (uniform footprint) was more severe and caused higher temperature gradients both through the thickness and along the length. While the elliptic footprint had a smaller average cooling intensity over the surface than the other one. For the elliptic footprint, the cooling was severe only at the central zone of cooling pattern and made the total cooling intensity (over the surface) gentler. However, this was not always positive because poor mechanical properties can be obtained for the heattreatable alloys with thick wall. 6.1.1. Study of the effects of the parameters The parameters of the elliptic pattern (Table 1 and Fig. 6) were also studied, and the results were as follows. 6.1.1.1. Angle. This parameter affected the distortion significantly (Fig. 14a). If the angle was not chosen properly (e.g., 80 ), the final distortion caused by the elliptic footprint was much worse than that by the rectangular one. Between 23 and 50 , better results were obtained. As shown by the red curve (in web version) in Fig. 14b, the 50 angle gave the least distortion. When the angle was
Fig. 10. The temperature distribution on the surface cooled by a) a rectangular footprint at time ¼ 4.3 s and b) an elliptic footprint at time ¼ 8.8 s (maximum distortion time).
Fig. 11. Comparison of the corner displacement history during the cooling process.
20
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
Fig. 15. The temperature distribution when the angle is 80 .
Fig. 12. Comparison of the residual distortion of sections cooled by the elliptic and the rectangular cooling patterns.
Fig. 16. The effect of the pitch on the distortion.
Fig. 13. Comparison of the cooling history at the puller side corners (Fig. 10).
80 , the cooling was concentrated in the middle of the plate (Fig. 15) and resulted in large distortions due to the severe non-uniformity across the width (y-axis). 6.1.1.2. Pitch. This parameter had a similar effect to the angle and should be set as small as possible to have minimum distortion (Fig. 16). Larger pitches introduced more non-uniformity across the width. When the pitch was large, the cooling was concentrated closer to the outer edges of the plate. 6.1.1.3. 2a. This parameter had an insignificant effect compared to the other parameters (Fig. 17). However, when the ellipses were elongated (2a increased), the footprint was more homogenous, many ellipses contributed in the cooling process across the width
(a)
Fig. 17. The major diameter effect on the final distortion.
(b)
Fig. 14. The effect of the angle on the final distortion.
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
21
and the cooling uniformity across the width improved. Its sensitivity was not that strong within the selected limits. 6.2. Second step e realistic HTCs for each footprint
Fig. 18. The cooling history for the selected footprints.
As it was mentioned in the section “Lab experiments”, the selected example of the nozzles for the elliptic footprint provided higher HTC than the other footprint’s example. In this part, the higher HTC (Fig. 4a) was applied to the elliptic footprint while the HTC of the rectangular footprint was same as the previous step. Therefore, in this step the combined effect of the HTCs and the footprints were considered. Comparing the cooling history of the elliptic footprint with the rectangular footprint’s (with lower HTC as it was in the first step), one can see that the average HTC of the elliptic footprint was larger than that of the other one (Fig. 18), resulting in larger distortions as well (Fig. 19). This result is caused by the higher temperature gradients through the thickness and the higher non-uniformity of the cooling across the width (Fig. 20). Consequently, approaching to flexible uniform cooling with proper HTC is ideal as it can minimize the distortions due to cooling with providing required mechanical properties. In case of an elliptic footprint usage, the arrangement of nozzles needs to be optimized for geometrically accurate and cooling-ratesensitive sections. 7. Discussion
Fig. 19. The distortion across the width due to cooling for the rectangular footprint (HTCmax ¼ 10,000 W/(m2 K)) and the elliptic footprint (HTCmax ¼ 30,000 W/(m2K)).
The distortions induced to the section were introduced by local contractions on the plate surface and throughout the thickness. These distortions were caused by uneven cooling. When a set of nozzles with elliptic footprints was applied to cool down the sample, the temperature distribution (or the cooling intensity) was complex and resultant of the effect of every individual footprint
Fig. 20. The history of the maximum temperature difference across the width for the elliptic footprint a) high HTC b) low HTC.
Fig. 21. Temperature distribution over the plate surface cooled with the set of the elliptic footprint at a) time ¼ 1 s b) Time ¼ 2 s.
22
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
Fig. 22. Effect of pitch on the resultant HTC fraction.
(Fig. 21). As shown in the figure, the cooling concentration started at three zones (indicated in Fig. 21a by arrows) across the width, forming by three elliptic footprint central areas. The areas outside of these zones were cooled gently until the following footprints covered them (Fig. 21b). The simulation results showed that the elliptic cooling pattern with higher HTC (30,000 W/(m2 K) in this study) resulted in larger distortions and was not a proper choice for the shape sensitive sections. The distortions were significantly reduced when the lower HTC (maximum 10,000 W/(m2 K)) was applied. This solution is not suitable for the thick walled sections or alloys demanding faster cooling. As it was demonstrated in previous section, to decrease the distortion induced by cooling, the minimum value of the pitch (when there is no overlapping) was desired. According to the nozzles manufacturers’ experience, and the article by Hall and Mudawar [21], the water-flux in the overlapped area is the sum of the water flux of the overlapped nozzles. Mudawar and Valentine in 1989 [25] showed that the water flux has a dominant effect on the heat transfer. These postulates allow us to suggest that the resultant HTC of the adjacent nozzles can be as shown in Fig. 22. In this case, the distribution becomes more uniform across the width. This result can be optimized and a uniform distributed and a higher HTC can be reached (the right hand side in Fig. 22). With this technique, we can also improve the flexibility (e.g., the cooling controllability across the section width) of the elliptic nozzles and the productivity (higher HTC will allow to cool the sections faster). An essential advantage can be gained if each nozzle is controlled individually. In this case, the cooling can be tailored to complex sections such as those with non-uniform thickness (Fig. 23).
8. Conclusion We simulated the cooling processes by the elliptic and the rectangular footprints and compared the induced distortions due to such cooling processes. We also studied the effect of the parameters in the elliptic case. The simulation results showed that the final distortions introduced by the elliptic footprint were less than those introduced by the rectangular type, assuming that both have the same lower level of HTC (maximum 10,000 W/(m2 K)). However, according to the experimental measurements, the elliptic type manufactured by Lechler had a higher level of HTC (30,000 W/ (m2 K)) than the rectangular type. This difference was valuable from the metallurgical point of view for thick sections and special alloys. However, the elliptic distribution of the water-flux with higher HTC was not advantageous for the geometrically sensitive sections because it introduced larger distortions to the studied section due to the higher cooling non-uniformities across the width. This non-uniformity can be eliminated by using the elliptic nozzles with an overlapping area and creating a uniform distribution of water-flux across the width with higher HTC. The parametric study of the elliptic type footprint showed the followings: The ellipse angle had an optimum value of 50 in this paper. If higher angles were set, the cooling was concentrated at the middle across the width. Therefore, high temperature gradients across the width led to larger distortions. The pitch was the second studied parameter. As it was decreasing (the nozzles were getting closer), the final distortion was also decreasing. The transverse diameter, which was called “2a” in this paper, had minor effect. However, larger values (elongated ellipse) were in favor for establishing homogeneous distribution of cooling across the width and for decreasing the distortion. Based on the facts presented here we can also draw the following general conclusions: Gentle (softer) cooling caused smaller distortions. Cooling footprints should not leave any area of the profile surface unsprayed. That was the reason for extreme distortions found for the cases with 90 angle and too high pitch. Acknowledgements
Fig. 23. Controlling cooling intensity across the width.
This work was a part of the project run by Hydro in cooperation with SINTEF Materials and Chemistry. The work is partly funded by the Research Council of Norway (RCN). The authors would like to greatly acknowledge their support. Part of the experimental data
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
was supplied acknowledged.
by
Lechler
GmbH,
which
hereby
is
also
References [1] J. Rivallin, S. Viannay, General principles of controlled water cooling for metallurgical on-line hot rolling processes, forced flow and sprayed surfaces with film boiling regime and rewetting phenomena, International Journal of Thermal Sciences 40 (2001) 263e272. [2] M. De Oliveira, J. Ward, D.R. Garwood, R.A. Wallis, Quenching of aerospace forgings from high temperatures using air-assisted atomized water sprays, Journal of Materials Engineering and Performance 11 (2002) 80e85. [3] T.A. Shedd, Next generation spray cooling, high heat flux management in compact spaces, Heat Transfer Engineering 28 (2007) 87e92. [4] S.S. Hsieh, T.C. Fan, H.H. Tsai, Spray cooling characteristics of water and R134a. Part I: nucleate boiling, International Journal of Heat and Mass Transfer 47 (2004) 5703e5712. [5] M. Plata, J. Piwnik, A. Patejuk, Modern precise high-power water-cooling systems for press quenching, Archives of Foundry Engineering 9 (2009) 205e210. [6] J.D. Bernardin, C.J. Stebbins, I. Mudawar, Mapping of impact and heat transfer regimes of water drops impinging on a polished surface, International Journal of Heat and Mass Transfer 40 (1997) 247e267. [7] Full Cone Nozzles, General Industry, Products, Lechler GmbH, www.lechler. com, Metzingen, Germany. [8] N. Järvstråt, S. Tjøtta, A process model for on-line quenching of aluminum extrusions, Metallurgical and Materials Transactions B 27 (1996) 501e508. [9] R. Becker, M.E. Karabin, J.C. Liu, Distortion and residual stress in quenched aluminum bars, Journal of Applied Mechanics 63 (1996) 699. [10] M. Kristoffersen, Studies of Shape Variations During Cooling of Flat Extruded Sections. PhD thesis, University of Oslo, Norway, 2004. [11] Y. Kaymak, Simulation of Metal Quenching Processes for the Minimization of Distortion and Stresses, Otto-von-Guericke-Universität Magdeburg, Germany, 2007. [12] R. Pietzsch, M. Brzoza, Y. Kaymak, E. Specht, A. Bertram, Simulation of the distortion of long steel profiles during cooling, Journal of Applied Mechanics 74 (2007) 427.
23
[13] T. Walle, Shape Variations Due to Cooling. PhD thesis, University of Oslo, Norway, 2007. [14] S. Bikass, B. Andersson, A. Pilipenko, H.P. Langtangen, Simulation of the distortion mechanisms due to non-uniform cooling in the aluminum extrusion process, International Journal of Thermal Sciences 52 (2012) 50e58. [15] R.H. Pereira, S.L. Braga, J.A.R. Parise, Single phase cooling of large surfaces with square arrays of impinging water sprays, Applied Thermal Engineering 36 (2012) 161e170. [16] W.J.J. Vorster, S.A. Schwindt, J. Schupp, A.M. Korsunsky, Analysis of the spray field development on a vertical surface during water spray-quenching using a flat spray nozzle, Applied Thermal Engineering 29 (2009) 1406e 1416. [17] Nozzle Performance and Service Data, Metallurgical Industry, Nozzle Technology, Lechler GmbH, www.lechler.com, Metzingen, Germany. [18] S. Schürmann, Measurement and mathematical approximation of the impact of descaling nozzles, in: Hydraulic Descaling Conference, 2000, London, UK. [19] E.U. Becker, G. Birkemeier, W. Buchele, M. Degner, L. Devrient, M. Nowak, G. Thiemann, Optimization of descaling nozzles in a hot strip mill, Metallurgical Plant and Technology International 23 (2000) 92e94. [20] L. Bendig, M. Raudenský, J. Horský, Descaling with High Pressure Nozzles, ILASS-Europe, Zurich, Switzerland, 2001. [21] D.D. Hall, I. Mudawar, Experimental and numerical study of quenching complex-shaped metallic alloys with multiple, overlapping sprays, International Journal of Heat and Mass Transfer 38 (1995) 1201e1216. [22] D. Tucker, A Quantitative Study Into the Comparative Efficiency of Flat Jet and Full Cone Spray Nozzles When Applied to the Application of Work Roll Cooling, Service, Lechler GmbH, www.lechler.com, Metzingen, Germany. [23] S. Bikass, B. Andersson, A. Pilipenko, Uncertainties on HTC measurement of water spray quenching of aluminum alloys, in: ASME/JSME 2011 8th Thermal Engineering Joint Conference, AJTEC2011, 2011, T10155-T10155-5, Honolulu, Hawaii, USA. [24] ABAQUS Standard Version 6.5, User Manual, Theory Manual, Hibbitt, Karlsson & Sorensen Inc, USA, 2005. [25] I. Mudawar, W. Valentine, Determination of the local quench curve for spraycooled metallic surfaces, Journal of Heat Treating 7 (1989) 107e121.
Contents lists available at SciVerse ScienceDirect
Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Spray footprint effect on the induced distortion by the cooling process in the aluminum extrusion process Saeed Bikass a, b, *, Bjørn Andersson a, Artem Pilipenko a, Hans Petter Langtangen b, c a
SINTEF Material and Chemistry, PB 124 Blindern, 0134 Oslo, Norway University of Oslo, Department of Informatics, Oslo, Norway c Center for Biomedical Computing at Simula Research Laboratory, Oslo, Norway b
h i g h l i g h t s Elliptic distribution of the cooling intensity on the surface was modeled. Input data were real and laboratorial data. The model was evaluated by experimental results. Process parameters were like a real process. The distortion mechanism was illustrated and discussed.
a r t i c l e i n f o
a b s t r a c t
Article history: Received 6 May 2012 Accepted 14 March 2013 Available online 28 March 2013
An efficient cooling following an aluminum extrusion process can eliminate the extra solution heat treatment operation prior to the age hardening and produce sound extrudates (if properly controlled) without large distortions and residual stresses. Generally, a single nozzle provides a spatially varying water flux distribution and causes a non-uniform cooling rate over the surface, which is called a “cooling pattern” or a “footprint” in this study. In quench boxes, a set of several nozzles are used in order to spray over a hot extrudate (usually initial temperature is above 500 C). The spray of the adjacent nozzles can overlap and provide an ideally uniform water flux distribution. In this study, distortions of a free plate due to cooling by two types of nozzles with different footprints were compared: the elliptic and the uniform patterns of singular nozzles and multiple overlapped nozzles, respectively. The goal was to find the benefits and the disadvantages of the two footprints by comparing the distortions due to cooling. Some lab experiments were done in order to measure the impact force distribution of the sprays and the heat transfer coefficient (HTC). These experimental results were used to simulate the cooling process. Finally, it was shown that the setting parameters (i.e., ellipse angle, adjacent nozzles center to center distance called “pitch” and ellipse major diameter) of the elliptic footprint had a significant effect on the resulting distortion and should be selected properly. Using a set of nozzles with elliptic footprint overlapping each other can be advantageous because the cooling pattern can be tailored over the section being cooled and the higher cooling rates can be applied. Ó 2013 Elsevier Ltd. All rights reserved.
Keywords: Aluminum extrusion Cooling process Distortion Cooling footprint Nozzle type
1. Introduction Cooling process in high temperature manufacturing processes like extrusion is an important issue. A controlled cooling in-line after the process can be incorporated to improve the mechanical
* Corresponding author. SINTEF Material and Chemistry, PB 124 Blindern, 0134 Oslo, Norway. Tel.: þ47 40310386. E-mail addresses: [email protected], [email protected] (S. Bikass). 1359-4311/$ e see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.applthermaleng.2013.03.036
properties without too much compromising [1]. Forced convective air cooling can be used to provide an appropriate cooling rate if the components are sufficiently thin. However, there are many alloys and components for which air cooling cannot provide a sufficiently high cooling rate [2]. For these situations, water spray could be a proper choice. Spray cooling is a technique in which droplets are generated by a nozzle. The droplets are directed toward a surface, and the heat is removed via the corresponding heating and the subsequent evaporation of the liquid [3]. Spray cooling with phase change has long
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
been recognized [3] as a powerful method to remove high heat flux (>1000 W/cm2) from surfaces [4]. Air-assisted atomization has an advantage over conventional water sprays as the variation of the air pressure and the water pressure can provide well dispersed sprays over a wide range of flow rates [2]. Then it is possible to adjust the quenching according to the necessary cooling conditions. However, problems with the geometrical distortions after quenching and variations in mechanical properties may arise due to rapid and uneven cooling. A distorted extrudate can be straightened, but it incurs extra costs and will induce internal stresses [5] of unknown consequences, whereas the varying mechanical properties can be critical to the product quality. The surface temperature is generally the most important parameter in quenching and is used to define four distinct heat transfer regimes (Fig. 1): (1) film boiling, (2) transition boiling, (3) nucleate boiling, and (4) natural convection (single-phase liquid cooling) [6]. A single hollow cone nozzle provides a bell shape water flux distribution over the surface and causes a non-uniform cooling rate over the surface, which is called a “cooling pattern” or a “footprint”. To get a relatively uniform water spray in quench boxes, the cooling can be performed with sets of nozzles. If neighboring nozzles are close enough to overlap, then it is possible to create more or less uniform water flux over the cooled surface. It is important to have the possibility to control the nozzles separately. This is especially advantageous for the sections with a varying thickness across the width, several newer developments use a high number of small nozzles for quenching [5]. This study aimed to compare the effects of two types of footprints; I) an elliptic footprint caused by using singular nozzles and II) a uniform/rectangular pattern which might be approached by overlapped nozzles. The Bertin nozzle technology was selected as an example with the rectangular footprint. This kind of sprayer has been developed in such a way that a uniform pattern of water is generated by low air pressure and constant airflow rate [1]. It was shown that the corresponding cooling was uniform as well [1]. A set of nozzles that provided elliptic footprints as proposed by Lechler [7] was selected for the study of effect of the elliptic footprint. In this study, we will refer to these two cooling patterns as the “rectangular” and the “elliptic” footprints. The finite element (FE) method was applied to simulate the cooling process. The FE method’s quality has been proven before; for example Järvstråt and Tjøtta made a FE model in ABAQUS to
15
study the thickness effect on the distortion during cooling of a section [8]. Becker et al. [9] carried out a fundamental study on distortions and stresses induced by cooling of an aluminum sample from one side and found an agreement between the residual stress predictions and experiment results. Similar work has been done by Kristoffersen later [10]; he showed that there was a good agreement between the theoretical and the empirical results, too. Kaymak [11] and Pietzsch et al. [12] also studied longitudinal distortions of a L-shape section. They showed that an enhanced cooling at the mass-lumped region combined with a reduced cooling intensity at the edges and thin parts could reduce both distortion and residual stresses. Walle investigated on symmetric non-uniform cooling across the width by 3D FE simulations [13]. He used moving cooling source and illustrated that the temperature gradients in the length direction significantly affected on the extrudate distortions. In this study, a free plate was exposed to cooling from one side using the rectangular and the elliptic footprints. The quality of the prepared FE model had been validated by the experimental results previously [14]. Some lab experiments were implemented to measure the impact force distribution and to inversely calculate the heat transfer coefficient HTC (based on the measured cooling history). The HTC distribution over the considered surface was determined by using the relationship between HTC and the water flow rate [15,16]. The main goal of this study was to evaluate these two types of footprints (rectangular and elliptic) and find the benefits and the disadvantages of them by comparing the simulated distortions caused by the non-uniform cooling. 2. Lab experiments Two sets of experiments had been done by Lechler GmbH to illustrate two relevant specifications of the nozzles: 1. Footprint: A useful experiment is “the impact force” test of the spray [17], which is measured based on a resistance strain gage [18] and [19]. The impact force addresses the water flow rate, the nozzle feed pressure and the impingement area over the cooled surface [20]. The impact force test indirectly characterizes the cooling intensity experienced directly on the surface [21] (higher flow rate and feed pressure had higher impact
Fig. 1. The water spray cooling regimes (“a” and “q” are the heat transfer coefficient and the heat flux, respectively) [3].
16
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
force on the surface [20] and consequently higher cooling intensity [20] and [22]). Therefore this experiment was used to identify the footprint. The impact force value and distribution for the selected nozzle settings are illustrated in Figs. 2 and 3. As it is shown in the figures, the impact force distributed almost evenly for the Bertin nozzle, while the other nozzles gives higher impact force at the nozzle center. 2. HTC: In another set of experiments, the temperature close to the surface of a sample was measured and the cooling histories of the cooled surface for both nozzle types were recorded. Then the HTC of every experiment was calculated by solving the heat transfer equations (by the finite difference method) in the range of 100 Ce500 C (Fig. 4) [23].
the effect of the cooling pattern (uniform and elliptic) on the distortion without interference from the HTC difference. In the second step, both the characteristic HTC and the pattern introduced by each nozzle were considered. Therefore, in this step, the HTC and the cooling pattern of the rectangular pattern were the same as in the first step (low HTC and uniform footprint), whereas the other one had a higher HTC (Fig. 4a) and an elliptic footprint. The goal of this step was to study the combined effect of the cooling pattern and the HTC on the resultant distortion. Some parameters for arrangement of the elliptic type footprints were selected (Fig. 6).
As shown in "fig2 fig3 fig4"Figs. 2e4, both the HTC and the impact force of the elliptic footprint were higher than those of the rectangular one. For the rectangular footprint, the impact force was almost uniform (Fig. 3). Therefore, for the finite element simulation, it was considered to be ideally uniform over the sprayed surface. For the other footprint, the pattern was idealized by considering it as an elliptic pattern with a Gaussian distribution that had a maximum value at the center and gradually decreased to zero toward the boundary of the ellipse. The cooling pattern over the sample surface was shown in Fig. 5. This is more clarified later in this paper.
This was the angle of the ellipse major diameter with the extrusion direction (x-axis). This parameter is responsible for adjustment of the heat transfer distribution over the surface. Its values were considered from 23 up to 80 . 3.2. Ellipse pitch
3. Investigation method and modeling case design
3.3. Major diameter
A 350 200 2 mm3 (length width thickness) free plate was to be quenched from the top surface by the moving cooling source. The initial temperature of the sample and the sprayed water were 540 C and 17 C, respectively. The cooling source was moving along the length of the sample with a 0.25 mm/s velocity. The cooling source was introduced by the HTC as a function of the surface temperature of the selected section (T) and positions on the surface (x,y):
The size of the transverse (larger) diameter of the elliptic pattern was named by “2a”. In this paper, only this diameter (2a) was changing between 200 e350 mm while the other one (2b) was considered constant. As 2a increased, the ellipse was elongated (the ellipse eccentricity approached 1).
HTC ¼ f ðT; x; yÞ The cooling pattern was either uniform or elliptic. Both HTC curves were simplified to decrease the calculation time (see the green line and the blue line in Fig. 4 in web version). As mentioned before, the HTC curves and the footprints of the nozzle examples were significantly different. The changing shape response could be caused by either the differences in HTCs or the footprints. To separate these factors, the study was done in two steps. In the first step, one of the simplified HTCs (simplified Bertin in Fig. 4b) together with different footprints were considered to study
3.1. Ellipse angle
The center-to-center distance of the adjacent ellipses (or the nozzles) across the extrudate width direction (y-axis) was called the “pitch” (Fig. 6). This parameter was studied in the range of 84 e 130 mm.
3.4. Minor diameter The conjugate diameter 2b was 52 mm for all studied cases. A “reference case” with suggested values for the selected parameters by the manufacturer was also defined (Table 1). To study the first and the third parameters (Angle and 2a), the pitch was chosen as small as possible without overlapping of two adjacent ellipses. 4. Material and process parameters The sample material was aluminum alloy 6082. The physical properties of the alloy are given in Table 2. The mechanical properties (including strength, elastic modulus and Poisson ratio) were
Fig. 2. The impact force distribution over the surface for the selected Lechler nozzle (water pressure ¼ 4 bar, air pressure ¼ 4 bar, distance between nozzle and cooled surface ¼ 200 mm).
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
17
Fig. 3. The impact force distribution over the surface for the Bertin nozzle (water pressure ¼ 2 bar, air pressure ¼ 0.1 bar, distance between nozzle and cooled surface ¼ 200 mm).
Fig. 4. Real (an example) and simplified (the implemented and claimed by the manufacturer) HTC curves of a) Lechler and b) Bertin nozzles.
5. FE model The program ABAQUS/Standard was used to simulate the cooling process and the distortions in 3D. The analysis type was “coupled temperature-displacement”, with the brick element, called C3D20T in the ABAQUS element library. This is a second order hexahedral solid element with 20 nodes. Four elements were used throughout the thickness (Fig. 8) [13,14]. The elasticeplastic material model was used in order to introduce the selected material behavior under thermal loads. The data achieved from laboratorial tests was entered as a table in the software. The software solves the following equation in order to calculate temperature distribution in the sample [24]. Fig. 5. Elliptic cooling pattern over the considered sample.
defined to be temperature and strain rate dependent. The temperature dependency of the elastic modulus and the strengths (yield and ultimate strengths) are shown in Fig. 7. The extrusion process parameters were selected close to the real process in the plants (Table 3).
VðkVTÞ ¼
rcp vT vt
Table 1 Details of the reference case and the selected parameters. Parameter
Range
Reference case
Angle ( ) Pitch (mm) 2a (mm) 2b (mm)
23e80 84e130 200e350 52
50 84 270 52
Table 2 Physical properties of the selected alloy.
Fig. 6. The selected parameters of the elliptic footprint (dimensions are in mm).
Density [kg/m3]
Thermal expansion coefficient [1/ C]
Thermal conductivity [W/m K]
Specific heat capacity [J/(kg C)]
2700
2.55e-05
200
900
18
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
80 70
250
Stress (MPa)
60 200
50
150
40 30
100
20
50
Elastic Modulus (GPa)
300
YS, SR= 0 YS, SR= 1 UTS, SR= 0 UTS, SR= 1 EM
10 YS=yield stress (plastic strain=0)
0
0
0
200 400 Temperature (°C)
600
UTS=ultimate stress (plastic strain=31%) SR= strain rate EM= elastic modulus
Fig. 7. Stress and elastic modulus curve versus temperature for different plastic strains.
Table 3 Extrusion process parameters. Run-out speed [m/s]
Initial profile temperature [ C]
Water temperature [ C]
Ambient temperature [ C]
0.25
540
17
17
where r is the density, cp is the specific heat capacity, k is the thermal conductivity. T and t are temperature and time, respectively. For coupled temperature-displacement systems ABAQUS solves a system shown as following [24]:
Kuu KTu
KuT KTT
Du DT
¼
Ru RT
where Du and DT are corrections to the incremental displacement and temperature, respectively. Kij are the sub-matrices of the coupled stiffness matrices and Ru and RT are the thermal residual vectors, respectively. These systems are solved simultaneously using the full Newton’s method. The cooling source was moving along the x-axis from the puller side (right hand side in Fig. 8) toward the die side. As previously mentioned, the studied case was a free plate, and the die and the puller sides were only about cooling source moving direction; there was no mechanical boundary condition representing the die and the puller. The FILM subroutine was used to define the moving cooling source with either the uniform or the elliptic footprint. This
subroutine was used to provide heat transfer coefficients and sink temperatures for fully coupled thermal-stress analysis. An HTC (called “h” in this subroutine) could be conditionally defined versus temperature and coordinates. Therefore an elliptic footprint was introduced by ellipse function and conditional statements (if... then) over the surface coordinates (x and y). Other conditional statements were used to define the HTC dependency on the surface temperature. The elliptic footprint was addressed by means of several scaledconcentric ellipses representing the cooling intensity boundaries. The resultant HTC and the heat flux distribution are shown in Fig. 9. In the studied cases, several similar distributions moved from the one end to the other end along the x-axis (similar to Fig. 5). 5.1. Boundary conditions The plate was fixed at the width middle points of the two edges in the y- and z-directions (Fig. 8) to represent a free plate behavior. Simultaneously, the edge at the puller side was constrained in the x-direction to make the model more sensitive to the distortion. Generally, thermal boundary condition is defined as below:
kvT ¼ hðT T0 Þ vn where n is the normal direction of the sprayed surface, h is the HTC between the surface and the cooling water, T0 is the water temperature and T is the surface temperature of the sample.
Fig. 8. The prepared meshed model for FE simulation (“Ui” is displacement in “i” direction).
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
19
Fig. 9. a) The HTC distribution b) the heat flux distribution on the surface evaluated in the FE model.
A uniform initial temperature was set at 540 C throughout the sample. 6. Results and discussion 6.1. First step e the same (and lower) HTC The goal of this step was to investigate on the effect of the footprints on the resulted distortions due to cooling. Therefore, the samples were cooled by both the rectangular and the elliptic types. As previously mentioned, the HTC was the simplified Bertin’s HTC (Fig. 4b). An example of the temperature distribution over the cooled surface was shown in Fig. 10. The main interest was the induced distortions (the flat sample’s out of flatness due to cooling) across the width. Because it was difficult to compare the total 3D distortions, we decided to focus on the distortion at the crosssection at the puller side (right side in Fig. 10). To avoid the
influence of the constraint on the right-hand side (Fig. 8), the crosssection was selected 8 mm away from the edge of the puller side (Fig. 10). This position was representative for the sample distortion and was used to compare the distortions at different cooling conditions. The corner points of the cross-section were also used to monitor the cooling history. Cooling with elliptic footprint and lower HTC (like Bertin nozzle’s) gave significantly smaller distortions compared to the uniform spray distribution (as it was shown in Figs. 11 and 12). This result can be explained by reviewing the cooling history (Fig. 13). Cooling by the rectangular pattern (uniform footprint) was more severe and caused higher temperature gradients both through the thickness and along the length. While the elliptic footprint had a smaller average cooling intensity over the surface than the other one. For the elliptic footprint, the cooling was severe only at the central zone of cooling pattern and made the total cooling intensity (over the surface) gentler. However, this was not always positive because poor mechanical properties can be obtained for the heattreatable alloys with thick wall. 6.1.1. Study of the effects of the parameters The parameters of the elliptic pattern (Table 1 and Fig. 6) were also studied, and the results were as follows. 6.1.1.1. Angle. This parameter affected the distortion significantly (Fig. 14a). If the angle was not chosen properly (e.g., 80 ), the final distortion caused by the elliptic footprint was much worse than that by the rectangular one. Between 23 and 50 , better results were obtained. As shown by the red curve (in web version) in Fig. 14b, the 50 angle gave the least distortion. When the angle was
Fig. 10. The temperature distribution on the surface cooled by a) a rectangular footprint at time ¼ 4.3 s and b) an elliptic footprint at time ¼ 8.8 s (maximum distortion time).
Fig. 11. Comparison of the corner displacement history during the cooling process.
20
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
Fig. 15. The temperature distribution when the angle is 80 .
Fig. 12. Comparison of the residual distortion of sections cooled by the elliptic and the rectangular cooling patterns.
Fig. 16. The effect of the pitch on the distortion.
Fig. 13. Comparison of the cooling history at the puller side corners (Fig. 10).
80 , the cooling was concentrated in the middle of the plate (Fig. 15) and resulted in large distortions due to the severe non-uniformity across the width (y-axis). 6.1.1.2. Pitch. This parameter had a similar effect to the angle and should be set as small as possible to have minimum distortion (Fig. 16). Larger pitches introduced more non-uniformity across the width. When the pitch was large, the cooling was concentrated closer to the outer edges of the plate. 6.1.1.3. 2a. This parameter had an insignificant effect compared to the other parameters (Fig. 17). However, when the ellipses were elongated (2a increased), the footprint was more homogenous, many ellipses contributed in the cooling process across the width
(a)
Fig. 17. The major diameter effect on the final distortion.
(b)
Fig. 14. The effect of the angle on the final distortion.
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
21
and the cooling uniformity across the width improved. Its sensitivity was not that strong within the selected limits. 6.2. Second step e realistic HTCs for each footprint
Fig. 18. The cooling history for the selected footprints.
As it was mentioned in the section “Lab experiments”, the selected example of the nozzles for the elliptic footprint provided higher HTC than the other footprint’s example. In this part, the higher HTC (Fig. 4a) was applied to the elliptic footprint while the HTC of the rectangular footprint was same as the previous step. Therefore, in this step the combined effect of the HTCs and the footprints were considered. Comparing the cooling history of the elliptic footprint with the rectangular footprint’s (with lower HTC as it was in the first step), one can see that the average HTC of the elliptic footprint was larger than that of the other one (Fig. 18), resulting in larger distortions as well (Fig. 19). This result is caused by the higher temperature gradients through the thickness and the higher non-uniformity of the cooling across the width (Fig. 20). Consequently, approaching to flexible uniform cooling with proper HTC is ideal as it can minimize the distortions due to cooling with providing required mechanical properties. In case of an elliptic footprint usage, the arrangement of nozzles needs to be optimized for geometrically accurate and cooling-ratesensitive sections. 7. Discussion
Fig. 19. The distortion across the width due to cooling for the rectangular footprint (HTCmax ¼ 10,000 W/(m2 K)) and the elliptic footprint (HTCmax ¼ 30,000 W/(m2K)).
The distortions induced to the section were introduced by local contractions on the plate surface and throughout the thickness. These distortions were caused by uneven cooling. When a set of nozzles with elliptic footprints was applied to cool down the sample, the temperature distribution (or the cooling intensity) was complex and resultant of the effect of every individual footprint
Fig. 20. The history of the maximum temperature difference across the width for the elliptic footprint a) high HTC b) low HTC.
Fig. 21. Temperature distribution over the plate surface cooled with the set of the elliptic footprint at a) time ¼ 1 s b) Time ¼ 2 s.
22
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
Fig. 22. Effect of pitch on the resultant HTC fraction.
(Fig. 21). As shown in the figure, the cooling concentration started at three zones (indicated in Fig. 21a by arrows) across the width, forming by three elliptic footprint central areas. The areas outside of these zones were cooled gently until the following footprints covered them (Fig. 21b). The simulation results showed that the elliptic cooling pattern with higher HTC (30,000 W/(m2 K) in this study) resulted in larger distortions and was not a proper choice for the shape sensitive sections. The distortions were significantly reduced when the lower HTC (maximum 10,000 W/(m2 K)) was applied. This solution is not suitable for the thick walled sections or alloys demanding faster cooling. As it was demonstrated in previous section, to decrease the distortion induced by cooling, the minimum value of the pitch (when there is no overlapping) was desired. According to the nozzles manufacturers’ experience, and the article by Hall and Mudawar [21], the water-flux in the overlapped area is the sum of the water flux of the overlapped nozzles. Mudawar and Valentine in 1989 [25] showed that the water flux has a dominant effect on the heat transfer. These postulates allow us to suggest that the resultant HTC of the adjacent nozzles can be as shown in Fig. 22. In this case, the distribution becomes more uniform across the width. This result can be optimized and a uniform distributed and a higher HTC can be reached (the right hand side in Fig. 22). With this technique, we can also improve the flexibility (e.g., the cooling controllability across the section width) of the elliptic nozzles and the productivity (higher HTC will allow to cool the sections faster). An essential advantage can be gained if each nozzle is controlled individually. In this case, the cooling can be tailored to complex sections such as those with non-uniform thickness (Fig. 23).
8. Conclusion We simulated the cooling processes by the elliptic and the rectangular footprints and compared the induced distortions due to such cooling processes. We also studied the effect of the parameters in the elliptic case. The simulation results showed that the final distortions introduced by the elliptic footprint were less than those introduced by the rectangular type, assuming that both have the same lower level of HTC (maximum 10,000 W/(m2 K)). However, according to the experimental measurements, the elliptic type manufactured by Lechler had a higher level of HTC (30,000 W/ (m2 K)) than the rectangular type. This difference was valuable from the metallurgical point of view for thick sections and special alloys. However, the elliptic distribution of the water-flux with higher HTC was not advantageous for the geometrically sensitive sections because it introduced larger distortions to the studied section due to the higher cooling non-uniformities across the width. This non-uniformity can be eliminated by using the elliptic nozzles with an overlapping area and creating a uniform distribution of water-flux across the width with higher HTC. The parametric study of the elliptic type footprint showed the followings: The ellipse angle had an optimum value of 50 in this paper. If higher angles were set, the cooling was concentrated at the middle across the width. Therefore, high temperature gradients across the width led to larger distortions. The pitch was the second studied parameter. As it was decreasing (the nozzles were getting closer), the final distortion was also decreasing. The transverse diameter, which was called “2a” in this paper, had minor effect. However, larger values (elongated ellipse) were in favor for establishing homogeneous distribution of cooling across the width and for decreasing the distortion. Based on the facts presented here we can also draw the following general conclusions: Gentle (softer) cooling caused smaller distortions. Cooling footprints should not leave any area of the profile surface unsprayed. That was the reason for extreme distortions found for the cases with 90 angle and too high pitch. Acknowledgements
Fig. 23. Controlling cooling intensity across the width.
This work was a part of the project run by Hydro in cooperation with SINTEF Materials and Chemistry. The work is partly funded by the Research Council of Norway (RCN). The authors would like to greatly acknowledge their support. Part of the experimental data
S. Bikass et al. / Applied Thermal Engineering 57 (2013) 14e23
was supplied acknowledged.
by
Lechler
GmbH,
which
hereby
is
also
References [1] J. Rivallin, S. Viannay, General principles of controlled water cooling for metallurgical on-line hot rolling processes, forced flow and sprayed surfaces with film boiling regime and rewetting phenomena, International Journal of Thermal Sciences 40 (2001) 263e272. [2] M. De Oliveira, J. Ward, D.R. Garwood, R.A. Wallis, Quenching of aerospace forgings from high temperatures using air-assisted atomized water sprays, Journal of Materials Engineering and Performance 11 (2002) 80e85. [3] T.A. Shedd, Next generation spray cooling, high heat flux management in compact spaces, Heat Transfer Engineering 28 (2007) 87e92. [4] S.S. Hsieh, T.C. Fan, H.H. Tsai, Spray cooling characteristics of water and R134a. Part I: nucleate boiling, International Journal of Heat and Mass Transfer 47 (2004) 5703e5712. [5] M. Plata, J. Piwnik, A. Patejuk, Modern precise high-power water-cooling systems for press quenching, Archives of Foundry Engineering 9 (2009) 205e210. [6] J.D. Bernardin, C.J. Stebbins, I. Mudawar, Mapping of impact and heat transfer regimes of water drops impinging on a polished surface, International Journal of Heat and Mass Transfer 40 (1997) 247e267. [7] Full Cone Nozzles, General Industry, Products, Lechler GmbH, www.lechler. com, Metzingen, Germany. [8] N. Järvstråt, S. Tjøtta, A process model for on-line quenching of aluminum extrusions, Metallurgical and Materials Transactions B 27 (1996) 501e508. [9] R. Becker, M.E. Karabin, J.C. Liu, Distortion and residual stress in quenched aluminum bars, Journal of Applied Mechanics 63 (1996) 699. [10] M. Kristoffersen, Studies of Shape Variations During Cooling of Flat Extruded Sections. PhD thesis, University of Oslo, Norway, 2004. [11] Y. Kaymak, Simulation of Metal Quenching Processes for the Minimization of Distortion and Stresses, Otto-von-Guericke-Universität Magdeburg, Germany, 2007. [12] R. Pietzsch, M. Brzoza, Y. Kaymak, E. Specht, A. Bertram, Simulation of the distortion of long steel profiles during cooling, Journal of Applied Mechanics 74 (2007) 427.
23
[13] T. Walle, Shape Variations Due to Cooling. PhD thesis, University of Oslo, Norway, 2007. [14] S. Bikass, B. Andersson, A. Pilipenko, H.P. Langtangen, Simulation of the distortion mechanisms due to non-uniform cooling in the aluminum extrusion process, International Journal of Thermal Sciences 52 (2012) 50e58. [15] R.H. Pereira, S.L. Braga, J.A.R. Parise, Single phase cooling of large surfaces with square arrays of impinging water sprays, Applied Thermal Engineering 36 (2012) 161e170. [16] W.J.J. Vorster, S.A. Schwindt, J. Schupp, A.M. Korsunsky, Analysis of the spray field development on a vertical surface during water spray-quenching using a flat spray nozzle, Applied Thermal Engineering 29 (2009) 1406e 1416. [17] Nozzle Performance and Service Data, Metallurgical Industry, Nozzle Technology, Lechler GmbH, www.lechler.com, Metzingen, Germany. [18] S. Schürmann, Measurement and mathematical approximation of the impact of descaling nozzles, in: Hydraulic Descaling Conference, 2000, London, UK. [19] E.U. Becker, G. Birkemeier, W. Buchele, M. Degner, L. Devrient, M. Nowak, G. Thiemann, Optimization of descaling nozzles in a hot strip mill, Metallurgical Plant and Technology International 23 (2000) 92e94. [20] L. Bendig, M. Raudenský, J. Horský, Descaling with High Pressure Nozzles, ILASS-Europe, Zurich, Switzerland, 2001. [21] D.D. Hall, I. Mudawar, Experimental and numerical study of quenching complex-shaped metallic alloys with multiple, overlapping sprays, International Journal of Heat and Mass Transfer 38 (1995) 1201e1216. [22] D. Tucker, A Quantitative Study Into the Comparative Efficiency of Flat Jet and Full Cone Spray Nozzles When Applied to the Application of Work Roll Cooling, Service, Lechler GmbH, www.lechler.com, Metzingen, Germany. [23] S. Bikass, B. Andersson, A. Pilipenko, Uncertainties on HTC measurement of water spray quenching of aluminum alloys, in: ASME/JSME 2011 8th Thermal Engineering Joint Conference, AJTEC2011, 2011, T10155-T10155-5, Honolulu, Hawaii, USA. [24] ABAQUS Standard Version 6.5, User Manual, Theory Manual, Hibbitt, Karlsson & Sorensen Inc, USA, 2005. [25] I. Mudawar, W. Valentine, Determination of the local quench curve for spraycooled metallic surfaces, Journal of Heat Treating 7 (1989) 107e121.