Mizuno reports increased excess heat

    Quote from Daniel_G

    1) the Shanetsu insulation used in this paper is rather poor relative to the state of the art insulation (Vacuum insulated panels). This product (https://www.turvac.eu/0/Products/WhatisVIP.aspx) gives extremely low thermal conductivity (3~10x the product Mizuno used) (3,5mW/mK), with just 20 mm thickness, U value less than 0,22 W/(m2K) can be reached. This would allow more precise calibrations at higher temperatures and more capture of the reactor heat in the air flow.


    2) for replicators with sufficient funding a mass flow meter such as this (https://www.sierrainstruments.com/products/oem-probes.html) would help put to bed about 70% of the discussions here. This device provides 1% accuracy in air mass flow and 0.2% in repeatability. This slight modification would leave zero doubt about velocity profiles and turbulent flow, etc. since it measures mass flow not velocity. Also the turndown rate is 1000:1 and dynamic range is 0 to 20,000 SFPM so this device should allow replicators to measure a higher power level at higher temperatures accurately all the way up to 3000W.


    If someone is successful replicating exactly as in the paper, if it were me, I would improve the air mass flow measurements and improve insulation with VIP technology and run the experiment again.


    That panel has a max working temperature of +80C - a bit dangerous. I'd suggest anyone doing this stuff goes for materials that will withstand +150C at least to be safer, especially because better insulation => higher inner temperatures. Although airflow will cool this, the cooling will not be even and there will be hot spots.

    Quote from Daniel_G

    1) the Shanetsu insulation used in this paper is rather poor relative to the state of the art insulation (Vacuum insulated panels). This product (https://www.turvac.eu/0/Products/WhatisVIP.aspx) gives extremely low thermal conductivity (3~10x the product Mizuno used) (3,5mW/mK), with just 20 mm thickness, U value less than 0,22 W/(m2K) can be reached. This would allow more precise calibrations at higher temperatures and more capture of the reactor heat in the air flow.


    As long as the insulation performance does not change, and the losses from the box can be measured, this makes no difference. There is no great advantage to capturing more of the reactor heat in the air flow.

    • Official Post
    Quote from JedRothwell

    As long as the insulation performance does not change, and the losses from the box can be measured, this makes no difference. There is no great advantage to capturing more of the reactor heat in the air flow.


    Not true Jed. The more efficiently the calorimeter captures heat from the reactor and delivers it to the outlet where it is measured the bigger the signal and the lower the 'noise' (lost heat ).

    Quote from Alan Smith

    Not true Jed. The more efficiently the calorimeter captures heat from the reactor and delivers it to the outlet where it is measured the bigger the signal and the lower the 'noise' (lost heat ).


    In addition: lower efficiency means more chance for different conditions between active and control runs to alter efficiency and hence provide false positives. A good failsafe is that differences between control and active cannot normally be larger than calorimeter losses. Of course it may be posisble to show a much tighter bound than this.

    Blow, blow, blow the man down.


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    We need to get very deep into that box.


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    Quote from Alan Smith

    The more efficiently the calorimeter captures heat from the reactor and delivers it to the outlet where it is measured the bigger the signal and the lower the 'noise' (lost heat ).


    Lost heat is not noise. Ambient temperature changes . . . THAT'S noise, by golly. It is a big problem with this calorimeter, as you see in Fig. 7.


    A high recovery rate (not much lost heat) is desirable for various reasons described by McKubre and others, but I do not think this problem should be called "noise." If the recovery rate fluctuates for the same power level, that's a big problem.


    Hemminger and Hohne call this "thermal leakage" and they say it must be measured as a function of time. Calibrated, in other words. They also say "If losses by thermal leakage are known or negligible, the specific heat capacity of the medium can be determined in a flow calorimeter by the release of a known heat flux (reverse calorimetry)." They say "known"; they don't say you have to toss out the results.


    There are lots of problems with flow calorimeters, including low recovery rates, insufficient mixing, and ambient noise.



    . . . Okay H&H say:


    "For an adiabatic operation the escape of heat from the temperature measurement sites to the surroundings must be reduced to a minimum. This can be achieved by the use of tube materials of low thermal conductivity on the one hand, and high reaction and flow rates on the other."


    It don't say operation has to be adiabatic . . . but I guess that is implied.



    H&H conclude that section (8.2.2 Flow Calorimeters) by saying:


    Here is a summary of the difficulties encountered in flow calorimeters:

    1. The thermal leakage depends strongly on sample and apparatus parameters.
    2. There are no well-defined thermodynamic relationships.
    3. The changes of many parameters in the course of the reaction can hardly be determined.
    4. There exist major calibration problems.

    Some time ago here, I reported that the reactor surface temperature is much lower during calibrations than excess heat runs. That was a mistake. There are two columns in the spreadsheet:


    G Control reactor surface temperature

    K Reactor surface temperature [Ni-mesh reactor]


    I got them mixed up. Both temperatures rise during calibrations and excess heat tests. Column G goes much higher during calibrations, K goes higher in excess heat tests. During excess heat tests, the Control reactor surface temperature is close to the outlet air temperature.


    I will investigate this and report on it in more detail, probably in two weeks.

    Quote from THHuxleynew

    YOU WILL NOT PUT ERROR BOUNDS ON THE DATA IN THE PAPER.


    THHnew has not even included the effect of area. why not?

    and has not shown the edge annulus velocity profile

    Even putting the edge area velocity unrealistically as completely zero gives an error of 17%

    Does THHnew think that the velocity in the entire 3mm edge annulus is zero?


    Rather than get into the nitty-gritty of the fleas on a dog's rear end

    its better to first look at the purpose of the dog

    Is it a police Alsatian a poodle or the next dish at a Shanghai restaurant?


    The focus of the calorimetry is output/input.

    Calorimetry measures mass flowrate m and temp difference delta T

    A rough guide to output/ input is m x Cp x delta T

    Lets say for the calibration/active the delta Ts are 3/3.5


    Output = m x Cp x 3.5

    Input= m x Cp x 3


    Lets assume the error in m is 17% ..The systematic error due to

    the edge annulus overestimate of velocity is 17%... unrealistically


    Dividing Output/Input gives m x (1+17%) Cp x 3.5 divided by m x (1+17%)Cp x 3

    Equals 3.5/3 ....COP = 1.17


    The effect of a 17 % systematic error on the O/I is negligible.

    COP equals 1.17 regardless.


    Rather then a decontextualized myopic 24% pls look at the whole dog.


    For further myopic investigation THHnew might look at the effect on COP

    of an increased air density/ calorific value due to increasing the temperature by 0.5 C.

    Then one can get some minute idea of the error bounds in the COP estimate

    due to THHnew's purported 24% error


    So: RB - let us stick to the point.


    Your contribution above seems to be saying that a 17% error in the airflow does not matter.


    Always good, in technical work, to derive results from the data, rather than work out first what results you want. That way you stay honest.


    What are your error bounds for the airflow figures, stating which errors you consider?


    No error bounds => you have no confidence in those figures. At least I've got some confidence!


    RB - this is like talking to a wall.


    Working out error bounds on data in an experiment is pretty basic. Either you can do it, or you can't. If you can't, then the data has no integrity, and cannot be used to derive anything.


    Now of course in this case error bounds exist - you are just unhappy at my trying to work out what they might be.


    Well - if you don't like mine, propose your own, with reasons. As I have.


    To answer your questions: I don't know what is the velocity profile. Nor do you. That is why you cannot give bounds.


    I've stated why the traverse results cannot in this case be trusted, and why they indicate - if taken at face value - a non-standard airflow where the air near the edge has higher velocity than expected for the air speed and pipe diameter.


    But - if you do trust them, please address these issues and state what error bounds you put on the traverse figures, what bounds you put on the "edge" part of the velocity distribution, and therefore what bounds you put on the airflow figures?


    For example:

    What size is the hot wire anenometer? For the 3cm measurement (3mm from side of tube) how is the (non-point) wire distributed radially?

    What is the disturbance on airflow at the deg caused by inserting the anenometer? Does this increase or decrease the measured velocity?


    I maintain that till these questions are answered you cannot know that the airflow here is non-standard and somehow faster near the edges than turbulent flow of this reynolds number normally is.

    Quote from RobertBryant

    state your velocity profile in the edge annulus

    Is it zero over the entire 3mm

    Yes or No?


    That is impossible, of course. The velocity cannot drop abruptly to zero. Somewhere very close to the edge, it is zero. That's basic physics. But the non-moving layer might be a few atoms thick, for all I know. I am sure it is a tiny fraction of a millimeter.


    You can throw in zero for the whole 3 mm zone. That would be a "worst case" analysis. But it is physically impossible. So, 17% is a gigantic overestimate.


    RB - this is like talking to a wall.


    Working out error bounds on data in an experiment is pretty basic. Either you can do it, or you can't. If you can't, then the data has no integrity, and cannot be used to derive anything.


    Now of course in this case error bounds exist - you are just unhappy at my trying to work out what they might be.


    Well - if you don't like mine, propose your own, with reasons. As I have.


    To answer your questions: I don't know what is the velocity profile. Nor do you. That is why you cannot give bounds.


    I've stated why the traverse results cannot in this case be trusted, and why they indicate - if taken at face value - a non-standard airflow where the air near the edge has higher velocity than expected for the air speed and pipe diameter.


    But - if you do trust them, please address these issues and state what error bounds you put on the traverse figures, what bounds you put on the "edge" part of the velocity distribution, and therefore what bounds you put on the airflow figures?


    For example:

    What size is the hot wire anenometer? For the 3cm measurement (3mm from side of tube) how is the (non-point) wire distributed radially?

    What is the disturbance on airflow at the deg caused by inserting the anenometer? Does this increase or decrease the measured velocity?


    I maintain that till these questions are answered you cannot know that the airflow here is non-standard and somehow faster near the edges than turbulent flow of this reynolds number normally is.

    Quote from JedRothwell

    That is impossible, of course. The velocity cannot drop abruptly to zero. Somewhere very close to the edge, it is zero. That's basic physics. But the non-moving layer might be a few atoms thick, for all I know. I am sure it is a tiny fraction of a millimeter.


    You can throw in zero for the whole 3 mm zone. That would be a "worst case" analysis. But it is physically impossible. So, 17% is a gigantic overestimate.


    Jed:


    You told me that Mizuno had investigated the theory of turbulent flow in this case - so he must know the equations, and that turbulent flow with this Reynolds number (or indeed any) has speed that decreases progressively towards the edge. In this case the average is around 20% less than the peak value. What you said - that the velocity profile must be flat at this Reynolds number - is just untrue. The Re is ~ 12,000 and at anything like this (say 10,000 - 20,000) flow is fully developed turbulent but such flow does not have a flat velocity profile in a pipe. It is just flatter than laminar flow.


    If you answer precisely the above questions about the anenometer we could start to investigate what the traverse results mean - BUT - you cannot say how inserting the anenometer near the edge of a pipe alters the velocity profile in the pipe. It sort of makes sense that it would make it locally more turbulent and therefore flatter than for real.


    So: why are you so confident in results that do not correspond to normal turbulent flow? Of course it is possible that the fan creates much more turbulence than normal in a pipe, and that the small distance to the pipe is not long enough for a normal flow velocity profile to be attained. I'm not saying i know the theory is correct here. But, you are on sticky ground believing that the anenometer traversal measurements are reliable: you just can't know. Further, you'd need to make sure the distance between blower and anenometer was the same when doing the traverse as when doing the blower calibration, since given a greater distance the pipe flow would become more typical, even if it was unuusally flat at the entry of the pipe. Can you confirm this? Also when was the blower calibration done?


    Quote from THHuxleynew

    To answer your questions: I don't know what is the velocity profile. Nor do you. That is why you cannot give bounds.


    Of course you know what it is. Of course you can give bounds. It is right there in Fig. 4. The velocity is uniform across 60 mm of the 66 mm orifice. The velocity cannot possibly drop abruptly to zero in the last 3 mm. Make a reasonable estimate based on common sense and basic physics. You don't have to know exactly. Just look at the velocity profile for similar round, smooth pipes with turbulent air flowing through them. Or you can do it the rigorous way. Here:


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    This ends with the equation right there on the screen. Go for it!


    Or, see the figures and equations here. If you can't do the maths, I think you can use Fig. 4:


    https://jingweizhu.weebly.com/…bulent_flow_modelling.pdf



    2 Velocity profiles: the inner, outer, and overlap layers


    We have seen in Fig. 3 that there are three regions in turbulent flow near a wall:
    1. Wall layer: Viscous shear dominates.
    2. Outer layer: Turbulent shear dominates.
    3. Overlap layer: Both types of shear are important.

    Let τw be the wall shear stress, and let δ and U represent the thickness and velocity at the edge of
    the outer layer, y = δ. For the wall layer, Prandtl deduced in 1930 that u must be independent of
    the shear–layer thickness . . .



    You want bounds? The error has to be far less than 17%.

    Quote from THHuxleynew

    I maintain that till these questions are answered you cannot know that the airflow here is non-standard

    You can predict the average velocity flow in the annulus region based on standard physics with good accuracy.


    In practice we used in our engineering department in a huge dirty and sprawling pulp and paper plant in the 1980's


    a series of nomographs..


    On occasion we would have to derive stuff from equations


    it was horses for courses ...sometimes you needed to be 5% accurate sometime 10% with a simple triangular approximation


    Its not rocket science


    although it may appear so to THHnew for whom fluid mechanics is weirdly novel

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