In recent years, there has been a significant increase in the customer demands for improved induction coil lifetime. This has led to several publications in recent years by induction tooling manufacturers [1-4]. The main conclusion in these papers is that besides mechanical crashes the cause of most induction coil failures is localized overheating of the coil copper due to insufficient cooling. What is lacking from these publications is any way to determine what is sufficient cooling. In this paper, a scientific method for determining local copper temperatures will be presented. This will include evaluations of heat transfer coefficients for different sections of a multi-component inductor, dependence of heat transfer coefficient on water pressure and water passage cross-section, non-uniform power density distributions in various 2-D cross-sections and the resulting temperature distribution in the copper winding. The effects of duty cycle on optimal design will also be considered.

1. ASM International Conference, September 2007, Detroit
Increasing Inductor Lifetime by Predicting Coil Copper Temperatures
R. Goldstein, V. Nemkov
Fluxtrol Inc., Auburn Hills, MI, USA
Abstract
In recent years, there has been a significant increase in the customer demands for improved induction coil lifetime. This has led to several publications in recent years by induction tooling manufacturers [1-4]. The main conclusion in these papers is that besides mechanical crashes the cause of most induction coil failures is localized overheating of the coil copper due to insufficient cooling.
What is lacking from these publications is any way to determine what is sufficient cooling. In this paper, a scientific method for determining local copper temperatures will be presented. This will include evaluations of heat transfer coefficients for different sections of a multi-component inductor, dependence of heat transfer coefficient on water pressure and water passage cross-section, non-uniform power density distributions in various 2-D cross-sections and the resulting temperature distribution in the copper winding. The effects of duty cycle on optimal design will also be considered.
This method may be incorporated into the standard coil design and development process, which can be used to prevent costly tooling lifetime issues during the early stages of a production run or avoid the purchase and use of unnecessary oversized or booster pumps. It can also be used as a basis for more advanced future studies into temperature distributions and stresses within the induction coil itself, which lead to failure.
Introduction
Like most industrial processes, tooling failure is a leading cause of induction heating machine downtime. This leads to many challenges in today’s manufacturing environment. In most cases, this downtime is unplanned for and results in significant costs to ensure on-time delivery of parts to a customer.
Induction heat treating tooling failures can be broken into three main classes:
 Mechanical Damage
 Thermal Degradation
 Electrical Break
Electrical breaks may be caused by different factors: insufficient insulation between the coil turns, insulation wearing, magnetic chips attracted to the conductors etc. This failure mode may be prevented by proper design and maintenance of the coil [3].
Mechanical damage may be caused by inaccurate coil setup resulting in the part impact, by incorrect machine operation, by the electromagnetic forces and by the thermal distortion of the coil components. These factors can damage the coil instantaneously, breaking the coil integrity or causing water leakage or they can change the coil dimensions gradually. For example, coil copper may be out of specifications due to creeping, which can result in loss of the heat pattern. In many cases, mechanical failure is preventable with the proper precautions and maintenance on the machine.
Failures due to thermal degradation are more challenging to resolve. Thermal degradation is caused by local or total overheating of the coil head due to eddy-current losses in the copper, magnetic losses in the flux concentrator and by heat transfer from the hot surface of the part by convection and radiation. Overheating can result in copper cracking or deformation as well as the concentrator material degradation. Copper cracking happens usually in hardening coils with a short cycle due to thermal stresses, while a gradual coil deformation is more typical for continuous processes. Thermal effects can strongly increase the effects of electromagnetic forces and accelerate the electrical insulation aging.
Copper overheating is the leading cause of failure in heavy- loaded hardening inductors and the present article is focused mainly on this mode of failure. For consistently manufactured inductors, copper cracks occur in nearly the same place on an inductor each time in a certain range of parts produced. There are several approaches to increase the coil life: provide additional cooling, reduce the density of heat sources or change the coil design completely.
Additional cooling is provided by either increasing the water flow rate, by adjustment of the water pocket or introduction of additional cooling circuits. Water flow rate is the first step

2. and it will be increased until the pump output limit is reached. Inductor manufacturers generally have internal guidelines related to best practices for water pocket and cooling circuit design, which have been developed over the years. Once these standard options have been exhausted, the next step is to replace the existing pump with a larger one or introduce a booster pump to the system.
Sometimes on very high power density coils, a limit is reached; where even with the best cooling circuit design and very large pumps coil lifetime is still unsatisfactory. At this point, it is necessary to try to minimize the localized power density in the weak point of the inductor. This is oftentimes challenging in complex inductors, as changing this section may have some effect on the heat pattern in this area of the part, as well as in the rest of the part.
The response to thermal degradation failures of inductors are based upon practical experience and in many cases will require several iterations to resolve. There is significant cost associated with these each of these tests (prototype coils, additional pumps, machine downtime, test parts, met lab time, etc.), none of which were budgeted for.
A more scientific method, which would reduce the number of iterations and optimize equipment purchases is of great value. Furthermore, if this method is used up front in the coil development, it can avoid troubleshooting due to insufficient coil lifetime occurring during the early stages of production.
Computer simulation is a natural tool for analyzing all of these factors (fluid dynamics, heat transfer, electromagnetic, thermal, stresses and structural transformations) that go into inductor failure by thermal degradation. However, there is no one program on the market with strong 3-D coupling of all of these phenomena at this time.
At the same time, there are simulation tools and analytical formulae available now for tackling the problem in steps. The following sections describe such a methodology currently being used at Fluxtrol Inc. for predicting copper temperatures. This method is viewed as a good first approach to the problem and a basis for further improvements with future software improvements. A case story is shown to demonstrate representative results.
Technical Description
The current approach used by Fluxtrol Inc. is an iterative 7 step approach:
1. Electromagnetic + Thermal simulation to optimize part heating and coil parameters
2. Coil Engineering using CAD program
3. Analytical hydraulic calculations for the coil cooling circuit
4. Calculation of localized heat transfer coefficients in high power density areas of the inductor
5. Electromagnetic + Thermal simulation of coil component heating
6. If elevated component temperatures exist, return to step 2 to improve cooling circuit
7. If elevated component temperature exist, return to step 1 to improve induction coil geometry from a cooling perspective with minimal sacrifice in part heating or coil parameters
.
Computer simulation for optimization of the induction coil based upon the part heating and coil parameters should be done first. If the coil designed on this criterion will have satisfactory lifetime, it will lead to the minimum cost production.
In the coil engineering stage, the busswork, leads and complete inductor cooling circuit should be layed out. This is essential for the hydraulic calculations. The water cooling circuit can be broken down into several components. Pressure drops in every component of the circuit along with the additional pressure drop due to directional or flow passage geometry change should be calculated. These numbers will give a flow rate and localized water velocities based upon the pump pressure.
Using the water velocities and tubing cross-sections, it is possible to calculate Reynolds, Prandtl and Nusselt numbers for all of the different heating areas of the inductor. Heat transfer coefficients are calculated from the respective Nusselt numbers with equation 1.
ά = Nu k / De Equation 1
ά – heat transfer coefficient
Nu – Nusselt Number
k – thermal conductivity
De – Equivalent diameter
Using an Excel spreadsheet, it is possible make a good approximation for the hydraulic and heat transfer coefficient calculations. Both the hydraulic values and heat transfer coefficients are dependent upon the bulk temperature of the cooling water in that section of the inductor. It is possible to incorporate in this spreadsheet the power dissipated in the different sections of the inductor along with simple calorimetric calculations for local water temperature.
Using the base data derived from the first four steps, we now have all the required information to simulate the inductor temperatures. An electromagnetic plus thermal simulation program should be used for these calculations.
In many cases, the procedure will only be five steps and further improvement will not be required. All inductor temperatures will be acceptable and no further modification will be required.

3. In very heavy loaded applications, there will be a need to have an iterative process to resolve elevated temperatures in local areas of the inductor. The first step for this should be to return coil engineering and see if there is a way to improve the cooling circuit to improve the areas of the inductor with high temperature.
After updating the cooling circuit, the hydraulic and heat transfer coefficient calculations should be remade. Then simulation for the inductor temperature should be repeated.
If unacceptably high temperatures still exist in the same components of the inductor, then it is necessary to return to the first step, computer simulation for induction coil design. Changes will need to be made to the coil copper profile and it will be necessary to find a compromise between inductor performance and coil component heating. After coming up with a new design, the rest of the above steps need to be made.
This method should be repeated until satisfactory temperatures are reached in all components of the induction coil. To demonstrate this method, the case of a seam annealing coil after arc welding of heavy wall tube is considered.
Case Story – Seam Annealing
In the tube and pipe industry, heavy walled tubes are commonly arc welded. The arc welding process changes the structure of the material in the seam area and this is oftentimes restored using a local annealing process (seam annealing). Induction heating is the preferred method for seam annealing after welding.
Figures 1 and 2 shows parts of a coil drawing for seam annealing after spiral welding of large diameter tubes used in the oil and gas industry. Figure 1 shows the cross-section in the heating area. There is a recess in the middle of the inductor to allow room for the bead that is formed during the welding process. Figure 2 shows the top view of the inductor. In this projection, the inductor is shown as flat, but on this side view (not shown) it is arced to follow the contour of the pipe.
Figure 1: Cross-section of induction coil for spiral welded tube seam annealing
Figure 2: Top view of induction coil for spiral welded tube seam annealing
For the case story, we can consider heating of a large diameter pipe with ¼” wall thickness. The arclength of the “active” area (does not include the cross-overs or leads area) of the winding is 32.5”. The process is continuous with a feed rate of the pipe is 7.4” / second. Frequency used for simulation is 1 kHz. Two different types of magnetic flux controller, laminations and Fluxtrol A will be considered. Flux 2D electromagnetic plus thermal program will be used for all heating simulation.
Comparison will be made with the same maximum weld seam temperature of 1200 C (± 20 C) exiting the inductor and desired equalized annealing temperature of 1000 – 1050 C shortly after exiting the inductor. This area of high temperature should extend beyond both sides of the weld bead.
The first step in the procedure is electromagnetic and thermal simulation of the heating process. The coil is long and can be considered as a 2-D, plane parallel system. Due to symmettry, it is only necessary to simulate half of the induction coil.
Figure 3 shows the temperature of the pipe at the exit of the induction coil for a coil with laminations (left) and 3 seconds after exiting the coil (right). The inductor power required was 600 kW. The coil current was 18.8 kA in the main leg, 9.4 kA on each of the return legs.
Figure 3: Temperature distribution for the induction coil with laminations at the exit of the seam annealing coil(left) and 3 seconds later (right)
To compare the performance of laminations to Fluxtrol A, we’ll change the magnetic flux controller properties and the coil current required to reach the same temperature in the Color Shade ResultsQuantity : Temperature Deg. Celsius Time (s.) : 4.4 Phase (Deg): 0Scale / Color42.3855 / 115.03716115.03716 / 187.6888187.6888 / 260.34045260.34045 / 332.99213332.99213 / 405.6438405.6438 / 478.29541478.29541 / 550.94708550.94708 / 623.59875623.59875 / 696.25043696.25043 / 768.90204768.90204 / 841.55371841.55371 / 914.20538914.20538 / 986.85699986.85699 / 1.05951E31.05951E3 / 1.13216E31.13216E3 / 1.20481E3Color Shade ResultsQuantity : Temperature Deg. Celsius Time (s.) : 7.599999 Phase (Deg): 0Scale / Color45.1254 / 108.46457108.46457 / 171.80374171.80374 / 235.14291235.14291 / 298.48209298.48209 / 361.82126361.82126 / 425.16043425.16043 / 488.49957488.49957 / 551.83881551.83881 / 615.17792615.17792 / 678.51715678.51715 / 741.85626741.85626 / 805.1955805.1955 / 868.53461868.53461 / 931.87384931.87384 / 995.21295995.21295 / 1.05855E3

4. same amount of time. Figure 4 shows the temperature of the
pipe at the exit of the induction coil for a coil with Fluxtrol A
(left) and 3 seconds after exiting the coil (right). The
distribution is nearly identical. The inductor power required
was the same as for laminations, 600 kW. The coil current
was slightly higher, 20 kA in the main leg, 10 kA on each of
the return legs.
Figure 4: Temperature distribution for the induction coil with
Fluxtrol A at the exit of the seam annealing coil(left) and 3
seconds later (right)
After simulation of the part heating, the next step is the coil
engineering. We are using a real inductor where the drawings
are already completed, so it is not necessary in this case.
The next step is the hydraulic calculations. The magnetic flux
controller will have no effect on the hydraulic calculations.
There are 2 water inlets and 4 outlets on the inductor. Since
there are 2 separate water channels in the main leg, we can
consider the inductor as having 4 separate water circuits on
this inductor for the hydraulic calculations.
If we have 30 psi applied each of the 4 water circuits, the total
water flow rate should be around 23.2 gpm total or 5.8 gpm
per circuit. This calculation takes into account the pressure
drop in the inlet hose, the buss tubing, the bending areas and
the two active legs.
After the hydraulic calculations, we can derive the local heat
transfer coefficients. The areas of interest are the main
heating leg and the return leg. In the main water pocket, the
water velocity is 12 ft/sec and in the return leg the velocity is
close to 19 ft/sec. Putting the losses in the inductor sections
into the spreadsheet yields a temperature rise in the bulk water
of around 13º F (7º C), which does not have a significant
impact on heat transfer coefficients. Therefore, the heat
transfer coefficients for the main leg and return leg should be
14,000 and 22,500 W/m2K respectively.
To simulate the localized inductor temperatures, it is necessary
to consider the inductor construction for calculation of the
localized inductor loading. The power calculated in both
cases is the same. It is necessary to recalculate for the space
factor considering space unusable around the leads area, cross-overs
and copper keepers. This will not have an effect on the
total power, only on the required coil current.
For laminations, there needs to be a 1/8” copper keeper every
4”. Also, there should be at least a ¼” gap between the
laminate stack and the leads or the cross-overs to prevent
overheating from the 3-D magnetic fields. This means that the
effective length of the inductor is about 0.89 times ideal. To
achieve the results calculated above, it is necessary to increase
the coil current 5.7%.
For Fluxtrol A, copper keepers are not required. There is still
a ¾” wide busswork. The concentrator can come within
around 1/16” of an inch due to the better performance in 3-D
magnetic fields. This means the effective length of the
inductor is over 0.99 times ideal. Therefore, it is only
necessary to increase the coil current by 0.4%.
After recalculation, the current in the coil with laminations is
1% lower than the coil with Fluxtrol A. This is a much
smaller difference than 6% from the ideal inductors.
Therefore, a current of 19.9 kA will be used for the coil with
laminations and 20.1 kA for the coil with Fluxtrol A.
Besides current recalculation, it is also necessary to include
losses in the concentrators and heat transfer between the
copper and concentrator. Losses in the concentrator should be
calculated based upon flux density values from simulation of
the part heating and fed into the thermal block as a constant
power source in Flux 2D. The process is continuous, so
simulation should be run until steady state temperatures are
achieved in all components.
Figure 5: Steady state temperature distribution in the
induction coil with laminations- T scale 20 – 250 C
Figure 5 shows the temperature distribution after 2000
seconds for the coil with laminations. At this point, the
inductor is definitely at steady state. The maximum
temperature in the copper is 322º F (161º C). This is in the
corner of the main leg. The temperature in the lower half of
the laminations is nearly identical. This is due to poor heat
transfer between the laminations and the copper due to
uncertain thermal contact and low conductivity filling resins.
The temperature in the return leg is significantly lower due to
Color Shade Results
Quantity : Temperature Deg. Celsius
Time (s.) : 0.002E6 Phase (Deg): 0
Scale / Color
20 / 34.375
34.375 / 48.75
48.75 / 63.125
63.125 / 77.5
77.5 / 91.875
91.875 / 106.25
106.25 / 120.625
120.625 / 135
135 / 149.375
149.375 / 163.75
163.75 / 178.125
178.125 / 192.5
192.5 / 206.875
206.875 / 221.25
221.25 / 235.625
235.625 / 250
Color Shade Results
Quantity : Temperature Deg. Celsius
Time (s.) : 4.4 Phase (Deg): 0
Scale / Color
41.60269 / 113.62705
113.62705 / 185.65143
185.65143 / 257.67578
257.67578 / 329.70013
329.70013 / 401.72449
401.72449 / 473.74884
473.74884 / 545.77325
545.77325 / 617.79761
617.79761 / 689.82196
689.82196 / 761.84631
761.84631 / 833.87067
833.87067 / 905.89502
905.89502 / 977.91937
977.91937 / 1.04994E3
1.04994E3 / 1.12197E3
1.12197E3 / 1.19399E3
Color Shade Results
Quantity : Temperature Deg. Celsius
Time (s.) : 7.599999 Phase (Deg): 0
Scale / Color
44.25912 / 106.98604
106.98604 / 169.71295
169.71295 / 232.43987
232.43987 / 295.16675
295.16675 / 357.89368
357.89368 / 420.62061
420.62061 / 483.34753
483.34753 / 546.0744
546.0744 / 608.80133
608.80133 / 671.52826
671.52826 / 734.25513
734.25513 / 796.98206
796.98206 / 859.70898
859.70898 / 922.43591
922.43591 / 985.16284
985.16284 / 1.04789E3

5. lower losses, higher heat transfer coefficients and shorter
distance to the water cooling.
Figure 6 shows the temperature distribution in the coil with
Fluxtrol A after 2000 seconds. Steady state was definitely
achieved. The maximum temperature in the copper is slightly
lower, 311º F (155º C) compared to the coil with laminations.
The maximum temperature is again in the corner of the main
leg. The overall temperature of the Fluxtrol A is significantly
lower than the laminations due to good thermal contact with
the copper for heat extraction due to exact material dimensions
and the use of high thermal conductivity epoxies (Duralco
4525) [5]. As before, the temperature of the return leg is
significantly lower than those of the main leg.
Figure 6: Steady state temperature distribution in the
induction coil with Fluxtrol A- T scale 20 to 250 C
Based upon these temperatures, it is the opinion of the authors
that both the inductor with laminations or Fluxtrol A would
have good lifetime and no further optimization would be
required.
For the sake of study, let’s consider the case of 50% higher
power in the inductor for both cases. Coil current would be
increased 22% for both cases and the losses in the
concentrator would increase by 50%. Heat transfer
coefficients used will remain the same.
Figure 7 shows the temperature distribution after 2000
seconds for the coil with laminations. The maximum
temperature in the copper is 466º F (241º C). This is in the
corner of the main leg. The temperature in the lower half of
the laminations is nearly identical and there is even a small
spike on the outer corner. The temperature on the return leg is
relatively low.
Figure 7: Steady state temperature distribution in the
induction coil with laminations with 50% higher power- T
scale 20 – 250 C
Figure 8 shows the temperature distribution after 2000
seconds for the coil with Fluxtrol A. The maximum
temperature in the copper is 433º F (223º C). This is in the
corner of the main leg. The temperature in the lower half of
the Fluxtrol A is lower than in the copper, but still elevated.
There is no peak at the outer corner.
Figure 8: Steady state temperature distribution in the
induction coil with Fluxtrol A with 50% higher power- T scale
20 – 250 C
Figure 9 shows the temperature evolution over time for the the
coil with laminations and with Fluxtrol A in two critical areas
of the induction coil, the copper corner and the center of the
bottom of the concentrator pole for a continuous heating
process.
Color Shade Results
Quantity : Temperature Deg. Celsius
Time (s.) : 0.002E6 Phase (Deg): 0
Scale / Color
20 / 34.375
34.375 / 48.75
48.75 / 63.125
63.125 / 77.5
77.5 / 91.875
91.875 / 106.25
106.25 / 120.625
120.625 / 135
135 / 149.375
149.375 / 163.75
163.75 / 178.125
178.125 / 192.5
192.5 / 206.875
206.875 / 221.25
221.25 / 235.625
235.625 / 250
Color Shade Results
Quantity : Temperature Deg. Celsius
Time (s.) : 0.002E6 Phase (Deg): 0
Scale / Color
20 / 34.375
34.375 / 48.75
48.75 / 63.125
63.125 / 77.5
77.5 / 91.875
91.875 / 106.25
106.25 / 120.625
120.625 / 135
135 / 149.375
149.375 / 163.75
163.75 / 178.125
178.125 / 192.5
192.5 / 206.875
206.875 / 221.25
221.25 / 235.625
235.625 / 250
Color Shade Results
Quantity : Temperature Deg. Celsius
Time (s.) : 0.002E6 Phase (Deg): 0
Scale / Color
20 / 34.375
34.375 / 48.75
48.75 / 63.125
63.125 / 77.5
77.5 / 91.875
91.875 / 106.25
106.25 / 120.625
120.625 / 135
135 / 149.375
149.375 / 163.75
163.75 / 178.125
178.125 / 192.5
192.5 / 206.875
206.875 / 221.25
221.25 / 235.625
235.625 / 250

6. Coil Heating Data
0
25
50
75
100
125
150
175
200
225
250
0 200 400 600 800 1000
Time (seconds)
Temperature (C)
Cu Corner,
Fluxtrol A,
HP
Fluxtrol A
bottom, HP
Cu Corner,
Laminations,
HP
Laminations
Bottom, HP
Figure 9: Temperature evlolution in critical areas of the
induction coil with laminations and with Fluxtrol A for
continuous heating
When 50% higher power is used, both of these cases could be
considered to be in the critical temperature range for thermal
degradation. The copper temperature is around 33º F (18º C)
less for the coil with Fluxtrol A. The temperature on the
bottom of the Fluxtrol A is around 74º F (41º C) lower than on
laminations. The difference in overall temperature of the
concentrator is even greater. The lower temperatures present
on the coil with Fluxtrol A should lead to extended coil
lifetime.
The above considerations are for a continuous heating process.
From the above data we can make some short evaluations of
what we could expect during an intermittent heating process,
like single shot hardening. Figure 10 shows the heating
dynamics in the first 10 seconds of heating.
Coil Heating Data
0
25
50
75
100
125
150
175
200
225
250
0 2 4 6 8 10
Time (seconds)
Temperature (C)
Cu Corner,
Fluxtrol A,
HP
Fluxtrol A
bottom, HP
Cu Corner,
Laminations,
HP
Laminations
Bottom, HP
Figure 10: Temperature evlolution in critical areas of the
induction coil with laminations and with Fluxtrol A for
intermittent heating with 10 second on-time
As we could expect, all temperatures are lower than for the
case of steady state. It is interesting to note though, that the
difference between the copper temperatures is much higher
than in the case of a continuous process. The temperature of
the copper in the coil with laminations is 72º F (40º C) higher
than for the coil with Fluxtrol A. This will have a dramatic
effect on the coil lifetime.
The temperature of the concentrators in this case is much
lower than for continuous and can be considered as well in the
safe range. This is only for one heating cycle. Further
analysis will be made in a future paper of the effect of cycling
on the coil temperatures and include running an intermittent
process to steady state.
One thing to note though, is that due to the shape of the curves
it is absolutely clear that the main source of heating in the
concentrator is conduction from the high temperature copper.
Therefore, the concentrator temperatures should be lower than
the high temperature corner of the copper. From this, it can
stated that keeping the coil copper cool will enhance not only
the lifetime of the copper, but the magnetic flux controller
also.
Conclusions
Improving induction coil lifetime in heavy loaded applications
is an area of growing importance. Coil lifetime can be
compromised by mechanical damage, electrical breaks and
thermal degradation. Problems related to mechanical damage
or electrical break can almost always be solved with the
proper maintenance procedures and machine design.
Induction coil failures due to thermal degradation are more
complicated. Reducing the induction coil component
temperatures has a dramatic effect on the coil lifetime in these
cases. Practical methods for increasing the coil lifetime
include raising water pressure and changes to the water circuit
in the inductor. These adjustments are commonly made
empirically based upon the experience of the induction coil
designer.
A more scientific method of predicting induction coil
temperatures is described. The method is based upon an
iterative approach using a combination of computer
simulation, good engineering practices and analytical
calculations. This method can be used for improvement to
existing design or in the process of new induction coil
development.
A case story for continuous seam annealing of heavy walled
tube is presented to demonstrate representative results. In this
case story, the effect of magnetic flux controller type on coil
temperatures is evaluated. Two types of magnetic flux
controller were considered: Fluxtrol A and laminations.
The induction coil with Fluxtrol A had lower overall
temperatures, both in the copper and magnetic flux controller,
than the one with laminations. This should lead to the coil
with Fluxtrol A lasting longer before thermal degradation

7. occurs than the one with laminations in a continuous operation.
A quick evaluation was made for the case of a short heating time with this inductor. For a shorter cycle, the difference between the copper temperatures was much more pronounced. Further study is required for determining the intermittent cycling characteristics to lead to the steady state conditions.
It was clear from the temperature evolution curves that for this case the main source of heating in the concentrators was conduction from the high temperature copper corners. Therefore, lowering the copper temperature will also lower the concentrator temperature.
This paper sets a foundation for the study of induction coil heating and further improves the state of the art in induction coil design process. The method described here may be applied to any induction coil. Future studies on different induction heating applications using this method will lead to further improvements in induction coil lifetime.
References
1. Pfaffman, G.D., (July 2006) Advancements in Induction Heating Tooling Technology, J. Heat Treating Progress
2. Stuehr, W.I. and Lynch, D.C., (Jan-Feb. 2006), How to Improve Inductor Life, J. Heat Treating Progress
3. Haimbaugh, R.E., (2001), Practical Induction Heat Treating, ASM Int.,
4. Rudnev, V.I., Loveless, D.L., Cook, R.L., Black, M.R., (2003), Handbook of Induction Heating, NY, Marcel Dekker
5. Nemkov, V.S., (2006), Resource Guide for Induction Heating, CD-R, Fluxtrol Inc., MI