So how hot was the reactor, really? It must have had the emissivity messed with, since here we are with the settings..
Rossi Lugano/early demo's revisited. (technical)


Reiterating away with the same image frame, we get this:
(The Caps are still far too cool relative to the main body.)
Emissivities set from 0.78 to 0.69 as required by the Lugano Method.

So when were the Rods 235 C and still had Kapton "dots" on them?
The minimum temperature possible in the 2001500 C range is 175 C, so either the dummy Rods were done with another camera (there were 2) or under a different calibration setting.

Initially that didn't work also
If I loaded another profile then the one embedded in the Ravi file, then the colored temperature information disappeared.
Also the configuration menu was not properly working and I had many crashes.
This all without warnings.
Have now installed the software on another, almost identical computer, and it is now working properly.
Could also load the Lugano profile and use it on the MFMP Ravi data
Thanks for the above link

Radiated and convective power calculation of the MFMP Dogbone thermal simulation 500W, 700W and 900W runs.
In the attached spreadsheet the total radiated and convective power of the second MFMP dogbone thermal test was calculated.
Temperatures of the same sections as in the Lugano report where taken from the Ravi files made available by the MFMP. This was done by replacing the embedded profile in the Ravi files by a Lugano style profile.
The dogbone did not show up totally horizental on the Optris thermal pictures and as such the Lugano profile did not match the camera picture perfectly.
I did not correct the profile for this. since the introduced errors where in my opinion not very significant for this comparison.
From each file (500, 700 and 900 Watt) the temperatures for all sections where taken at the 5 minute mark .
Then the average temperature of the ribbed area of each run was taken and this value was used to do new CFD simulations to determine the convective heat correction factor of the ribbed area for that average temperature. This since the fin spacing where the maximum heat transfer occurs is temperature dependent and this optimum spacing increases with increasing temperature.
Otherwise stated, due to the optimal spacing shift to the right with increasing temperatures, the correction factor decreases with increasing temperatures, but is also dependent on the change in the convective heat transfer coefficient.
For the MFMP 500,700 and 900 watt runs the correction factor was found to be relative constant with an average value of about .695
With the above established convective heat correction factor for the ribbed area and the section temperatures the total thermal power (convective and radiated) for each run was calculated
These thermal powers where then compared with the input powers.
The results are :
Run Input power (Watt)Thermal power (Watt) Error (%)
500 Watt502 591.1416.3
700 Watt713916.6727.3
900 Watt8951208.7733.9
The results confirm the earlier statement in a previous post that for the MFMP dogbone thermal test the total convective and radiated power is much more then the applied power.
Update 3112018
For the 500 watt run the error was stated as 6.3%, however it was 16.3 %


Error (%)
Good spreadsheet work based on limited data
By error I thought you meant 'exptal' error
33.9% difference at 900W looks pretty good...
But confirm needs error estimate ,

MFMP dogbone thermal versus applied power. A case study
In post #445 in this forum thread we have shown that for the second MFMP dogbone thermal test the applied electrical power differs largely from the calulated thermal power.
This can be possibly due to one of the following reasons :
Wrong power measurement
Wrong temperature measurement
Emissivity of the casting material is different
Concerning the power measurement we can state that for the first MFMP dogbone thermal test we are able to calculate from the posted data the heater coil resistance for each setting.
This resistance has about the same value for each run and calculating the wire resistance from the supplied heater coil winding data gives about the same value.
Thus it is likely that the power measurments of the MFMP where correct.
Concerning the temperature measurment we know that at least during one of the MFMP thermal tests the temperatures where measured with the Optris thermal camera and the Williamson pyrometer.
Both where in close agreement and as such the measured temperatures with the Optris are supposed to be (near) accurate.
The differences between applied power and calculated thermal power can also be explained if the emissivity of the material used to cast the dogbone is much lower then that for standard alumina.
This possibility will now be investigated in this case study.
Using the data in post #445 a rough estimate was made by which factor the emissivities of the 900 Watt run had to be lowered to bring the applied power in line with the total thermal power.
This factor was about .7 . This means that emissivities for the high temperatures would have been closer to .30 then to the .45 for the ribbed area.
The question now is which ceramic material can have such low emissivity values. A likely candidate is Magnesia (Magnesium Oxide, MgO).
Total emissivity of magnesia as a function of temperaure is given in the following figure which was taken from the
"HANDBOOK OF THE INFRARED OPTICAL PROPERTIES OF AL2O3, CARBON, MGO, AND ZrO2, VOLUME 1" by Milo E. Whitson, Jr. (1975)
For magnesia the handbook gives the following figure :
The dots in the figure are representing measured emissivities by several sources. As can be seen there is a wide spread between those measured values.
The curve represents the curve fit for all these measurements. The curve is in the document also presented as a table and we take that data as a starting point .
We now do a linear transformation on the published emissivity values of magnesia such that for the MFMP dogbone thermal tests for 500, 700 and 900 Watt the applied power is about in agreement with the thermal power. The translation used was :
ε' = 1.02 x ε  0.04
The part of the original curve between 200K and 1800K together with the translated curve is given in the following figure :
The new curve has almost the same shape as the original curve but is somewhat lower but falls largely within the upper and lower limits shown by the measurements dots in the first figure.
We now apply the translated emissivity values of magnesia to our calculations of the thermal powers of the MFMP 500, 700 and 900 Watt runs and determine the difference between applied powers and thermal powers.
The table below shows the differences between the calculated applied power and thermal power for both the original calculation with alumina and for the translated magnesia curve
RunDifference alumina (%)Difference magnesia (%)
500 Watt16.30.94
700 Watt27.30.09
900 Watt33.90.95
Conclusion of this case study is that the large differences between the applied and calculated thermal powers of the MFMP dogbone thermal test almost disappear if the MFMP dogbone was casted from a lower emssivity ceramic then alumina. Especially magnesia is a good candidate to bring applied and calculated thermal powers in line.
Note that if instead of alumina magnesia or an other low emissivity ceramic was used it would also result in the higher surface temperatures measured during the MFMP test.
Update for typo error : Constant factor in emissivity translation is .04 instead of .03
Update 08112018 : Added Excel sheet with calculation

MFMP dogbone thermal test, earlier FEM simulation result
In support of the finding that the MFMP dogbone thermal power is not in agreement with the applied power, I post here the result of an early FEM simulation (Was I think posted on ECW before I moved to LF) that was in my opinion accurate enough to be indicative for the relationship between temperature and power.
Below is shown the FEM model of the dogbone created.
With that model the temperature versus power curve was calculated and combined in a single graph with the results of the MFMP dogbone thermal test.
The graph is shown below
From the graph it can be seen that the temperatures measured by the MFMP are much higher then those calculated by the FEM simulation program.
For the about 480 watt power of the Lugano dummy run the temperature is about 500 degree C while the MFMP measured about 700 degree C.
Note that the 500 degree C is not too far of from the about 450 degree C measured in the Lugano test.
The conclusion of the simulation is that the temperatures measured by the MFMP are not representative for the Lugano ECAT.
I will now starting work on upgrading the FEM model with the latest findings for view factor and convective heat correction and taking into account the power dissipated into the rods.
I also intend to increase the number of different sections the model is made up from in order to increase the accuracy of the simulations.
After doing so I expect that the simulated temperatures will be even closer to the ones in the Lugano report.

Lugano dummy run thermal FEM simulations  FEM model adaptionsHeater coil
Preparing for new thermal FEM simulation of the Lugano dummy run i did some some preliminary FEM simulations and found the following issues :
1. Heat transfer into the rods
Measuring the total convective and radiated heat transfer of the sides of the caps from some inital FEM simulation runs show that the sides of the caps can not supply enough power into the rods by convection and radiation alone.
Temperature of the end cap at the side will be about the same as the measured end cap temperature closest to the end, or about 320 degree C.
Area of the side of the end cap is pi * 2 x 2 = 12.57 cm^2 or 12.57E4 m^2
Convective heat transfer coefficient for the vertical wall is 21.28 and the emissivity is 0.790
This gives 15.25 Watt total thermal power for one side, or 30.50 watt total.
From the dummy run recalculation we found a total thermal power of the rods being 118.38 Watt.
This is much more then the convective and radiated heat from the sides can supply.
The conclusion then is that in order to get enough power transferred into the rods, the rods must have been in good physical contact with, maybe even cemented, to the end caps during the Lugano test. This because radiation and convection can not alone supply enough heat into the rods.
2. Heat transfer from heating element to the side of the end cap
Other preliminary thermal FEM simulations showed that if the heating element is only present under the ribbed area, then not enough power can be transferred from the heating element to the outher side of the end cap to supply (by thermal conduction) the power to the rods. This because the thermal conductivity of alumina is too low resulting in a too high thermal resistance between heating element and side of the cap. This limits the maximum possible heat transfer.
The conclusion therefore is that the heating element must have continued in the end caps till (almost) the end of the caps.
In that case the thermal resistance between heating element and side of the end cap becomes much lower.
New FEM simulations showed that if the heating element is extended to the full length of the ECAT, then enough thermal power can be conducted into the rods.
Conclusion is that the heating coils of the Lugano ECAT must have covered the (almost) full length of the ECAT.
The adapted FEM model wil thus have its heating element running over the full length of the ECAT.
Besides the above issues i have done and still am doing quite a lot of FEM simulations in preparation of a new Lugano dummy run simulation, since the physical configuration of the dogbone (core size, depth of the heating element, material specifications ) all have influence on the final temperature profile and I want to make sure that I understand the influence of all.
(It seems that it has not much effect on the average temperatures).
So it still may take some time before I am ready for an FEM rerun.

LDM ,
About 5 cm leads of each of three twisted heater coil wires exit the Caps at both ends of the “reactor”. About 1 cm each is used at the clamp connection to the C2 cables. This leaves about 4 cm each, or 12 cm combined of Kanthal resistance wire inserted into the rods, at each end, heating the Rod ends.
About 30% of the total heater coil twisted wiring is contained in the Caps and the extensions into the Rods.
Therefore, roughly 30% of the total input power is fed to the combined Caps and Rods, or ~ 15 % into each end (but not 100% of that actually goes to the Rods due to the loose fit of the Rods to Caps).
The glowing twisted wires are quite visible in the Lugano photos. The caption of one photo suggests that the twisted leads are glowing due to conduction from the reactor, which is silly. The wires are glowing orange hot because up to 50 A are passing through them.


The glowing twisted wires are quite visible in the Lugano photos. The caption of one photo suggests that the twisted leads are glowing due to conduction from the reactor, which is silly. The wires are glowing orange hot because up to 50 A are passing through them.
Correct. The heater wires all produce the same amount of heat per cm from wherever they are connected to the positive rail and back to where they are connected to the negative rail.

Correct. The heater wires all produce the same amount of heat per cm from wherever they are connected to the positive rail and back to where they are connected to the negative rail.
If the wire has no temperature dependent resistance and the diameter is uniform, what could change after excessive (over) heating.

About 5 cm leads of each of three twisted heater coil wires exit the Caps at both ends of the “reactor”. About 1 cm each is used at the clamp connection to the C2 cables. This leaves about 4 cm each, or 12 cm combined of Kanthal resistance wire inserted into the rods, at each end, heating the Rod ends.
Correct
So for both sides the total length of heating wires in the rods totals 2 x 12 = 24 cmAbout 30% of the total heater coil twisted wiring is contained in the Caps and the extensions into the Rods.
We consider two situations
1. Heating coil extends under the end caps
Since the length of the caps totals 80 mm and the total length of the ECAT is 280 mm
80/280 = .285 or close to 30% of the total power is dissipated in the end caps
This amounts to .3 * 480 = 144 watts
Again that is if the heating coils continue within the end caps
To take also into account the length of the heating wires in the rods then we need to calculate with heater wire lengths instead of the section lengths above.
If we assume the heater coil spacing being equal to the rib spacing (69 ribs) and the coil diameter is assumed to be approximate 2 cm, then the total heater wire length under the ribs is 438 cm and the total heater wire length under the end caps 175 cm
Total heater wire length is then 24 + 175 + 438 = 637 cm
Amount of power in the rods is (24/637)x480 = 18 Watt
Total power in the end caps is (175/637)x 480 = 132 Watt
Total power in end caps and rods = 18 + 132 = 150 Watt
2. Heating coil ends at the end caps
In this case there is 2 x 3 x 4 cm = 24 cm of heater wire under the end caps
24 cm is in the rods
438 cm is under the ribs
Total heater wire length in this case is 24 + 24 + 438 = 486 cm
Amount of power in the rods is (24/486)x480 = 23.7 Watt
Total power in the end caps is (24/486)x 480 = 23.7 Watt
Total power in end caps and rods = 23.7 + 23.7 = 47.4 watt Watt
The 47.4 Watt in the rods and the end caps in this case is not enough to supply about 110 Watts measured in the rods
So despite what is shown in figure 2, the heating coils must have continued under the end caps if we have 69 windings of heater coil under the ribs.
Therefore, roughly 30% of the total input power is fed to the combined Caps and Rods, or ~ 15 % into each end, or, 7% into the end of each Rod bundle (but not 100% of that actually goes to the Rods due to the loose fit of the Rods to Caps.
Case 1 (heating coils extend under the end caps)
We have 100 x (150/480) = 31.25 % in rods and caps
Of which 100 x (18/480) = 3.75 % in the rods
And 100 x (132/480) = 27.5 % in the caps
Case 2 (Heating coil ends at the end caps)
We have 100 x (47.4/480) = 9.88 % in rods and caps
Of which 100 x (23.7/480) = 4.94 % in the rods
And 100 x (23.7/480) = 4.94 % in the caps
The glowing twisted wires are quite visible in the Lugano photos. The caption of one photo suggests that the twisted leads are glowing due to conduction from the reactor, which is silly.
Indeed
The wires are glowing orange hot because up to 50 A are passing through them.
Note that the above calculations and the simulations I did are when the coil under the ribs has 69 windings. (Same as the number of ribs)
However figure 2 in your post (Was that from the patent application ?) shows 3 x 9 = 27 windings of the heater coil and that can make a large difference for the calculations.
Wonder if figure 2 is a concept drawing or that the details are also correct
Maybe someone can tell me more from photographs ?
I will anyway also analyze the case with 27 windings and do also a FEM simulation for that situation and see if in that case enough power can be get into the rods. (May take some days)
Thanks for your input ! (And thanks for the many edits so that i was able to refer to figure 2 )

LDM ,
When I worked out the specifications for the windings, using steamengine.org calculator, I ended up with 30 wraps in the reactor main body area, 10 wraps for each of the three windings. This in turn gives 7 mm spacing between each adjacent winding (21 mm spacing for each coil winding).
The number of twists unit length for the heater windings is important to getting the resistance and overal length of each winding into the correct range (fit the reactor space and have the correct amount sticking out the ends, using the correct Kanthal wire size).
(This from memory, but I posted the steamengine page with filledin parameters in here quite some time ago, and made such a winding to verify the calculator [which works very well indeed]. I will have to find the steamengine calculations to see if the spacing included one of the windings’ diameter within the spacing, or not).
Edit: The steam engine website had some upgrades, so old bookmarks don’t seem to insert the values into the boxes correctly, and new bookmarks aren’t saving the specifications.
The calculator credits the leads with 23.8 % of the total Joule heat/length of the windings. Wrap spacing is 21 mm between each coil loop for each coil, so the spacing between adjacent coils is close to 4 mm, with a twisted wire diameter of 2.9 mm.
One coil: (settings for steamengine). Three required.
Kanthal A1
Round,twisted
Twist pitch 2.9 mm
2 strands
AWG 15
Target resistance 0.41 ohms
Inner diameter of coil 10 mm
Leg length total 175 mm
Wrap spacing 21 mm
Results:
Resistance wire length (twisted pair) 734 mm
Wire length 559 mm (coil part)
Wraps 10.15
Helix angle 25.7 degrees
Leg power loss 23.8 %
Heat capacity 10059.67 mJ/K
Heat flux 200 mW/mm^{2}
Stable output: 1700 W @ 26.4 V

Lugano dummy run thermal FEM simulations  FEM model adaptionsHeater coil update 1
Thanks to the input from Paradigmnoia a preliminary FEM model upgrade will be made to test the effect of a heater coil with 3 x 10 windings.
The test will be to discover if with such a heater coil configuration enough power can be transferred into the rods.
In the calculations below I did not use the steamengine.org calculator as PARA did, but instead used a spreadsheet to calculate the coil parameters necessary for the FEM model update.
A minor difference with my calculation compared to the data PARA supplied is that my heater coil is 200 mm in length while he calculated for a coil length of 210 mm. (Makes it easier for me to update my current FEM model).
I don't think this will result in a major difference for the prelimary FEM simulations.
Also I will use a coil diameter of 1 cm in calculating the powers to be applied to the different sections.
If we have 3 coils of 10 windings each then my spreadsheet calulations are giving the following data for a coil length of 200 mm and coil diameter of 1 cm
Under the ribs each of the three coils has a wire length of 37.24 cm
Total coil wire length under the ribs 3 x 37.24 = 111.73 cm
Total heating wire length under both end caps 24 cm
Total heating wire lengths in both sets of rods 24 cm
Total heating wire length is then 24 + 24 + 111.73 = 159.73 cm
Power dissipated under the ribs (111.73/159.73) x 480 = 335.76 Watt
Power dissipated under the end caps (24/159.73) x 480 = 72.12 Watt
Power dissipated in the rods (24/159.73) x 480 = 72.12 Watt
Total power in end caps and rods is 72.12 + 72.12 = 144.24 Watt
The total electrical power dissipated in the ECAT is 335.75 + 72.12 = 407.88 Watt
Electrical power dissipated in the rods is 72.12 Watt
Total power dissipated by the rods was calculated in the dummy run recalc as 118.38 watt
So 118.38  72.12 = 46.26 Watt into the rods has to come from the ECAT as thermal power.
(46.26 / 2 = 23.13 Watt for each end cap)
This seems already quite feasible and possible much easier to accomplish then my old calculation with 69 turns and 2 cm coil diameter.
So I now have to work on an update of the FEM model and then do some preliminary simulations with the above data to see where we stand.


LDM ,
Using steamengine, the wire twist rate, lead length, coil spacing etc, could be finetuned to get a 20 cm coil exactly... 🙂
I had to do that last one from memory, in a few minutes, and didn’t spend much time dialling in the coil size.
BTW, the Rods I found to be a nearly intractable problem (especially with no FEM). There is poor to incomplete information on them, and far too many opinion answers to questions than data answers to questions associated with them for my liking.
I am still interested to see if it looks like the 2/3 factor was applied (or not) to the Rods in the Active Runs in the report (even if the method used was flawed anyway), since the numbers reported and the supporting data supplied are not clear on that point.

Thank you so much for doing this,

Using steamengine, the wire twist rate, lead length, coil spacing etc, could be finetuned to get a 20 cm coil exactly... 🙂
I had to do that last one from memory, in a few minutes, and didn’t spend much time dialling in the coil size.
BTW, the Rods I found to be a nearly intractable problem (especially with no FEM). There is poor to incomplete information on them, and far too many opinion answers to questions than data answers to questions associated with them for my liking.
As far as the dummy run FEM simulation is concerned I do not know how much of the normal convective and radiated heat of the sides of the caps is absorbed by the rods and how much just disspated in the environment. Otherwise stated : We do not know the percentage which is going to the rods,
Also we need to add the possible heat transfer to the rods by thermal conduction for which the value is also not known.
It thus will be some guess work which part of convective and radiated power goes to the rods and how much will be transferred by convection.
However from the many preliminary simulations I already did I have the impression that an introduced error in assigning the proper powers makes no large difference for the total profile. (except maybe somewhat for the outer points).
So I am optimistic that from the simulation results we can gain some additional knowledge of the dummy run temperature profile.
I am still interested to see if it looks like the 2/3 factor was applied (or not) to the Rods in the Active Runs in the report (even if the method used was flawed anyway), since the numbers reported and the supporting data supplied are not clear on that point.
I wonder if we can ever make any conclusions on that since the data of the active runs is much less detailed then that of the dummy run.
But lets first finish the dummy run and then determine if the FEM model can be used to gain some additional knowledge of the active runs.
If you, or any other interested person has any ideas on what kind of simulation tests we could do for the active runs then let me know and I will see what I can do.
I am progressing with an update of the FEM model. Am also bringing the dimensions of the model more in line with the drawing in the IH patent application.
(But have seen from the many preliminary simulations I already did that internal dimensional changes have marginal effects as long as the outher dimensions and used materials stay the same).
I found that to obtain accurate results it is important to also simulate the temperature dependency of the material properties. (Alumina, Air) instead of using fixed average values. (but average values still do amazing well for a quick evaluation). So I now use a piecewise linear approximation for those temperature dependent physical material properties.
A problem I am having is that Durapot 810 seems to have a higher thermal conductivity then standard alumina and I have no information what the thermal conductivity value and it's temperature dependency is.
So any information on this point would be welcome.

LDM ,
Durapot properties link:
www.cotronics.com/vo/cotr/pdf/801.pdf
Not sure about dependence of thermal conductivity on temperature but the overall curve shape is likely similar to pure alumina, offset by the engineered conductivity improvement. That is a guess however.

LDM ,
Durapot properties link:
www.cotronics.com/vo/cotr/pdf/801.pdf
Not sure about dependence of thermal conductivity on temperature but the overall curve shape is likely similar to pure alumina, offset by the engineered conductivity improvement. That is a guess however.
The datasheet give a thermal conductivity of 15 BTUin/hour.F.Ft^2
Converting that value to SI units gives 2.16 W/mK (Using the engineering toolbox converter)
However different sources are giving a value of about 35 W/mk at room temperature for Al2O3
That's a large difference !
Any clue ?
