Comparison of Impact, Velocity, Drop Size and Heat Flux … · Comparison of Impact, Velocity, Drop...

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Comparison of Impact, Velocity, Drop Size and Heat Flux to Redefine

Nozzle Performance in the Caster

Kristy TannerSpraying Systems Co.

Wheaton, IL 60187 USA

Abstract

Impact testing performed to empirically determine the impact and velocity of airmist nozzles used in continuous casting. Previous data from heat flux testing above Leidenfrost temperature is used to compare the affects of impact, velocity, dropsize on heat flux. A close look at the

film boiling regime shows how impact and velocity can influence localized cooling. Flux distribution is compared to impact distribution to show different parameters defining traditional air mist nozzle performance.

As presented at: American Iron and Steel Technology Conference in Nashville, TN, September 2004

Materials Science and Technology Conference in New Orleans, September 2004

2 www.sprayconsultants.com

Comparison of Impact, Velocity, Drop Size and Heat Flux to Redefine Nozzle Performance in the Caster

Introduction

Spray performance is very important in final steel quality. Traditionally, air mist nozzles used in the caster are defined with spray density and spray density distribution. Previous data from heat flux testing above the Leidenfrost temperature along with spray density and density distribution is examined to determine if even cooling is achieved throughout the spray as the density distribution would suggest. Furthermore, drop size data and impact distributions are reviewed and relationships are established with respect to heat flux of the spray nozzle at given operating conditions. A close look at the film boiling regime shows how impact and velocity can influence localized cooling. Overall, new parameters need to be reviewed to evaluate air mist nozzle performance in the caster. Hydraulic nozzles will also be examined to determine how dropsize and velocity can affect heat transfer characteristics.

As seen in previous testing [1], even cooling is required to prevent destabilization of the thin steam layer between the spray water and the surface of the steel. During film boiling it is critical not to disrupt

the stable steam layer and prevent localized cooling of the steel surface. Localized cooling will lead to cracking of the final product and steel quality is then compromised. Studies have found during film boiling, the heat transfer coefficient is more dependent on the mass water flux and less dependent on the steel surface temperature. It is important to find the correct mass water flux to optimize heat transfer while preventing localized cooling. Drop size and impact also need to be taken into consideration when looking at cooling in the film boiling regime as well. A relationship must be drawn between dropsize, impact and heat transfer to create a reliable model for evaluating heat transfer rates in continuous casting.

Previous testing has focused on low density nozzles [3-5]. However, larger densities which are typically used in industrial applications have not been examined extensively. Heat transfer rates of larger density nozzles along with corresponding dropsize, impact and velocity needs further study to determine the existing relationships.

Nozzle Testing Matrix

Traditionally airmist nozzles are used for spray cooling in continuous slab casting as well as billet and bloom casting. The nozzles used for study are Spraying Systems Company’s Next Generation CasterJet nozzles. These are air assisted flat spray nozzles and have an even spray density distribution. Sprays are positioned in the caster to slightly overlap and create on overall even spray density across the steel surface.

The airmist nozzle chosen has an air inlet entering the nozzle in line with the exiting spray tip. The water inlet enters 90° to the air inlet and impinges on a concave target bolt surface. The water and air then mix in the mixing chamber and exit at the tip orifice. Testing was performed on the following tip and inlet combinations: 3/8NCJ-2.5-110-SS, 3/8NCJ-3-120-SS, 3/8NCJ-4-110-SS, 3/8NCJ-5-120-SS and 3/8NCJ-6.5-95-SS. Water

pressure ranged from 5 to 100 psi (.34 to 6.89 bar) and air pressure ranged from 10 to 60 psi (.69 to 4.14 bar).

Hydraulic nozzles were also studied to strengthen the database and provide further depth to the relationships established. Fullcone nozzles, hydraulic flat sprays and extra thick flat sprays were all tested at pressures from 10 to 100 psi (.69 to 6.89 bar). These tests gave a much larger dropsize range than using airmist nozzles only. It also provided an argument comparing hydraulic to airmist nozzles based on heat transfer rates achieved.

Testing Procedures

Heat Transfer Testing [1]

Heat transfer testing was performed using circular stainless steel plates 4.92 inches (125 mm) in diameter and 0.98 inches (25 mm) thick. The plates are heated to 1650°F in a laboratory box furnace, shown in Figure

EQUIPMENT & METHODS



EQUIPMENT & METHODS

1, then moved via a rail to the spray cooling chamber. The testing plates are placed directly under the spraying nozzle. Two thermal couples are placed in the plate. Both are positioned 0.098 (2.5 mm) inches from the surface of the plate. This is as close to the steel surface as possible for an accurate reading. If they are placed too close to the surface, the reading is just recording the heat of the thermal couple material and not the steel. One thermal couple is directly in the center of the plate. The other is located 2 inches (50.8 mm) off center. Two additional thermal couples are placed inches below the first two. The temperature readings at these different depths show heat conduction could be treated as one dimensional. Direct heat conduction calculations are used for modeling between thermal couples while inverse heat conduction calculations are used for modeling between the top thermal couples and the surface of the steel.

Direct conduction can be used for modeling in a steady state condition where there is no loss of heat from the side or bottom of the plate.

( )tTCq

x px ∂∂

=∂∂ ρ 1 -dimensional

Transient Conduction

xTkqx ∂∂

= First Law of Conductivity

In these equations T is temperature, t is the time

variable, x is the spatial variable, Cp is specific heat and q is heat flux. This numerical scheme has been checked for stability and convergence.

Since conditions are only known in the direct conduction region, inverse conduction modeling must be used to determine the ultimate goal of surface temperature. The Weber method is used for solving this inverse heat conduction problem:

tT

xT

∂∂

=∂∂

2

2

α

Alpha (α) is the thermal diffusivity, T is temperature, t is the time variable and x is the special variable. The use of the inverse conduction numerical scheme is necessary to determine the maximum heat flux and to identify an accurate Leidenfrost point. If direct conduction is used for the region between the top thermal couples and the steel surface, the maximum heat flux and Leidenfrost points will be grossly distorted.

During testing, data is recorded using a data collection system and a database is created for all testing. Testing is conducted for 300 seconds with a time variable of 0.1 seconds. This gives us data to derive a heat transfer curve for each nozzle at each testing condition. This clearly defines the separate regimes of pool boiling heat transfer for each test. At high surface temperatures, where film boiling is dominant, a vapor film adjacent to the hot surface prevents or minimizes intimate droplet contact with the surface. This results in low heat transfer. The heat flux decreases only slightly with decreasing temperature. As surface temperature gets lower, the droplets begin to penetrate the vapor film and a sharp increase in heat transfer begins. This is the start of the transition boiling heat transfer regime. Just before transition boiling the heat flux experiences a minimum at the Leidenfrost point. Sharp increases in heat flux due to surface wetting is obvious below the Leidenfrost temperature. After reaching the maximum critical heat transfer rate the curve starts decreasing as it goes through Nucleate boiling and finally convection cooling close to the saturation temperature as clearly defined through testing and evaluation.

Figure 1. Steel Plate Transfer to Rail



EQUIPMENT & METHODS

Drop Size Testing

Testing was done using a Malvern particle analyzer. This is a special sampling device that measures how much light scatters as it moves through the droplet to determine the diameter of the droplet. The intensity of the scattered light is measured with photo diodes. Both the volume mean diameter (VMD) and the sauter mean diameter (SMD). The VMD is the drop size with respect to the volume of water sprayed. The value obtained is that diameter where 50% of the total volume has drops with larger diameters and 50% has drops with smaller diameters. The SMD is calculated using the surface area. This is the diameter of the drop with the volume to surface ratio the same as the total volume to total surface area of all drops.

Drop size testing was performed on nozzles in the middle of the spray as well as 2 inches (50.8 mm) from the edge of the spray at a 6 inch (152.4 mm) spray height.

Impact Testing

Total theoretical impact of the nozzle can be calculated by:

Im𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑡𝑡𝑝𝑝𝑝𝑝𝑡𝑡𝑝𝑝ℎ𝑒𝑒𝑡𝑡 = 0.0527∗ 𝑓𝑓𝑡𝑡𝑡𝑡𝑓𝑓𝑓𝑓𝑝𝑝𝑝𝑝𝑒𝑒∗ Pr𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑓𝑓𝑒𝑒

However, the distribution of the total impact is important, especially in the film boiling regime where localized cooling is avoided. High impact could break through the steam layer and compromise the final quality of the steel.

Impact testing is done with a patent pending process used to measure load throughout the entire spray pattern, shown in Figure 2. A small pin is attached to a load cell and moves throughout the spray pattern.

The load cell is mounted on a test grid railing where two small servo motors move the load cell in the x and y directions. The load cell measures the force of the spray pushing downwards on the top surface of the pin. This load is recorded using a data acquisition program. Data is then graphed to determine the impact distribution across the spray. Spacing between recorded data points can be set by the operator to control resolution of graphs.

Figure 2. CasterJet Spraying During Impact Testing

Impact tests were performed at a six inch spray height. Step size (distance between force readings) was 0.10 inches (2.54 mm). When comparing impact test results between nozzles it is important to compare tests of same resolution or step size. A large step size may smooth out any low or high readings and make the distribution look more even than it actually is.

RESULTS & DISCUSSION

Testing Results

Looking at each nozzle type individually it is observed that an increase in mass flux leads to a direct increase in heat flux. This trend was seen on all nozzle

types (both hydraulic and airmist). If nozzle types are compared to one another however, a different trend is observed. Air mist nozzles had higher heat flux than hydraulic nozzles for a given mass flux range.



0

500

1000

1500

2000

2500

3000

0 5 10 15 20

Spray Density (kg/m2s)

Qm

ax (k

w/m

2)

3-3.25

4.4-4.75

9.5-9.6

12-12.5

14-15

Air mist Nozzles

Hydraulic Nozzles

Figure 4. Spray Density vs. Qmax for Selected Density Groups Comparing Air Mist Nozzles to Hydraulic Nozzles

Figure 3. Spray Density vs. Qmin for Selected Density Groups Comparing Air Mist Nozzles to Hydraulic Nozzles

0

200

400

600

800

1000

1200

1400

1600

0 5 10 15 20


Qm

in (k

w/m

2) 3-3.25

4.4-4.75

9.5-9.6

12-12.5

14-15

Air mist nozzles

Hydraulic Nozzles


At high surface temperatures, where film boiling is dominant, a vapor film adjacent to the hot surface prevents or minimizes intimate droplet contact with the surface. This results in low heat transfer. The heat flux decreases only slightly with decreasing temperature. As surface temperature gets lower, the droplets begin to penetrate the vapor film and a sharp increase in heat transfer begins. This is the start of the transition boiling heat transfer regime. Just before transition boiling the heat flux experiences a minimum at the Leidenfrost point. Figure 3 shows the Leidenfrost point (Qmin) for five groups of airmist and hydraulic nozzles with the same spray densities. It is

observed that the air mist nozzles have a higher heat flux at the Leidenfrost point than the hydraulic nozzles. Sharp increases in heat flux due to surface wetting is obvious below the Liedenfrost temperature. After reaching the maximum critical heat transfer rate the curve starts decreasing as it goes through Nucleate boiling and finally convection cooling close to the saturation temperature. Figure 4 shows the same nozzles as in Figure 3 but compares the maximum heat flux transitioning into nucleate boiling. Air mist nozzles have higher Qmax values however, as the spray density decreases, the difference between the airmist and hydraulic becomes less obvious.




After observing the difference between airmist and hydraulic nozzle heat transfer rates, we now look at the drop size difference with the same nozzles. Figure 5 shows the drop size for the same groups of nozzles. The hydraulic nozzles have much larger drops, and in turn smaller heat transfer rates. Note, drop size difference is of the same magnitude for the lower

density nozzles but there is not a large difference in the heat transfer rate. Drop size is a contributing factor in heat transfer rates for larger density nozzles.

Figure 6 compares the Leidenfrost temperature of the same nozzles. The Leidenfrost temperature does not differ greatly for any nozzle density group, hydraulic or airmist.

0

100

200

300

400

500

600

700

0 5 10 15 20


Leid

enfro

st T

empe

ratu

re (C

)

3-3.25

4.4-4.75

9.5-9.6

12-12.5

14-15

Figure 6. Spray Density vs. Leidenfrost Temperature for Selected Density Groups For Air Mist and Hydraulic Nozzles

Figure 5. Spray Density vs. Dropsize for Selected Density Groups Comparing Air Mist Nozzles to Hydraulic Nozzles

1.0

10.0

100.0

1000.0

10000.0

0 5 10 15 20


Drop

size

VM

D (m

icro

ns)

3-3.25

4.4-4.75

9.5-9.6

12-12.5

14-15

Air mist Nozzles

Hydraulic Nozzles




Figure 7 shows the impact results for a 6.5-95 CasterJet nozzle. Figure 8 shows the density distribution for the same nozzle. The density is even across the coverage of the nozzle. However the impact distribution shows the middle to have a much higher impact that tapers out to the sides. A velocity distribution of the nozzle would match the impact distribution shape since this nozzle has equal droplet size across the spray pattern. Our heat transfer testing looked at the center and 2 inches from the center. The difference was very slight, but this is limited data. The next phase needs to include an entire heat transfer profile across the coverage of the

nozzle. It is common practice in industry to stagger nozzles from row to row in the caster. The surface becomes too cool under the center of the nozzles if there is no staggering. This suggests a relationship between heat transfer and impact. If the spray density alone could determine the distribution of the heat transfer coefficient of the spray, further study would not be needed. Ideally, to achieve a constant heat transfer coefficient throughout a spray pattern with constant drop size, the impact needs to remain constant. Another way to keep the heat transfer even may be to have a heavy edged density distribution where lower velocity droplets are on the edge.

0 1.5 3 4.5 6 7.5 9 10.5 12 13.5 15

0

0.001

0.002

0.003

0.004

0.005

0.006

0.007

0.008

0.009

0.01

Forc

e (lb

f)

Coverage (Inches)

Figure 7. Impact Test Results for 6.5-95 CasterJet Nozzle Spraying at 45 psi (3.1 bar) air and 100 psi (6.9 bar) water at a 6 inch (152 mm) Spray Height

0

1

2

3

4

5

6

7

8

9

10

% o

f Flo

w

Figure 8. Flow Distribution Test Results for 6.5-95 CasterJet Nozzle Spraying at 45 psi (3.1 bar) air and 100 psi (6.9 bar) water at a 6 inch (152 mm) Spray Height




Using the heat transfer data collected, we can look at the case of a fullcone nozzle. Fullcone nozzles are typically used in billet and bloom casting but can also be found in slab casting. For the 1/4 HHCC8 we have the same dropsize in the center as well as at 2 inches off center, VMD = 3198 microns. Also, spray density is similar (10% lower for side). The Qmax has a much larger difference than the Qmin between the center or the spray and 2 inches to the side. This could be due to the large difference in the impact between the center and the side. The impact is twice as large in the middle of the spray than on the side.

Center SideQmax 2050 1700Qmin 600 580Leidenfrost Temp. 650 700Dropsize 3198 3198Spray Density 4.5 4

Table 1. Values for 1/4HHCC8 Nozzle Spraying at 60 psi (4.1 bar) and a Spray Height of 6 inches (152 mm)

Figure 9 shows the density distribution for the 1/4 HHCC8 fullcone nozzle spraying at 60 psi (4.1 bar). Figure 10 shows a typical impact distribution for the HHCC fullcone nozzle. Table 1 compares values for the nozzle at the side and in the center of the spray.

Figure 9. Density Distribution for Fullcone Nozzle

Figure 10. Common Impact Distribution for HHCC Fullcone Nozzle

Conclusions

While spray density is the single largest factor for heat transfer in spray nozzles, it is apparent that velocity, dropsize and impact play an important roll as well. At low spray density (<8kg/m2s), the density of the spray seems to be the main factor in heat removal. At the lower densities, the heat transfer in the film boiling regime is less dependent on droplet size, impact and velocity. This is parallel to the Yao and Choi’s research of heat transfer of low mass density sprays [7]. While with larger spray density nozzles, other factors need to be examined. It is obvious the heat transfer of these sprays is strongly related to dropsize.

There was a strong relationship between Qmin and dropsize for all the nozzles tested, with dependency increasing with increasing spray density. Less of a relationship was found between Qmax and dropsize for low spray density nozzles. However, there was a larger dependency with increasing spray density.

Heat transfer is dependent on impact and velocity under constant dropsize and spray density with increasing dependency with increasing spray density. The impact and velocity relationship needs to be further studied to determine how spray density can be adjusted to even out the spray impact distribution.



References

1. Hamed M. Al-Ahamdi, S.C. Yao, Kristy Tanner and Ken Kasperski, “Experimental Study on the Spray Cooling of High Temperature Metal Using Full Cone Industrial Sprays,” ISS Conference, Nashville, TN (2002).

2. S.G. Hibbins and J.K. Brimacombe, “Characterization of Heat Transfer in the Secondary Cooling System of a Continuous Slab Caster,” Trans. ISS, Vol. 3, pp 77-89.

3. R.E. Gaugler, “Experimental Investigation of Spray Cooling of High Temperature Surfaces,” Doctoral thesis, Mech. Eng. Dept., Carnegie Mellon University, July (1966).

4. E. Mizikar, Spray cooling investigation for continuous casting of billets and blooms, Iron Steel Engng, June, 53-70 (1970).

5. M. Ciofalo, I. Piazza, , and V. Brucato, , Investigation of the cooling of hot walls by liquid water sprays, Int. J. Heat Mass Transfer 42, pp. 1157-1175, (1999).

6. S.C. Yao and K.J. Choi, “Heat Transfer Experiments of Mono-Dispersed Vertically Impacting Sprays”, Int. J. Multiphase Flow Vol. 12, No. 5, pp. 639-648 (1987).

7. K.J. Choi and S.C. Yoa, “Mechanisms of Film Boiling Heat Transfer of Normally Impacting Spray”, Int. J. Multiphase Flow Vol. 30, No. 2, pp 311-318 (1987).

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