Embed Size (px)
Citation preview
INFORMATION TO USERS
This manuscript has been reproduced trom the microfilm master. UMI films
the text direetly trom the original or copy submitted. Thus. some thesis and
dissertation copies are in typewriter face, while others may be trom any type of
computer printer.
The quality of this reproduction is dependent upon the quality of the
copy submitted. Broken or indistinct print, colored or poor quality illustrations
and photographs, print bleedthrough, substandard marglns, and improper
alignment can adversely affect reproduction.
ln the unlikely event that the author did not send UMI a complete manuscript
and there are missing pages, these will be noted. Also, if unauthorized
copyright material had to be removed, a note will indicate the deletion.
Oversize materials (e.g., maps, drawings. charts) are reproduced by
sectioning the original, beginning at the upper left·hand corner and continuing
tram left ta right in equal sections with small overlaps.
ProQuest Information and Leaming300 North Zeeb Raad. Ann Arbor. MI 48106-1346 USA
800-521-0600
®
NOTE TO USERS
Page(s) not included in the original manuscriptare unavailable from the author or university. The
manuscript was microfilmed as received.
6
This is reproduction is the best copy available
•
•
•
Reat Transfer Augmentation by ~Iodification of Impingement
Flow Region for Planar and Non-planar Target Surfaces
By
Zhi Yuan Bi
Department of Chemical Engineering
~1cGill University
Montreal, Quebec, Canada
~..
~::.' '.;" .." :.... . -
~.~;~.
A thesis submitted ta the Faculty of Graduate Studies and Research
in partial fulfillment of the requirements for the degree of
Master of Engineering
© Zhi Yuan Bi, June, 2001
1+1 National Libraryof Canada
Acquisitions andBibliographie Services
395 Weling10n StreetOttawa ON K1A 0N4Canada
BibliothèQue nationaledu Canada
Acquisitions etservices bibliographiQues
395, rue Wellington0IIawa ON K1A 0N4canada
The author bas granted a nonexclusive licence allowing theNational Library of Canada toreproduce, loan, distribute or selIcopies of tbis thesis in microform,paper or electronic formats.
The author retains ownership of thecopyright in this thesis. Neither thethesis nar substantial extracts frOID itmay he printed or otherwisereproduced without the author'spenmSSlon.
L'auteur a accordé une licence nonexclusive permettant à laBibliothèque nationale du Canada dereproduire, prêter, distribuer ouvendre des copies de cette thèse sousla fonne de microfiche/film, dereproduction sur papier ou sur formatélectronique.
L'auteur conseIVe la propriété dudroit d'auteur qui protège cette thèse.Ni la thèse ni des extraits substantielsde celle-ci ne doivent être imprimésou autrement reproduits sans sonautorisation.
0-612-75263-1
Canada
• ABSTRACT
Heat transfer enhancement under a single semi-confined turbulent slot jet for
planar and non-planar surfaces was studied experimentally by modification of How
rei!ion using several configurations of nozzles and turbulence generators (turbulators).- - - ......
The results show both positive and negative effects of heat transfer compared with those
for a smooth slot nozzle with the same equivalent width (and length) under the same
experimental conditions.
The experimental results show that the heat transfer rate for the narrower nozzle
of 5 mm width is much higher than that for the wider nozzle of 7.5 mm width at same jet
Reynolds number. In most cases. 50~/O higher heat transfer coefficient cao be obsef',ied
tûr the 5 mm nozzle than that for the ï 5 mm nozzle. while the difference ben,veen the
computed ~usselt numbers at the same Reynolds number and nozzle-to-surtàce spacing
is under 30~·o. For planar surfaces the insertion of turbulence generators in the jet now
was also examined ta see their effect on the impingement heat transfer. It was found that
• the enhancement of heat transfer induced by insertion of turbulators depends on the type
and location of the turbulator employed.
Three rough-edged nozzles. i. e. right-triangular. equilateral-triangular and
rectangular nozzles. were tested tûr bath planar and non-planar impingement surtàces.
For a planar surface. the 1'vo triangular nozzles \vere found ta be capable of enhancing
the impingement heat transfer by up ta 40~/o: the right-triangular nozzle results in more
pronounced enhancement effects. while the rectangular nozzle deteriorates the heat
transfer rate by up ta 10%. The phenomenon is believed to be caused by the differences
in turbulence intensities of the jets issuing from the raugh nozzle and three-dimensional
effects. Interestingly, for non-planar surfaces aIl three rough nozzles generally produced
negative effects on the non-planar impingement heat transfer by up to 50~/0 .
•A
•
•
•
RÉSUMÉ
L'optimisation du transfert thermique sous un seul injecteur turbulent semI
confiné de fente pour surfaces planaires et non-planaires a été étudié expérimentalement
par modification de la région d'écoulement en utilisant plusieurs configurations de becs et
de générateurs de turbulence (turbulateurs). Les résultats démontrent à la fois des effets
positifs et négatifs sur le transfert de chaleur comparés à ceux obtenus avec un bec avec
fente lisse de même largeur équivalente (et longueur) et sous les mêmes conditions
expérimentales.
Les résultats expérimentaux prouvent que le taux de transfen thermique pour le
bec plus étroit de largeur de 5 millimètres est beaucoup plus important que celui pour le
bec plus large de largeur de 7,5 millimètres sous le même nombre de Reynolds de
gicleur. Dans la plupart des cas, on peut obser'ier que le coefficient de transfen de
chaleur est 50 '% plus élevé pour le bec de 5 millimètres que celui pour le bec de 7,5
millimètres, alors que la différence entre les nombres de Nusselt calculés au même
nombre de Reynolds et interligne de bec-à-surface est inférieure à 30~/o. Pour les surfaces
planaires, la mise en place de générateurs de turbulence dans l'écoulement de gicleur a
également été examiné pour voir leur effet sur le transfert thermique. On a constaté que la
hausse du transfert thermique provoqué par la mise en place des turbulateux dépend du
type et de l'emplacement du turbulateur utilisé.
Trois becs rugueux. c.-à-d.des becs à triangle droit. triangle équilatéral et
rectangulaire. ont été testés pour des surfaces d'impact planaires et non-planaires. Pour
une surface planaire, les deux becs triangulaires se sont avérés capables d'augmenter le
transfert thermique par jusqu'à 40%~ le bec à triangle droit a des effets plus prononcés de
perfectionnement, alors que le bec rectangulaire détériore le taux de transfert thermique
par jusqu'à 1D%. Ce phénomène pourrait être attribuable aux différences dans les
intensités de turbulence émises par les gicleurs à bec rugueux et à des effets
tridimensionnels. Il est intéressant de constater que dans le cas de surfaces non-planaires,
les trois becs rugueux ont engendré des effets négatifs sur le transfert thermique non
planaire allant jusqu'à 50%.
B
•
•
•
Lt\CKNOWLEDGE~IENTS
l would like to express my sincere gratitude to Prof Arun S. Ylujumdar. my thesis
supervisar. for his conscientiaus and constructive advice and partial financial support
through this work .
Nlany thanks to A. Siripon. for the construction of the main part of the established
equipment used in this work and his kind help in perforrning and tàmiliarizing with the
experimental procedure.
l wouid also like to thank D. Sakamon. for his useful suggestions and discussions
at various stages of my studies since [ came to NlcGill. and most importantly for his
friendship.
~lany thanks ta Pumima ~[ujumdar in praafreading and revising the thesis draft.
and to Sonar Shah as weil for her help
Thanks ta Darshan Prabhu for assisting sorne of collected data. and ta Shi Yuling
for the simulation data used in this thesis .
Thanks aisa ta the members of the Chemical Engineering ;\t[achine Shop for their
heip in the tàbricatian of the experimental equipment. especial1y to \-Valter. Gagnan and
Frank.
Gratitude ta the members of the Chemical Engineering graduate tàcultv and
department staff \vho helped me in many ways .
c
•
•
•
TABLE OF CO~lENTS
ABSTR.:\CT _ _ -\
RÉSLrwlÉ .. __ _ _ _ ._. __ _. B
ACKl'iO'NLEDGMENTS _._ _ _ _ C
T.AJ3LE OF CONTENTS _ - 0
LIST OF TABLES ... H
LIST OF FIGlJRES . .. l
NO~NCLAn'RE _ .. ~
CK-\PTER 1 I~TRODUCTIO~ 1
1.1 INlPIl\IGEyŒNT JET FLOW·.................................. .. .. 1
1.2 OBJECTrVES .-\;''IT) SCOPE _ .3
1.3 THESIS OL~LI0iE '., .. ~
CK-\PTER 2 PRIOR \VORK 5
1.1 ~TRODCCTION __ .._.5
2.2 INlPNGEl\Œ?\iT JET FLO\V O!\j ~Oi\i-PL.-\.'\i.-\R SLRFACE .. 6
2.3 THE E~RA..l'\iCEWŒNT OF IMPC'iG~GJET
HEAT TRA'SFER _ _ __ _. __ .. _. Il
2..+ FLO\V STRUCTURE STLJDŒS OF HF .. .. __ " _ 15
2.5 CLOSURE ...1 ï
CIL-\PTER 3 EXPERI~IENT,AL SETUP AND
J\rIETHODOLOG\' 19
3.1 rNTRODUCTION _ 19
3 :2 EXPERIMENTAL APPARA.TUS.... ..19
3.2. 1 Overall Experimental Setup __ .. 19
3.2.2 Impingement Plate and Non-planar Target.................... _ 20
3.2.3 Design ofSlot Nozzles 21
o
• 3.2.3.1 Smooth Siot Nozzles 21
3.2.3.1 Rough Slot Nozzles 21
3.2..+ Turbulators 23
3.3 EXPERIMENTAL TEC~lQl.TES 2-l
3.3.1 Heat Transfer w1easurements........... . 2-l
3.3.2 Pressure Measurements .
3 ~ EXPERIMENTAL PROCEDURE :\.i"JTI PL.~"i .
3.5 DATA REDCCTION .
3.5.1 Analysis ofHeat Balance on Studied Surtàce .
35.2 Local Reat Transfer Coefficient .
3.5.3 Average Heat Transfer Coefficient and ~usselt ~umber ..
3.6 DATA REPRODCCIBILITY.................. . . .
.26
.. .. 26
. 28
... 28
29
. 3 l
. 31
•CH.-\PTER 4 HEAT TlL-\NSFER CfL-\RACTERISTICS-
PL -\N..-\R SliRF.-\CE 3..
-l.I ~TRODUCTION .
~.2 HEAT TR.A~'\iSFER ON THE PLA.."\i.-\R SL"RFACE
3-l
.... 34
•
-l.l.I Temperature Distribution along the Surtàce... .. 34
4.2.2 Reat Transfer Coefficients ofTwo Smooth ~ozzles..... 35
-+.2.2.1 Smooth N'ozzle ofW=5mm............................. . , 36
-+.2.2.2 Smooth 0Iozzle with \V=7.5mm 38
4.2.3 Nllsselt Number Distribution __ " -+0
4.1.4 Comparison of the Two Smooth Nozzles........ . 43
~.3 EFFECTS OF NOZZLE-TO-PLATE SPACING.............................. .. 45
4.4 EFFECT OF REYNOLDS NUMBER.............................................." 46
4.5 HEAT TRANSFER AUGMENTATlON SY ROUGR NOZZLES.... .. 48
4.5.1 Right-Triangular Rough Nozzle (Jaws 1) . . -+8
4.5.2 Eqllilateral-Triangular Rough Nozzle (laws 2) .,51
4.5.3 Rectangular Rough Nozzle (Jaws 3) " 54
4.6 EFFECTS OF fNSERTINGTURBULATORS lN JET FLOW " 55
E
• 4.6.1 Single Insertion of Square Rod or Strip 55
4.6.2 Multiple Insertion of Square Rods or Strips.. . 57
4.6.3. Insertion ofa Perforated Plate 60
4.6.3.1 Effect ofNonnal Distance.......................................................... .60
4.6.3.2 Effect of Horizontal Distance 62
4.7 COMPARlSON WITH SIJ.\;flTLATION RESLIl.TS..................... ... 63
4.8 POWERCONSUMPTION fOR NOZZLES................................ 67
4.9 CLOSU"RE 69
CIL\PTER 5 HEL-\T TR.-\NSFER CfL-\R.-\CTERISTICS
:\TON-PLL-\NL-\R SlTRFL-\CE.•..........•.••••.•.•.....•.....••.....•.....70
•5.1 INTRODCCTION .
5.2 THE~\10COtJPLEAR.R....\.~'IGENŒ~l .
5.3. TEMPERA.TIJRE DISTRIBUTION ALONG THE SL"RFA.CES .
54. HEAT TRA,,"ISFER fOR SNfOOTH :\iOZZLES .
541 Heat Transfer Coefficient Distribution __ .
5.4.2 Cornparison between the Smooth ~ozzIes .
5.5 HEAT TR..~'ISFERfOR PLAl"IA.R. A.~l) NON-PLA~'-rAR
SURFA.CES .
56 HEAT TR..~'\[SFER fOR ROUGH NüZZLES __
.. . 70
.70
-.,.. /-
...... 75
.... 75
-...,il
... 78
.... 81
•
5.6.1 Right-Triangular Rough NozzIe (Jaws 2) ... ........ ... ...82
5.6.2 Equilateral-Triangular Rough Nozzle (Jaws 2) __ .. . 83
5 6.3 Rectangular Rough Nozzle (Jaws 3)................................. .. 84
5.7 DATA COl\1PARlSON AND CORRELATION ............................................86
5.7.1Comparison with Previous Work. 86
5.7.2 Correlation 86
5.8 CLOSURE ...87
F
•
•
•
CHAPTER 6 CONCLUSIONS 88
REFERENCES ... [
APPENDLX l LT'N"CERTAINTY .~ALYSIS ", V
APPENDIX 2 ESTIMATION OF HEAT LQSSES \lII
APPENU[X 3 EXPERnvIENTAL DATA ON LOCAL
N1.JSSELT N1.;'MBER XI
G
•
•
•
LIST OF TABLES
2. 1 Summary of sorne work on non-planar surface jet tlow 6
2.2 Summary of sorne work on enhancement of impingement jet heat transfer Il
2.:? Briefsummary offlow field studies ofUF .. , 15
3.1 The dimensions ofthree types aftins ,.. 22
3.2 Geometrie parameters of the nozzles used 22
3.3 Ranges of experimental parameters , 27
4.1 Comparison of experimental conditions of CUITent study and previous work 42
4.2 Correlation parameters of present and previous work ,', , 46
4.3 Correlation parameters for different range ofHIW _ , , 47
5.1 Correlation parameters ofexperimental data of stagnation point
Nusselt numbers for the two smooth nozzles ' 85
H
• LIST OF FIGlfRES
t 1 Typical schematic of confined impingement of a single nozzle.. ". _l
31 Schematic of the overall experimental setup 19
32 Target protrusion and its location on the flat impingement plate .
3.3 Geometrical contigurations of three rough-edged nozzles .
... 20
..21
3.4 Slotjet impingement on an isotlux heated plate 23
3.5 Seetional view of impingement plate with thermocouples and insulation layers ..... 24
3.6 The schematic of thermocouple arrangement and test section _.
37 Schematic ofheat flux description ofthree consective sections of
the heated thermofoil. .
..,,,.... -- .. _.....
..29
3.8 Temperature data reproducibility ......,
.j-
... 0j " Heat transfer coefficient reproducibility .
3.10 Heat transfer coefficients measured at various tluxes ...
35
.35
.... 39
... 36
..36
Temperature distribution on flat surtàce (smooth nozzle) .
Temperature distribution on flat surface (right-triangular nozzle) .
Local heat transfer coefficient distribution (H/\V=2.2. \V=5 mm) .
Local heat transfer coefficient distribution (H/\V=6. \V=5mm)
4.16 Local heat transfer coefficient distribution (HI\V=12, W=7.5mm)
Local heat transfer coefficient distribution (HJ\V=9. \V=5mm) .36
Local heat transfer coefficient distribution (HJW= 18. W=5mm) ... 36
Comparison between local and average heat transfer coefficients (l-I/\V=2. 2) .. 37
Comparison between local and average heat transfer coefficients CHJ\V=6) .... 37
Comparison between local and average heat transfer coefficients (Ht\V=9) 38
4.10 Comparison betvieen local and average heat transfer coefficients (H/\V=18) .... 38
4.11 Local heat transfer coefficient distribution (H/W=1.6, W=7.5mm) ... 38
4.12 Local heat transfer coefficient distribution (H/W=4. \V=75mm) .. 38
4.13 Local heat transfer coefficient distribution (Hf\V=6. W=7.5mm) 39
4.14 Local heat transfer coefficient distribution (HI\V=8. W=7.5mm) . .39
4.15 Local heat transfer coetlicient distribution (HI\V=IO. W=7.5mm) . 39
• -+ 1
~.2
~.3
4,4
4.5
4.6
4.7
4.8
4.9
•
•
•
•
4.17 Local and average Nusselt number (HtW=2.2, W=7.5mm) 40
-+.18 Local and average Nusselt number (HIW=9. W=7.5mm)........................ 40
-+.19 Comparison with previous work .41
4.20 Comparison of experimental data with empinc equation (H/W=12) 42
4.21 Comparison of experimental data with empiric equation (HJW=6) 42
-+.22 Comparison of heat transfer between tv/o smooth nozzles (HJW=6) ... ~3
4.23 Comparison of heat transtèr between two smooth nozzles (HI\V=8 and 9) .. ~3
-+.24 Comparison ofNusselt numbers betv/een two smooth nozzles (HIW=6)... 4-+
-+.25 Comparison of Nusselt numbers betv/een !\NO smooth nozzles (H/\V=8 and 9). .44
-+.26 Effect of nozzle-to-plate spacings on stagnation point heat transtèr coefficients. .+)
-+.27 Effect of nozzle-to-plate spacings on Average ~usselt numbers.... 45
4.28 Comparison between measured and correlated Nusselt number.... .. ~7
4.29 Comparison of local heat transtèr coefficients bet\\-'een smooth and
right-triangular rough nozzles (MV=2.2)........ . -+8
-+.30 Comparison of local heat transfer coefficients between smooth and
right-triangular rough nozzles (H!'N=6)....................................... -+8
-+.31 Comparison of average :-';usselt numbers bet\veen smooth and
right-triangular rough nozzles (ThW=2.2) 49
4.32 Comparison of average Nusselt numbers bet\\-'een smooth and
right-triangular rough nozzles (H/\V=6)....... 49
4.33 Enhancement effect ofright-triangular rough nozzles (HJW=2.2) .. 50
4.34 Enhancement effect of right-triangular rough nozzles (HJW=6) 50
4.35 Comparison of local heat transfer coefficients between smooth and equilateral
rough nozzles (H!W=2.2) 51
4.36 Comparison of local heat transfer coefficients between smooth and equilateral
rough nozzles (H/W=6) 51
4.37 Comparison of average Nusselt numbers between smooth and equilateral
rough nozzles (H/W=2.2) ... 51
4.38 Comparison of average Nusselt numbers between smooth and equilateral
rough nozzles (HJW=6) . .51
4.39 Enhancement effect of equilateral-triangular rough nozzles (H/W=2.2) .,... .52
J
• 4.40 Enhancement effect of equilateral-triangular rough nozzles (HIW=6) .
4.41 Sketches ofnozzlejaw arrangements .
-,........... )-
· 53
4.42 Comparison between shifted and normal nozzles (for Jaws 2).......... . 53
4.43 Comparison beween smooth and rectangular rough nozzles (HIW=1.2) 54
4.44 Comparison beween smooth and rectangular rough nozzles (HJW=6) 54
4.45 Effect of inserting a rad turbulator in jet flow at different Reynolds numbers 55
4.46 Effect of strip turbulator and its inserting locations............................................ .. 56
4.47 Effect ofturbulator types at different locations . S7
4.48 Effect ofturbulator types and inserting locations (Rew = 12.000). """ .. 58
4.49 Effect of Reynolds numbers on heat transfer with rods (l-I/\V=-t) .59
4.50 Effect of insertion object and its locations .
4.51 Schernatic of jet flow affected by the insertion .
4.52 Effects of the plate and its locations .
. ... 60
· bl
· 62
•4.53 Comaprison \vith RS~1 and standard k-E models (Rew =12.000)
4.54 Comaprison \vith RSNI and standard k-E models (Rew =9000) ..
4.55 Comaprison with RSNI and standard k-E models (Rew =6000)
4.56 Comaprison Vv'ith RS~1 and standard k-E models (Re\\=3000)
4.57 Comaprison with RS~l and standard k-E models (Rew = 1500) ....
-+.58 0/ozzle discharge coefficients .
459 Comaprison of purnping power..
.. 64
64
.. aS
.66
68
68
· 73
Temperature difference profile for smooth nozzle (\V=7.5mm)... ..74
Local heat transfer coefficient distribution (HIW=6)....................... . 75
Temperature distribution for smooth nozzle (H;W=8)................ ..... .75
Temperature distribution for smooth nozzle (H/W=10)..................... . 75
Temperature distribution for smooth nozzle (H;W=12).................... ... 75
Local heat transfer coefficient distribution at various Re numbes (H/W=9).. ..76
Temperature distribution for smooth and rough nozzles (W=5 mm) .
51 Schematic of the thermocouple arrangement on a semi-cylinder surface and
the flat plate region " .. ' . ..71
Temperature distribution for smooth nozzle (\V=7.5mm).. . 735.2
5.3
5.4
5.5
5.6
5.7
5.8
• 59
K
.81
•
•
•
5.10 Local heat transfer coefficient distribution at various Re numbes (HI\V= 12) 76
5. Il Local heat transfer coefficient distribution at various Re numbes (HIW= 18) 77
5. 12 Comparison between local and average heat transfer coefficients
at various Re numbers (HIW=9) 77
5.13 Comparison oflocal heat transfer rates bet\veen tv/o smooth nozzles (H/\V=12) 78
5.14 Comparison of average ~usselt number between tviO smooth nozzles (H;W=12) .78
5.15 Comparison of local ~u numbers betvieen planar and non-planar surtàces
(I-IJW=6)............... . .79
5.16 Comparison of local )iu numbers between planar and non-planar surtàces
(HI\V=10)............................. . .79
5 17 Comparison of average Nu numbers bet\Veen planar and non-planar surtàces
(1-{/\\l=6)...................................................................................... 79
5.18 Comparison of average ::\u numbers bet\Veen planar and non-planar surtàces
fTh'\V=10) , "'. ,.,,79
5.19 Comparison of local heat transfer coefficients bet\veen smooth nozzle and
Jaws 1 (E-LW=9) .
520 Comparison of average ~u numbers bet,-veen smooth nozzle
and Jaws l(HJW=9).............. 81
5.21 Comparison of local heat transfer coefficients bet\\ieen smooth nozzle and
Jaws 1 (I-{;W=12) ''' "'" 81
5.22 Comparison of average Nusseit numbers between smooth nozzle and
Jaws 1 (HI\V=12) 81
523 Comparison of average Nusselt numbers between smooth nozzle and
Jaws 1 (HJW=18) ,............. 82
5.24 Comparison of average Nusselt numbers between smooth nozzle and
Jaws 1 (HIW=18) . 82
5.25 Comparison of averôJe Nusseit numbers between smooth nozzle and
Jaws 2 (HJW=9) .., 83
5.26 Comparison of average Nusselt numbers between smooth nozzle and
Jav\iS 2 CH/W=12) 83
L
•
•
•
5.17 Comparison of average Nusselt numbers between smooth nozzle and
Jaws 1 (HIW=18) 8~
5.28 Comparison of average Nusselt numbers between smooth nozzle and
Ja\vs 3 CH/W=9) __ _._... 85
5.29 Comparison of average Nusselt numbers between smooth nozzle and
Jaws 3 (HJW=12) _ _....................................................... ..85
5 30 Comparison of average ~usselt numbers between smooth nozzle and
Ja\VS 3 (HJW=18) _.. _.. ... .. ... ... .... .. ..... ... .. . .. .85
5.31 Comparison with previous work ".. _8S
M
•
•
A
8
C
0"
Os
-h
K
kr
k
L
NOR~IENCLATURE
area of heated strip surface (m::)
area of the nozzle exit. m::
width of divided section ofthermofoil under studv (m)
specifie heat of the impingement plate (Jlkg K)
nozzle discharge coefficient
nozzle diameter (mm)
thickness ofthermofoil or diameter of cylinder target (m)
hydraulic diameter of the nozzle exit. fi
hydraulic diameter of the slor. m
stagnation point heat transtèr coefficient (W/m:: K)
local heat transfer coefficient (W/m:: K)
area-average heat transtèr coefficient (\V/m:: K)
thermal conductivity of impingement plate (\V/mK)
thermal conductivity ofheating thermofoil (\V/mK)
thermal conductivitv of fiber glass (\V/mK)- ...thermal conductivity of air at nozzle exit temperature (\V/m K)
length of divided section ofthermofoil under study (m)
the length of total heating therrnofoil (m)
thickness of fiber glass layer (m)
stagnation point Nusselt number
local Nusselt number
•
lVu area-average Nusselt number
.:1P pressure drop across nozzle (Pa)
Ql.Q:: heat conduction between consecutive sections (W)
Q.: convective heat dissipated from the heated surface (\V)
Qi internal-energy change of the stainless steel sheet and insulation layer (\V)
Qk conductive heat 105S (W)
Qr radiative heat loss (W)
N
• Qr radiative heat loss (W)
Qt total heat generated by the thermofoil heater (W)
qc the local convective heat flux, Qc/A (W/m2)
r radius coordinate with ongin at the jet axis
R electric resistance of the foil (n)
Re dimensionless Reynolds number (W
Tt.,x surface temperature at location x and time t ~C)
Tw.x local wall temperature (oC)
Ti initial temperature of the impingement surface (oC)
T· jet exit temperature ~C)J
Ts surface temperature of the heating steel foil (oC)
T* dimensionless surface temperature
ti respective temperature of studied sections (oC)
Uj relative uncertainty of Xi
V volumetrie flow rate, m3/s
• w nozzle diameter or width (m)
X horizontal distance from the stagnation point (m)
x horizontal distance from the stagnation point (m)
Xi independent variable
Z a calculated quantity in discussion
Greek Syn,hols
a unit pumping power (Pa m2KJW)
E emissivity of the smooth stainless steel foil
y an intermediate parameter defined in Eqn. (l0)
II electric resistivity of metal foil (nm)
v kinematic viscosity of air (m2/s)
P density of the impingement plate (kg/m])
• Pj density of air, kg/m3
o
• ICHAPTER 11INTRODUCTION
1.1 fmpingen1ent Jet Flolv
Impinging jets are v.idely used in various fields of industry to obtain high local heat
and mass transfer rates. Sorne imponant applications include annealing of metal sheets.
tempering of glass. drying of textile and paper products. deicing of aircraft systems. and
cooling of high-temperature gas turbines and high-density micro-electronic chips. The
rypical fluid mechanical features of a submerged single phase jet impinging on a t1at
surtàce are displayed in Figure 1.1 .
• Dor W ConJinemenr hood
\ 1
~\.-"~.... Ir "elOCllV
proJiles ---.
Free
Pvtenrzal core
Stagnation =one orlmpmgemenr =one
1
1
~~~~~~~~~"""""""'~
H
fmprngemenl plate
•Figure 1. Typical schematic of confined impingernent of a single nozzle
•
•
•
As shown in figure 1.1. three mam reglons of the impinging jet flow can be
defined. viz.. the free jet region. the stagnation region and the wall regicn (of lateral or
radial flow outside the stagnation zone). In generaI. the free jet developing trom the
nozzle exit is turbulent and. at the nozzle exit. it is characterized by a uniform velocity
profile called the potential core where the viscous effects are not important. The free jet
region is termed 50 since the tlcw in trus zone is unaffected by the impingernent surtàce
\Vithin the stagnation zone. the tlow is intluenced by the target surtàce and is rapidly
decelerated and accelerated in the nonnal (y) and transverse (r or x) directions.
respectively. With increasing r or x. velocity components parallel to the surtàce increase
from a value of zero at the \vall te sorne maximum value. and subsequently decaying ta
zero agam.
Generally. tViC types of jet configurations are employed. \/1Z.. circular and slot (n'la
dimensional). The advantage of the slat jet over the circular jet is that it can provide a
larger impingemem zone and ensure relatively more uniform jet impingement on the
5urtàce. The jet impinges on a solid surtàce and spreads out along the 5urtàce \vith a high
transverse velocity. thus inducing high mass and heat transfer rates. Jets can also be
classified as submerged if they discharge into an ambient tluid of similar physical
propenies (e.g. air in air), and unsubmerged if the properties of the !\-\/o tluids are different
(e.g. \'later in air).
Confined IJF is differem from the unconfined type by having a hood (planar.
conical or other). which envelopes the nozzle. thus making the jet tlov...· unaffected by
entrainment from surrounding fluid. The first configuration is normally required for
industriai applications where the hot jet air must he collected and recirculated.
~'reasurements for unconfined jets. which are equipment specitic. are more complex.
because the transfer rates are affected by the amount of entrained fluid and bv its
temperature relative to that of the jet and the impingement surtàce.
For practical applications. numerous wcrks have been conducted on the heat
transfer behavior of impinging jets of various configurations. Ta improve the design of
such systems quantitative knowledge of the parameters affecting the heat transfer rate is
required. The heat transfer rate to or tram a jet impinging onto a surface is a complex
2
•
•
•
function of many factors: Reynolds number (Re), Nusselt number (Nu). Prandtl number
(Pr). the dimensionless nozzle-to-plate spacing (HAV or H/D). and the dimensionless
distance from the stagnation point C;(;W or X/D). In addition. the effects of nozzle
geometry. flow confinement. turbulence. recovery factor. fluid properties and surtàce
configurations have ail been shown to be significant. This thesis is concemed with an
experimental study of the enhancement of impingement heat transfer rates under a single
confined slot jet by introducing turbulence. The effect of pedestals mounted on the target
surface is also studied.
1.2 Objectives and Scope
The CUITent proJect fo110\V5 a prevlOUS study by Siripon (2000) and \lvill
experimemally investigate the effect of modification of the impingement tlo\v region on
the augmentation of heat transtèr tàr both planar and non-planar target surtàces (impinged
by smooth and rough-edged slot nozzles).
• Investigate experimentally the effects of weIl defined protruslons or
pedestals on the target surtàce on enhancement of heat transfer for a
confined impinging sial jet and ta compare qualitatively and quantitatively
the difference betVv"een the planar surtàce and non-planar surtàces
• Study the effect of jet turbulence on heat transfer by inserting a turbulence
generator between the jet nozzle and the impingemem surtàce tàr bath
smooth and rough slot nozzles.
The overall experiments are divided into two parts. namely. heat transtèr
measurements and tlow measurements .
3
•
•
•
1.3 Thesis Outline
This thesis includes six chapters staning with Introduction (Chapter 1). Chapter 2
presents a brief literature review. focusing on works conceming non-planar impingement
surface and etfects of turbulence on impingement heat transfer. Chapter 3 describes the
experimental apparatus. experimental techniques and data processing and analysis. In
Chapter 4. heat transfer characteristics of impinging jet flow on a tlat surface are presented
involving jet turbulence effects. ~on-planar surface impingement jet flo\v heat transtèr is
experimentally investigated in Chapter 5. Chapter 6 summarizes conclusions of CUITent
research.
PRIORWORK
2.1 Introduction
A large number of papers have been published on lJF trom as early as 1950' s. The
studies perfonned can be categorized into tv/o groups. in terrns of geometry of the jet
noules used. namely. 2-D slot jets and circular jets. [n these t\\/O groups. both single and
multiple jets have been covered. and the impingement surtàce may be planar. non-planar
(rib-roughened. convex and concave). and inclined (at an angle \-vith jet t10\V) Studies
reponed involve experimentaL analytical and computational investigations. The aim of
these studies is to find various factors affecting heat and mass transfer in IIF in order to
provide helpful foundation for practicaJ application.
Among sorne re'.iiews of IJF. y{artin (1977) published an extensive summary'
focusing on the engineering applications of impinging nozzles. Reasonable empiricaI
correlation equations have been presented for the prediction of heat and mass transtèr
coefficients for a broad and techno!ogicaI imponant range of variables on the basis of
sorne experimental data for single round and slot nozzles and arrays of round and slot
noules. ~lujumdar and Huang (1995) described a major practical application of lIF. i.e
impingement drying. A novel operation in which hot jets are directed normaJly onto thin
beds of pellets transported on a slo\'i-moving conveyor. was extensively examined.
conceming design of impingement dryers and heat transfer correlation.
For the mest parts. previeus studies have been focused on optimizing the transport
processes associated with steady impinging jets on flat surtàces. Trns review. however,
mainly covers tv/a key parts of lIF heat transfer: non-planar surface and enhancement
techniques. plus flow structure studies of IlF.
5
NOTE TO USERS
Page(s) nct included in the original manuscriptare unavailable from the author or university. The
manuscript was microfilmed as received.
6
This is reproduction is the best copy available
• Table 2.1 (continued)
TahakoffandClevenger( 1971)
o
Heat transfer for three.configurations of jet impinging onthe leading edge inner surface of ablade wall (the inside surface of a •haIf-circular cylinder)
Slot nozzle: \'/=0.070 and O.5in.length is 6in
Row of round nozzles: d:;:O .15~in •and the three ro\\"s had spacingsof 0.625. 1.250 and 2.200in
Four square arrays of round jets
The slot jet \\ith the smaller v.idthdemonstrated better average (100%more) heat transfer than the l~lrger slot
The round Jet array configurations tendedto give better overall heat transtèr thanthe ra\\" of round jets (300
/0 more) andslot jet (-15°/0 more) configuratlons
The presence of solid partlcles (sand) mthe impinging air decreased the heattransfer up to 30% by fornung ainsulation film on the model surface
M~tzger.
Saluer Jl1dJenkms(i 971)
•
•
1 Hp;cak
1 (981)1
1
1
i1
i1
1
GauChung
1 (199 i"')
p
and
"-.li
.,.1
( ,\1
2-D siot jets md single lines of •evenly-spaced circular Jetsunpinging on concave surtàce
Slot nozzle: Vi=O.64mmRound nozzle: d:;: 1.52mm
The ratio of centçr-to~enter
spaeing to hole diameter for single •lines of circular jets: 6.67. 3.33.1.67 and 125
Re:;:2000. -1.000 and 6000
Studies on heat transfcr from al.raw of impmgmg Jets to concavecylindrieal surfaces
Semi-eylindrical surface of 25.tmm dia. md of 127mm dia. •~ozzh~ d:;:952. 635 and 3 18mm 1
Hid: 1- 7 1
••
80th the concavc and the eonvex •surface made from a eyIindricalplexiglass tube eut in haIt: whichIS 16em in insidc diameter and16cm in length
Re=6000-35.000 •
Different SIZt: reetangular nozzlesof \vldth 0.35- lem
51 thermoeoupks uscd. Diff~rent
long strips use different OCpower for hcating unifonnly •
7
Large spacmg (l-fj\V) IS accomparued by a
decrease in cooling performance over theenure leading ~dge reglon \\ith larg~r
decreascs in general occumng at smallvalues of II\V n~ar the stagnatIon line.where 1 is half kngth of targct surface arc
Actual impingement of the Jet occurs onthe side wall of the cavity rather than in
the plane of s~mmetry
.-\t .tS' angular distance from stagnatIon 1
poinr. there oecurs a kmd of :l seeondarymaximum of heat transfer that :lppears tobe strongly depcndent on Relntensity of heat transt'er from IJF mayreach levels as hlgh as 10.000 \V/m:K.comparabk with evaporatlve coolmg
Flat surtàce: l'lu - Re<1; -:;'';-
Semi-cvlinder surtàcc: ~Vu - Recj'
For impingement cooling on a conn:xsurface. a series of 3-D eountcrroratingvortices arc mitiated near the wall on thestagnation point. which can enhance thehear transfer proeess (up to 50%
)
For the case on a concave surface.Tay lor- Gortk,. vortiecs are producedinstead. which Carl enhanœ the hcattransfer rates along the surface (up ta30%)
The mcrcase of surtàec curvature cmaecordingly increase the stagnation pomtNu number (up ta 40 %
)
\)
i\
"~liyake. '\)
Hiram andKasagi(1994)
.:'";,"./
Gabour and l'
Lienhard(1994 )
;
Table 2.1 (continued)•
•
•
Gau and Lee(1992)
Brahma.Padhy andPradhan(1994)
Pekdemirand Oavies(1994)
o
o
<)
Equally spaced ribs of 3mm •square elements having differentrib pitches. i.e. 9 and 12mmFour Slzes of slot nozzles withlength 15cm and different \\idths •of 0.35. 0.6. 1.1 and lemRe =1500 ta 1l.000HIW = 2 ta 16
Large-scale transverse repeated- •rib-~-pe roughness made ofsquare sectIon of copper black •attached on a flat surface.2-D nozzle of 50 by SOOHI\\'= 3-8 •Re =2..WO-.57.000
Roughness helghts is -+.7-28.2~. 1 •
dIe distance tS O.2-1.0nun k/d 1
effec! tS consideredRound nozzle of inner diameters4.4.6.0 and 9.Onun •Hld=10.8Re=10.000-84.000
Concave semicylindrical surtàce 1.wiili 100 mm ln diameter and 150mm in lengthSlot nozzJe width: Smm. 10mmand 14mm. Length: ISOmm •SmgIe-row round nozzles \\ithdiameter of 7mm. spacmgs: 1G. 15and 10mmHJ\V: 1.6 to 14.1
Cylinder was a tube of 10cm •0.0. (d). 49.4cm Icngth and ~.5
rn wall thicknessL/0=1-10. jet velocity 1S 4.5-27.5m/s. Slot nozzlc \\idth •=30.5crnRc=7.77xIO.l ta 1.778xl05
jet impact anglc( ~O-90.)
8
The stagnation point Nusselt number lSrelatively low and is significantly affectedby Re. the gap \\idth-to-rib helght ratioand locationThe maxunUffi ln the stagnation porntNusselt number oceurs mostly atapproximately H/\V= 10
Width of stagnation reglOn grows \"lthincrease in HiWHeat transtèr codficient IS generallylarger than that of a smooth surface (up ta
3000/0 in sorne cases!)Flow from the stagnatIon Ime ~hanges
faster from lammar to turbuknt tlow dueto rouclmess
Heat transfer can be slgnificantlyenhanced (up ta 500/0)by li~ presenœ ofroughness è1emenrs which can disrupt thechin thermal boundary layerThe enhancement tncreases \\lthincreasing Re number and deereasing Jetdiameter
Heat transt~r tàr slat nozzIe tncn~ases
\\ith Re but decreases as lie r::ltio of thearcas of the he::lt transfer su rt'ice ta thenozzJe ÎncreascsFor a single ro\\" of round jets the heattranstèr decreascs when lie ratio ofcenter-to-centa distance of the nozzle tothe diarneter of the nozzle mcreases.
The more nearly parallei is the flow tù thecvlinder ax.ls 50 the c1rcumtèœntlallyaveraged Sherwood number for the entIrccylinder becorncs largerLocal Shef\\"ood number cao increase upto 250% when Re number 15 from77.707 to 177.555
Table 1.1 (continued)•
•
•
~Iasood.
Baughn mdYap (1996)
L~~. Chungmd Kim( 1997)
Gt:unyo-ung.\-lansoo mdJoon (1999)
Chung YS ..L~e D.H.J.I1d Lçc 15( 1999)
Mansoo.Han.G~unyoung.
Joan andDong (2000)
o
, ...
)
:"\./
)
)
'l
/\./
',j
"'j
<)
Concave surface D:;::30.48cm •h~ating for 40-6 hrs and suddenlytaken out of a oven and ~xposed •to impinging jetRound nozzle of ·km diameterJet-to-surface distances (L/D=2A •and 6)
Re=23.000
Round nozzles \\ith three mner •diameters of 1.3. 2.15 md 3·kmDimt:nsionless surface curvature:0.03'+ - 0.0809 •Red: 11.000 - 50.000HJd:2-10
Senu-<:ircular concave surtàce •\vith 150 mm inner diarneter andt1at surtàce •Three different slot nozzles:round. rectangular and 2Dcontoured •5920<Re<25..500HI\\! =05-20
Three r:-pes of rib roughness •dements (2mm diameter) anachedon com'ex surfaœ. Different gapsbetween the ribs •Round nozzle of 2. 15 cm diameterRe=23.000. L/d=6-10Transient heat transtèr.rneasurement \\ith TLeS~mi-<:ircular concave surtàce •\\ith 150 mm diameter2-D siot nozzle with 5 mm \\idthReynolds number based on 2times the nozzle width: 1780.2960 and 4740 •Using LDA to mcasurc thedistributions of mean velocity andvelocit\' fluctuation on theconcave surfaceHIW =0.2-14
Increas~d entrJ..inment of warm air fromthe surface can decre~e heat transtèrTItree effects. thmmng of the boundary.reduced thenna1 development. and TaylorGartler vortices can increasc heat transtèrThe rcsults show mat curvature can affectthe local heat transtèr. gena::tllymcreasing the local heat transfer rate (upta 18% L whih~ the toul heat transfacoefficient. does not necessarilv increaseStagnation pornt ~ussdt numbermcreases up ta 2",,% \\ith mcrcasmgsurtàce cur\'atur~
The dependence of the mean ~usselt-
nwnber for larger H/d IS stronger (.V. x
Re) ~ for 6~HJd.:s; 10. and .V... x: R~ ., tor
2~J-U~6
Different nozzles result rn diffèrent t10wmd heat trmsfer charactcnstlcsThe secondary peak of local ~u numberoccurs marc prommently as the ReUlcreasesThe average heat transt~r rates \\crecnhanced (up co 16% tor HJ\V=12)
1
.-\fier the l ~t rib posltlOn ~u numbas arehigher than those on th~ smooth surface(up to 75 %
)
The average ~u number on the nb-jrou1!hened surfacc mcreascs bv maximum 1
14-34% for 3 rib r:"pes. respe~tively 1
Beyond r/d=~-5. the rib roughness doesnot affect the heat transfer any more
The occurrence of secondar:' peaks andtheir locations have becn explained fromthe variation of measured velocitytluctuations of the wall Jets evolvingalong the streamwise directionThe incrcase of stagnation hcat transt~r
rat~ for 2-3 < HiB < 5-6 has bcensucccssfully explaincd due to the stecpincrease of velocir:' fluctuations mcasuredin free and impinging jets
Table 2.1 (continued)
•
•
•
•
•
McDaniel -)and \V~bb
(1000)
,)
)-J
1:
1
Abdlmonem.-"lichel andChandr.1kant(2000)
Azad. )
Huang and 1
Han (2000)
!".1
Heated cylinder made from •oxygen-free copper rad ofdiameter d:::; 1.27cm and length15.1cm
Contoured and sharp-çdgedorifices usmg nozzle \\idths of •0.5. 1.0 and 1.5 cylinderdiameters
R~:nolds number. based on.cvlinder diameter rather thannozz1e \\idth. ranged from 600-SOOO •Cylinder diameter-to-jet \\ldthspacings are 0.66. Lü and 2.0Jet exit-to-nozzle distance are lI 1
The dIect of surtàce roughness •on the average heat transfercharacteristics of an unpmgmg:ur Jet WJ$ experimentallyinvestigated
The roughness took me shape of acircular array of protrusions of •0.5mm base and O.5mm height.1.8mm spacingRe: 9(,00 - 38..500HJd:::; 1 - 10: d:::; 6.S5mm
•Dimpled (rough) surfaces \Vere •made m (wo different panems(23x9:::;207 dimples and llx 5:::;55 •dimples)
Dimple dimention: 0.635 cm mdiameter (equal to the jet •diameter) and 0.3175 cm deep onthe urget surfacesRe :::; 4850 - 18.300
10
The slot jet ~ields considerably rnghaaverage heat transter (1.2-2 rimes) fromthe cylinder \vhen compared to me mfirut~
parallel tlow case on me basts of identlcalslot jet and infinite now average veloelty
\Vhile heat transtèr enhancement ratio forthe eontoured orifice IS reIatlvdyindependent of R~~TIolds nurnbers. theenhancement for the sharp-èdged onfieeincreases markedly \\ith Re~TIolds numberup to 100%
.
-"laximum in average ~usseit nurnber.more evident at higher Re:TIolds numbas.occurs at a nozzle-to....:ylinder spacmgbetween 3-7 nozzle \\ldths and theoptimum spacmg IS dosa ta the nozzleexit for the sharp-èdged onfice than forthe conroured onfie..:
Surface rouglmess disrupts the boundarylayer ~)rld promotes turbulence of me walljet whtch results m ID merease m heattransfer.-\n mcrcase of up to 6 OOn of the J.\erJ.ge~ussdt number due ta surt"ace roughness
The maximum dTect of the roughn~ss onthe average heat tr::msfa charactensttcswas most nOClceabk at R~~ 14.000 to34.000 for Hld::::4 to 8\7 R l)jll~ t' fl •.V:J- e - or at surtace-V R 1) 70S ' h cl •i :l - e - tor fOUs:! ene surtaceThe ~usse1t numbers for a dimpled and asmooth surtàce were about the SJ.IT1t:
The dimpled provldes a higher heattranst~r due ta an incrt:ased surtàce areawhen compared wim a smooth surface
The number of dimpks does have apositive mtluence on the heat transfercoeffiCient enhancement (20°1.. differencebet",een the many-dimples and the Iessdimplcs cases)
• 2.3 The Enhancement ofImpinging Jet Heat Transfer
Thought lJF is widely used for realizing high heat transfer rates between a t1uid
and a surface. the area of enhanced heat transfer is limited to the neighborhood of the
stagnation point. Heat transtèr is very intense near the stagnation zone due to direct
impact from air jet. but deteriorates quickly as the region extends away from there. [n
general (\iusselt numbers monotonically decays from a maximum value at the stagnation
point. This phenomenon is more notable for a stationary surtàce. while a moving surtàce
can effectively reduce the difference. Siripon and Mujumdar (1999) summarized sorne
enhancement techniques for IJF. namely. modification of jet. nozzle. confinement surtàce.
nozzle-to-place and impingement surface. Table 2.2 lists sorne of previous \vork on heat
transfer enhancement of lJF.
Table 2.2 Summary of sorne work on enhancement of impingement jet heat transtèr
;~ inch upstream from ItS ~Xlt.
:\-Iain results from the work\\idth • Stagnation point heat transfa coeffici~nt
can increase more:: than 30°!c) by msuLlingan 1S mesh scre::en sc::r\'~d as a rurbuk:ncc
promoter ln a nozzk (:- inch diarncter).
Experiment characteristics~ 2-D slot nozzles of which
were lI·t lI8 and 11 16 inch,:) Re = ~50 - 22.000
''; ~[easurements of the velocity andturbulence distnbutions U1
submerged jets
AuthorsGardon andAkfirat
1 (1965)
1
•
•
•
1 Saad.
\luJurndar.~kss~h andDouglas( 1980)
<) Staggered arra!"s of impinging air •jets \\ith cross-tle\\" effects
'~ Hole-te-hale spacmg III a •spanwise raw ranging from 2 ta .+jet hole diameters and row-to-ra\\'spacmg in the stream\\ise •direction ranging from 3 to 6 jethole diarncters
~ 0 = 2.5'+ - 5.08,) R~: 3350 - 21.500
.:, Hld= 1.2 and 3
The local variation Ln heat tr3l1st~r ISstrongly penodic Ln space
For a given array configuration and masst10w rate. higher Hld results Ln highero\'crall heat transtt:r ratç and amp htude
Denscr array glves higher overall heattransfer ratl: \\ith lower amplitude but atthe cast of higher mass tlo\\' rate
The eff~ct of superimposed cross-now isto decrease the magnitude of ~u~ and theanenuatlon of Nu,,; pc::a.k. IS affectedapprcciably on for a certain mass tlowrate condition
II
• Table 1.1 (continued)
•
•
•
Huang( 1988)
Yosfuda.Suenaga andEchigo(1990)
1 Polat.~lujumdar
and Douglas(1991 )
\lcCleave(1993 )
1 .:
1
')
<)
Confined mclined slot Jet •discharging from a sharp~dged
re-entry straight channel nozzle\\ith vanous inclinations of 00
•
15') and 30')
The rotating lmpingement cylinderis O.~82 m m diameter. 0.22 mlong and 3.2mm wall thicknessNozzIe \\idth: 6.17mm. 7~Omm.
925mm and 12.33mm. withrespective aspect ratio of 1329. •1LOS. 8.S6 and 6.66H/\V: 3. ~. 5 and 6
2-D impinging jet \\1th gas-solid •suspenslOns~lain nozzle (lOmm by 80mm)and [wo side nozzles (10 mm by3Omm) were usedHl\V=8 and Re= 10.000Loading ratio of parncles {rnean •diarneter: ~8.9~m): 0.1 - 0.8
Confined sIot jets of alf impinging 1 •
on a moving surface withJ\\ithoutthroughtlow at a surface of acylinder. O.~8m diameter2-D nozzle of 10mm x 0.2mHAV=5Re: 8000-58.000
Th~ method of turbulence •generation \Vas placing a bar v.itha diameter liS that of the nozzIe\\idth along the ccnterline of theslot nozz1e •Confined jets from sharp-çdgednozzles \vith \\idths of 6. 12. 18and 25mmHA\!= 1.0 - 1.5
12
\Vith negligible surface motion.stagnation and average Nusselt numberset either nonnal or 15') from normal. IS
similar ta that \\ith :\SN1E standardcontoured entry nozzles at Hi\V=6. \Vluleat HI\V=3. heat transfer becomes muchlarger for the re-Çntry channel nozzles. by
100% for average ~usselt number. by250% for stagnation Nusselt numberunder the conditions testedFor an irnpingernent surface moving athigh speed. lughest heat transfa lS
obtained with the nozzle set at mchnatlonsbetween normal and 15" wIth the Jetopposmg the impingement surfacemotions.
The turbulent mrensity nonnal to the platemcreases markedly (50 %
) near thestagnation point due to the presence of 1
particles rebounding from impmgemenr 1
plate and of the gas-phase ren~rse t10\\"
caused by entramed particles 1
.-\.round the stagnatIon paLOt. th~ ~ureaches 2.7 rimes as great as that of thesingle-phase tlow
lmpingcment surface motIon can dccrcasc~ussdt numbcr for slat Jcts by as much as250/0
Convective heat transfcr for bath smgle andmultiple slOl jets at H/\V<X IS linearlyenhanccd 17°;;) by throughflow.indepcndently of Re. surface motion andc:\1em of heat transfer surface
The method increased average h~at
transfer rates over those of the plainnozzle by 14%. \\ith only a 7'Yo lncreasein nozzle operating pressureThe mcreasc in local Nu can be in theorder of 50c~o. Whlk further from thestagnation point the local Nuis unaffectedby any change in the tlow structure at thenozzle exit
• Table 2.2 (continued)
Colucci andViskanta(1996)
Sing-\<lin. 1
Yasuo and 1
Jeong-Yun :(1995) ,
•
Dlsunile(1994)
\\j Examined an excited Jet.
impinging on a cUf\;ed plate \\itha constant surtàce heat flux andtwo levels of excitation •
2-D siot nozzle \\ith a 3.18mm\\ldth and an aspect ratio of-+-+
A 21cm loudspeaker was used ta •pro\-ide periodic excitation of thejet no\\'
Re=IO.OOO and Ht\V=2.0 1
2-D siot nozzk: 50 mm b,,· 500 1 •
mmRad array of square rods of sideIength: -+mrnRe = 6300 - 57.000 •
H'\V' 2 - 1-+
Sharp~dged orifice and.hyperbolic (contaured. diverging)nozzles of inlet diameter 1.27cm
Low nozzre-ta-plate spacmgs •(O.25<H/D<6.0)
Re: 10.000 - 50.000
•
'Wbeneva the plate was in an excltedt1owfidd. the surface temperature rose.compared to the unexcited reference case
the greatest percent decreases ln Viiappeared to occur in the frequency of100400 Hz range
Energy spectrums acquired undermaxunum excitation showed a drop ln
turbulent intensity between unexciœd andexcited cases
The generatlon of now acceleratlon andturbulence by the rad array near the \\3011.a large heat transfer enhancement (15%
50%) was praduced by the experirnent
The highest heat transfcr enhancementeffect (double that without the rod array)was obtained when the rad array \vas
positioned c10sest ta me wall ( 1mm)
For larger separation distance (H/D=6 Ù).
the local ~usselt numbers displayed weremdepc:ndent of the nozzk geometry
At small sc:paration distlnces (HiD <1.0). two local maXlIna in th~ :\usseltnumber arc evident.
The dTect of nozzle geometr: on thelocation of me second maXima wasconcluded to be dependent upon the Jetoutlet radius.
•
Xiaojun~adcr
( 1997)
o
", '
~leasurement of local convective •heat transfcr coefficients for anobliquely impinging circular aIrJet to a flat plate with liquidcrystal techniqueOblique angles were 90,1. 75'). 601). •and ~5l). \Vith 90.1 being a vertical
JetNozzlc di3.l11eta d=2.05cm
Re: 10.000 and 23.000
Hid: 2. 4. 7 and 10
13
Distribution of local heat transfer shows anon-axis:mmetric pattern. As HJdbecomes smaller. the as:mmetr:' of heattransfcr distnbution becomes morepronounced
The point of maximum heat transtèr shiftsfrom the geometrical impingement pointtoward me compression sicle of the plate.The scalcd shift of maximum heat transferincrcases with a decrcasing oblique angk(more inclination of the jet). The shift IS
found to be sensitive for smalkr HJd
• Table 2.2 (continued)
Hanecb.Tsuchiya.Nakabe andSuzukî09(8) .'\
1
2-D slot nozzle \\ith width of •l5mm and aspect ration of 33lnsertion of two types of ~ylindersin the jet flo\\"The distance between the cylinderinsertion and the target plate was •fixed at 8 mm.
Re =9100 - 10AOaHJW=3 and 5
The maximum :"J'usselt number attamedaround the stagnation point mcreased byabout 20% compared to the one tornormal impmging jet \\lthout the msertionof a cylinderlnsertion of rigidIy suspended cdinderdeteriorates the heat transtèr around thestagnation reglOn \vhile II dfectivetyenhances heat transtèr at the rim of thestammtion resrÎon up to -10%
•
2-D slot nozzk of 55 x 50 mm: •
Re: .+000 - 12.000: Hi\V' 8 -2'+
Inclination: 90 --ter
2-D slor nozzk. smooth md rough •(\\1th fine fins)
W\V=4 - 12. W=7'smm
Two types of turbulence •generators: square rod and thinplate
•
Da"id.Daniel andQlanli( 1999)
l, AbdImonem.
1 ~1ichel andChandrakant(2000)
Slripon(2000)
A puised air jet was lITlpingedupon a heated surface
Sharp-edged eXIt of thl: t~llIplpe
(d= 1-+mm) was used as nozzleDuty cycle representmg the ratIoof pulse cycle on-tune to totalcycle tune: 0.25.0 33 and 0.50
Re = 21.000 -31.000
HJd = -+. 6 and 8
•
•
•
Heat transfer enhancement up to 65°/c)was obtained for a variety of operatmgconditionsThe duty c.....ele W:lS tound to hesigruficant in determining the level of heattranstèr enhancementThe enhancement is most cvtdent tor thesmallest Jet-ta-plate spacmg tested(L/D=-+) while less not3ble If the spacmgincreases to 6 and 8
The distributIon of the local ~usselt 1
numba moves away from $:mmetry asthe mclination angle decreases
The shift of the maximum heat transfer 1
pomt moves towards the upiull sIde of theplate depending on the angle of inclinationand nozzle-to-plate SpaClfiQ.
The stagnatIon and average heat transferrates can be enhanced by up ta 15% and100/0. respectlvelyInsertion of two types of turbulencegenerator in the jet flo\\' provides up to15°/1) mcrease Ln avt::rage heat transtèr ratc
•
Luis andSuresh(2000)
Two sets of round nozzk plates •(diameter 15 3.18mm) \\ith twocharnfer angles of -+ 1.) and 60 l
) anda square-edged orifice (forcomparison)Chamfer depth: 06-+. l. ..n and2.31 mm •
HJd = 1 and -+Re = 5000 - 20.000
Comparcd ta square-edged nozz1es.chamfering the nozzle inld producessignificant reductions (over 20% ) mpressure drop while the average heattransfer coefficient is not strongly affected(bdow l°/ô).
The ratio of average heat transtèrcoefficient ta pressure drop was enhancedby as much as 30.8~~ as a result ofchamfèring the nozzle. with narro\\'chamfering providing the betterperformance.
• 2. -t Flow Stnlcture Studies oj'IJF
In IJF application turbulent jets are generally impinged on surfaces \Vith high
velocities. inducing forced heat convection between the jets and the surtàces.
Experimental results obtained in local heat transtèr measurements are closely related to the
velocity profiles and turbulence levels of the tlow field. Experimental and analytical
studies of the flow field of IIF have provided valuable information regarding its heat
transfer characteristics.
As aforementioned the flow field of an impinging jet can be divided into three
regions: the free-jet region. the impingement region and the wall region. The 11o\v in the
free-jet region is axial and is not affected by the presence of the impingernent plate: at the
nozzle exit. the axial velocity stans to decay and the jet spreads to the surroundings The
developmem of velocity and turbulence intensity of jets has been studied bv se\·eral
researchers. Table 1.3 sho\vs a brief summary of related studies.
Table 2.3 Briefsummary offlow field studies ofIJF•
•
Authors and experiment 1
characteristicsGardon and .\kfirat ( 1965) •I~ ~kasured nozzlt: c~ntcr
lin~ turbukncc intensitiesfor unpinging Jt:ts from •long channel slot nozzks
•
•
Gutmark. Wolfshtein and •VVygnanski (1978)
C Studied turbulentvdocity profile. and •turbulent intcnsity of a 2-D jet from an uncanfinedconraurcd cntry slatnozzle
:\tlain results from the work
Stagnatlon point heat transtèr increased \\1th mcreases m thenozzle-to-surface spacing up ta about 8\V. duc ta mcrèasmgturbulence pen~tration towards tht.: j~t a.'x..lS[ncrease m spacing beyond the potentIal core length results mgraduai decrease fi heat transfer becausè of the decrèJ,Sè ofmean velocity \\ith increasmg H\Vith HI\V > 14. thè initial turbulence lèVd at the nozzle eXIt
contributes much 1ess ta hèat transtèr at the surface than doesmixing induced turbulènceAt the nozzle exit center line for a long channd slot nozzlc.up ta LO~'O turbulence was rneasured
Turbult:nce intensity increascs sharply \\ith distance from thenazzle exit. then decreascs sharply \\Hh approach ta thesurfacèThe effect af the impingement surtàce on the turbulenceproperties dos not propagatc back inta the Ho\\" beyand 02Hfram the surface
L5
•
•
•
y osluda. Suenaga andEchigo (1 990)
,) The detailed turbulencestructure of the twodimensionai unpmgmgJet \\ith gas-solid wasmeasured by means ofLDA
~ Other features aredescribed in Table 2.2
L~tle and Vv'ebb (1994)
: The flo\\" structure wasinvestigated usmg LDVand wall pressuremeasurerncntsCircular tube nozzle:7 8mm diameœr And10.9mm diameterHtd:: 0.1 - 10.0Red = 3600 - 27.000
Janice and Suresh (1997)
: Experimenrally studiedusing LDV the t10w fieldof an axis:mmetric.confined and submergedturbulent jet unpmgmgnonnally on a tlat plate.through velocity andturbulence measurernents
: The nozzle aspect rationwas unity \\ith diametersof 3.18 and 6.35mm
" fi/d::; 4
•
•
•
•
•
•
•
•
•
•
•
•
•
Table 1.3 (continued)
The large difference bet\veen the Lnertla forces of gas and 1
solid phases is responsible for aU of the gas-solid interactlons 1
The most notable feature of the tlow 15 the presence of 1
particles rebounding from the irnpingernent plate and goingupstrearn against the oncoming flo\\" l'
The resuiting gas-solid interaction near the stagnatIon pomt IS
violent and induees gas-phase reverse tlO\v 1
The mechanism of the heat transtèr enhancement around me 1
stagnation point is attributable to the drastic change m the1
turbulence structure 1
Flow structure measurements reveal sigruficant mcre:lSes mboth mean velocity and R...'v(S turbulence tluctuauons as thenozzle-plate spacing is decrcasedConslderable he~lt transtèr enhancemenr \Vas observed due toglobal-<:onrinuity-torced acceleration of the unpmgmg tluidJ.S it escapes from the nozzle-plate gap. as weIl J.S slgrut1camincreases m the turbulence kvelThe above phenomena ~ield a stagnation point mlillmUITI and :an inner and outer peak in the local heat transfer ;The location of outer peak was found ta coinclde wIth a local:aximum Ln turbulence tluctuatlons. SUggcsting conslderabh· :
-- - • 1
hürher turbulent transport th~re '.-\ toroicial reclrculation zon~ was observed and mapped andthe location of the center of the torOld moved radiallyoutward~ both \\1th an increase in Re:TIolds number and \\ithan increast: in nozzIe-to-urget plate spacingThe center of the torOld moved nearer to the target plate \\lm :
an mcrease fi Re~nolds number !
The maximum velocity in the wall Jet reglon occurrcd veryclose ta the impingernent plate for r/d= 1. with the ma."'\.ill1umturbulence levels occurring at rid =1An increase in the nozzie-to-target plate spacing was sho\\nta reduce the magnitudes of the radial velociues as \\dl as thepeak turbulence Lntensities in the flow fieldIncreasing the nozzle diametcr resulted in a d~crease 10 peakradial velocitv. but an increase in peak turbulence k,ds
16
•
•
•
•
~liranda and Campos (1999) •
':' [nvestigated a laminar jetflow confined by a •conical wall and an
irnpinging plare) The Navier-Stokes.
~quations werenumerically solved by afiaire differencetechnique and the resultscompared with LDAdata. the latter aIsocovenng the transitionand turbuknt tlo\\"regunes
~lansoo. Han. Geunyoung. •100n md Dong (1000)
~ The distributlon of meanvelociry and vdocity •tlucruatlon on theconcave surtàce weremeasured ln three tlowregions by using a LDA •
~ Emphasis \\as placed onmeasunng turbulent Jettlo\\" characteristicsincluding impinging md •e\·olving wall jets andintcrprctmg heat transferdata. particuIarh-
2.5 Closure
Table 2.3 (continued)
The transition was found to start in the impmgemem regionat a jet R~!l1olds number of around 1600The jet Re~l1olds number and the inlet velociry profile have ::istrong influence on the whole laminar tlow. while the nozzkto-plate distance is influential oruy in the expansion reglon
At low jet Re~nolds numbers. the t1uid far from the plateacquires radial velociry and a short reclrculation zone closete the conical wall is observed.-\r high Jet Reynolds numbers. this reclrculatlng zont:enlarges. the fluid flows radially in a thin channd attached ta
the unpingement plate and a second recircularion zonedevelops in the expansion region. close ta the plateTransition from laminar to turbuknt tlow probably begins Ln
the impingement reglOn at an Re around 1600
The patencial core length becomes shorter for Re.....=2370 chan 1
for Re..~-=890 and 1~80. Stagnation region thlcknesses art: :approximately ~qual ta 2\V
For the case of HJ\V=O.~. thè effect of tluld J.ccd~ratlon hasbeen observed. The thickness of the wall jet 15 small~r forHAV=0. ~ tllan for HI\V=1. which is the reason why the hé:attranstèr r::ite for H/\V=O.~ IS hight:r than for Hi\V= 1The occurrence of sé:condary peaks and their locatIons ha\"~
bèen explained from the \·ariation of measured velocitytluctuations of the wall jets evol"ing J.long the stœam\\lSCdirectIonThe increase of stagnatIOn heat transfcr r~te for 2-3 < HJ\V <5-6 has been successfully explained due CO thé: stCèP mcreas~
of velOClty tluctuations mt:asurcd in fn::e and Impingmg J~ts
•
From above literature reVlew, it has been found that curvature. roughness and
turbulence can affect heat transfer ta sorne extent depending on parameter ranges
concemed. However. no enough comparison has been found to make systematically in
companson of various fluid nozzles, different types of curved sutface and various
protrusions and its neighbor region. AIso. there are no reports relating to the effects of
different types of turbulence generator insertion and their location in the jet tlO\V on heat
17
•
•
•
differem types of turbulence generator insenion and their location in the jet flow on heat
transtèr. The purpose of this project IS try to fill these gaps via experimental studies of
impinging jet heat transfer using various types of nozzles .
18
EXPERIMENTAL SETUP AND METHODOLOGY
3.1 Introduction
This chapter presents details of the experimental set-up. the design of the
impinging jet flow system. the rough nozzle and curved target protrusion used.
experimental techniques employed for the heat transfer as weil as the experimental
procedure. Data reduetion and reproducibility tests will be also covered.
•
•
3.2 Experimental.Apparatlls
3.2.1 Overafl t.-.xperimenral Setup
Power 5 upply
f .:~- _.~';~ i Impin~emeDt
1...:- ~_";J -1!I~t ! Confinement
~...... ~ .--- Surface
Thennocouples !1j Jet Plenum
~. 1 \ . ....1.. ~ambe,r
_"', ,.I! ~ =~ _ :.~~U. i~~C .~':== .' . Sict i :-------
. l j 'N0ZZ'l ......Il ~tainlessInsulation ~ U: 1 Steel -------
Layers ~ Foil
!!!-r----~Computerwith DAS
Figure 3.1 Schematic of the overall experimental setup
19
Blower
•
•
Figure 3.1 is a schematic diagram of the overall experimental set-up with the
auxiliary instruments for heat transfer measurements. The experimental set-up consists
of a centrifugaI blower (Industrial Model of BALDOR Electric Co., USA), circular
plastic pipes. two gate valves, an in-line pneumatic flow meter (FL7411, Omega
Company, USA) with a maximum flowrate of 40 SCFM (68 m3/h), a cylindrical plenum
chamber (20 cm in diameter and 30 cm in length), a slot nozzle (smooth or rough) and a
flat reetangular acrylic plate along with temperature measurement equipment
(thermocouples, DAS and power supply). In the case of non-planar surface experiment, a
semi-cylinder was attached on the flat plate.
3.2.2 lmpingement Plate and iVon-planar Target
The impingement plate is 300 mm long, 120 mm wide and la mm thick and is
made of acrylic. The edge effects caused by the difference (20mm) bet\veen plate width
and nozzle length is neglected. ft was machined with tvJo ears on each side. respectively,
50 as to be mountable at different horizontal locations on a sturdy platform with several
groove couples. The nozzle-te-plate spacing can be set in the range of 1.6 - 18 times
nozzle width (W=5.0 and 7.5 mm) by inserting the plate in one of the groove couples.
The impinging jet tlow is confined between two flat parallel surfaces. flush and extended.
respectively, from the nozzle exit and irnpingement surtace.
Semi-<ylinder
Impingement 120Plate
•300 -<- Symmetricalline
Figure 3.2. Target protrusion and its location on the fiat impingement plate(not ta scale)
20
•
•
The non-planar target employed here is a semi-cylinder. Figure 32 shows the
semi-cylinder and its location on the impingement plate. The length of the semi-cylinder
is 80 mm. and the it was eut from a cylinder with a diameter of 64 mm [t i5 located at
the geometric center of the flat base plate. The cylinder is attached tirmly ta the tlat plate
with double-sided adhesive tape.
3.2.3 Design of'SIol .Vo==les
3.2.3.1 Snl00lh SIOl ~Vo==les
The widths of the tViO smooth slot nozzles are 75 mm and 50 mm and the lengths
are bath 100mm. \vith aspect ratio of 13.3 and 20. respeetively They are considered as
2-D nozzles. Of the [\vO smooth nozzles. the \vider one. named S:\IS~O 1. i5 a circular
plate with a 7.5 x: 100 mm slot opening. exactIy the same as the rectangular cross-section
channel before the nozzle. Therefore the tlow exiting trom S~(S:\O 1 can be considered
nearly unidirectional. The narro\ver one. named Sy(S0i"02 is tàbricated by blocking the
slot opening with two metal strips. reducing the open \vidth of the nozzle to 5 Omm. The
air now from S~(S~O 1. is more complex due to the sudden contracting of the air tlO\V
passage width from 7.5 mm ta SOmm. The turbulence leveI of air t10\v exiting from
S~IS~02 is expected te be higher than that from S\150iO 1.
3.1.3.2 ROllgh Slo{ ~Vo==les
Figure 3.3 GeometricaI configurations of three rough-edged nozzles(the numbers of the jaws are not identical ta the real numbers)
• (3) Right Triangular (1aws 1) lb) Equilateral Triangular daws 2) le) Rcctangular (Jaws 3)
21
• To enhance the convective heat transfer between air jet tlow and impingement
surtàce. higher turbulence level of the air jet is desired. In this project three rough-edged
nozzies are designed for this purpose. aiming at promoting the turbulent level of the air
tlow. The three types of rough nozzles are right-triangular. equilateral-triangular and
rectangular rough nozzles. which are named Jaws 1. Jaws :2 and Jaws3. respectively. for
identification. They are fabricated by attaching different tins on each side of nozzle
SNIS0iü l. Figure 3.3 shows the configurations of the three rough nozzles.
In order to compare the heat transfer performance betvieen the smooth nozzles
and rough nozzles. the open areas of the three rough nozzies are made equal ta that of
nozzle S~[S0i02. i.e.. the narrO\Ver nozzle \vith a width of 5.0 mm. by measuring the totai
area of aIl tins. The actual dimensions of the three types of tins are listed in table .3
The ether geometric parameters are listed in table 3.2.
Table 3.1 The dimensions of three types of fins
•Right Triangular
(Ja\vs 1)
3mm
5mm
Equilateral Triangular(Ja\vs 2)
Rectangular(Ja\vs 3)
~mm
-+mm
Table 3.2 Geometric parameters of the nozzles used
~ozzIe 'Jietted perimeter Hydraulic Open area Decrease of open-area
(mm) diameter (mm) (mm:) compared with nozzieS~IS~OI
S~IS?'iO1 215 1~.0 750 0
srvrS0I02 210 9 5 500 500 /0
Jaws 1 402 5.0 500 50°'0
Jaws 2 354 5.6 500 50°'0
Jaws 3 'ÎÎ 62 500 50°'0• --'--
,.,
•
•
3.2. -1 Turbulators
ln arder to enhance the turbulence level in the jet flow locally. three types of
turbulence generators (turbulators) were used in the present study, i. e., square rads, strips
and perforated plate made of stainless steel. The first two are machined, respectively. as
5 mm x 5mm and 5 mm x 2 mm. The perforated plate is a 37 x 78 x 0.8 mm rectangular
plate with 54 apertures of 3 mm diarneter. The open area ratio is 17.660/0. The rods and
strips are fixed vertically at selected locations in the jet flow by supporting plates \Vith
approximately same-size hales. The perforated plate is welded with one end of a long
thin steel stick the other end of which is gripped by a clamp, thus fixing the plate at
desired positions. The turbulators were placed such that the main flowfield remained
nominally two directional.
Figure 3.4 shows schematically the location of the inserted turbulator. For the
cases of square rads and strips, the insertion location is defined by xJW and y/\V.
respectively, where x/W represents the ratio of the horizontal distance of the turbulator
from stagnation line ( X direction) to the \vidth of the nozzle and, y/W represents the ratio
of vertical distance from the impingement plate (Y direction) to the width of the nozzle,
as sho\vn in the figure ..
Siot
Jet Confinement
___\_V~~i ~ 1'---.-__-.::..-/_Surface
OutflowFlow Flow
-+- .. .. -. Outfla\\"/Y
Stagnation Line..........t ~ .... Turbulalor
,~ ~
Target Plate ConfinementSurface
•Figure 3.4 Slot jet impingement on an isoflux heated plate
•
•
For the rads and strips two values of y/W were tested, viz. is 1 and 2 while xIW
was either 0 or 2. For the perforated plate y/W values tested were l, 2, 5 and 8 while
xIW was either 0 or -5.2. Both single insertion of rod/strip and double insertion of them
were tested, expectedly producing asymmetrical or symmetrical jet flow on the
impingement surface. The dimensionless x!W and y/W are based on the geometrical axis
for the rod and strip and the left edge for the perforated plate. IDW is 4 for rods/strips and
it is 4 or 10 for the perforated plate turbulator.
3.3 Experimental Techniques
3. 3.1 Heat Transler lvfeasurenzents
Heat transfer rate is obtained here by measuring the temperatures of the jet exit
and the impingement surface. which is heated with a thin metal foil to give a uniform
heat flux, with 0.125 mm diameter type K thermocouples of which response time is less
than 5 ms. The jet temperature is measured by placing the tip of a thennocouple at the
nozzle exit while measuring impingement surface temperatures needs sorne skillful effort
ta minimize the expenmental errors and compensate for the heat loss. Total heat lasses
include conduction through the insulating materials, the lead w-ires of thermocouples, the
stainless steel foil and radiation loss. Total heat loss is estimated to be less than 4~/o of
the generated heat flux within the estimated uncertainty range. The uncertainty analysis
is discussed in Appendix 1.
O.0254mm thickstalnless steel shok
•..;.....;. ..:....:.:.:::.:::.:::.;::.;: ..;:..:...;:..;...:...:...: :'. .;. rt:·:::·::··:::·::··;.:·;·:·:::·;·..:··;· .. ·;:···::··:···~~
";"~>- '- \:~.~""''''~.:';\.~ ".:~>:,~ ~..":- :~'\.:."':~~.0:.'\."~~\.,,,:,,'\:::".)\:\ '''::-:'''''.~:;'':;'.<''>;''~.~-';;~:. ~.'." ..~. ,~ ;~."~ ..\,~~~.:s\.,"\.::--:,,,,');
Fiber glass layer(4mm lhick)
Acrylic layer(11mm thick)
Ceramic fiber layer(3cm thick)
$Iainless steel layer(12mm thick)
Figure 3.5 Sectional view of impingement plate with thermocouples and insulation layers•" / ('
l '/ ",';1 1'4-
.- .1 ,1Type K thermocouples(O. 125mm diameter)
• Figure 3.5 shows a sectional view of the impingement plate with insulation layers
and sorne thermocouples. The heating source is a thin type 302 stainless steel foil (SHI.Nl
STOCK QQ-S-766. LYON rNDUSTRIES) with an thickness of 0.025 mm (0.001"). 57
mm in width and 300 mm in length. The foil is glued with a 4 mm thick tiber glass
(thermal conductivity = 0.038 W/m-K) plate using double-sided adhesive tape. 31 fine
thermocouples were inserted one by one through holes of 0.95 mm machined through the
thickness of the fiber glass plate and led to outside between the fiber glass plate and
acrylic plate. The thermocouples tips are carefully set undemeath the stainless steel foil
in order not to produce bumps or dimples. which can adversely affect the experimental
results. on the impingemem surface. Other two layers are used to insulate further the
backside of the target plate and the heater assembly.
•Flat Base Plate
J }tecl foil
r' 1
;#Confinement Surfaces~
Thennocouplcs
•
"" -~ - - ..l i<t L );1 i 1
1 1 1l ~ 1
Figure 3.6 The schematic ofthermocoùple arrangement and test section(not aIl the thermocouples have been drawn)
Figure 3.6 is a schematic of the thermocouple arrangement and test section. The
x-y coordinate system is defined for the studv of the effects of turbulators on heat
transfer. At x!W is less than 9. the spacing of two consecutive thermocouples on the
impingement surface is half the width of nozzle SNISNO l, i. e. 3.75 mm, while out of the
region the spacing is the nozzle width, 7.5mm extend to xJW=16. One thermocouple is
25
•
•
•
located at the nozzle exit for measuring jet temperature. On the side where there are
more thermocouples, sorne two or three thermocouples are arranged at the same X
location to verify the assumption of 1-0 flow.
Electrical resistance heating was used to Impose a uniform heat tlux on the
impingement surface. The heat flux was controlled by regulating the electric CUITent in
the circuit. The type 302 steel foil is chosen because of its high electrical resistance
(specifie resistivity is 7.4 x lO-7Q.m) and uniform physical properties. Two drilled
square copper bars. served as electrodes on which the steel foil is scre\ved tirmly te
ensure good electrical contact betvleen them, are attached on bath sides of the plate. The
copper electrodes are connected to an adjustable OC power supply. allo\ving a adjustable
current up to 28 amp Preliminary testing was performed ta insure uniformity and
symmetry of the surtàce heat flux by recording temperature data of the plate at various
po\ver inputs without the jet impinging. It was found that the ternperature difference at
any t\VO locations on the plate was less than 0.6°(.
The detail of the thermocouple arrangement for non-planar IJF surtàce is shown
later in chapter 5.
3.3.2 Pressure J/easurenzenrs
Pressure drops across the nozzles \vere measured uSlng a pitot tube and an
inclined manometer. The pressure data \vere used te calculate the nozzle discharge
coefficients and pumping power.
3.-+ Experimental Procedure and Plan
The controlled parameters in the experiments are the nozzle-to-plate spacing. the
air flow rate, the electric power input. the location of the perforated plate. the types of
protrusion and of nozzles. AlI runs of the experiments are conducted under steady state
of heat transfer. During a run the air tlow rate and electrical CUITent are monitored and
adjusted to maintain the same values. The blower and power supply are switched on first
for a while (more than half an hour) untiI temperature distribution along the impingement
26
•
•
•
surface remains constant, as weil as the temperature of the air jet exit. After reaching the
steady state condition, ail the controlled parameters and temperature data are recorded.
Each data point in a run consists of moving the target plate to the desired horizontal
position (H/W), in the case of testing the effects of turbulators fixing the turbulators in
proper position~ waiting for the heater temperature to reach steady state. and signaling the
computer ta read and record the thermocouple data via a data acquisition analog-to
digital (AID) board (CIO-DAS 32/Jr type from OlVŒGA company) and a LABTECH
data sampling softvvare. into the hard disk of a Pentium-133 ~U-Iz PC for processing. A11
temperature data are recorded at a frequency of 5 Hz in 100 seconds. Each temperature is
calculated as an average of 500 transient temperature readings. sufficiently representing
the real temperature of that point. The experimental results \vere found to be independent
of sample range of over 500 points.
Table 3.3. Ranges of experimental parameters
Experimental Parameter1
Ramœ or T vpe
Reynolds number based on the jet exit1
1500 - 12.000
Jet velocity at nozzle exit (mis) 3 15 - 25.2
2-D nozzle size (mm)1 75 mm x 100 mm and 50 mm x 100 mm1
Corresponding flowrate (rn"!h)1
8.5 - 68
Type of nozzle (in terms of exit structure) Smooth. triangular tin and rectangular tin
~ozzle-to-plate spacing (LI\V) 1.6- 18
Type oftarget protrusion Semi-cylinder
Diameter of the cylinder protrusion (mm) 64 mm
Position of the protrusion Center or certain distance shift from center
Temperature ofair jet exit (uC) Below 30
Temperature on the surface (uC) 40 - 85
Electric CUITent for heating (ampere) la - 28
Heat flux on the impingement surface 800 - 6,200
(W/m:)
Type/size of insertion object Steel rods (5 x 5 mm-), strips (5 x 2 mm-)
and perforated plate (37 x 78 mm~)
,-_1
• Both parts were performed regarding the etfects of operating parameters on local
and average heat transfer coefficients and Nusselt number calculated trom the measured
temperature data along the impingement surface. The ranges ofexperimental parameters
tested are given in Table 3.3.
3.5 Data Reduction
3.5.1 Analysis ofHeat Balance on Studied Suiface
The objective of the data reduction in llF studies is to obtain the local and average
heat transfer coefficients on the impinged surface by using a model ta calculate the heat
lasses from the heated surface and thus to obtain an accurate value of the convective heat
flux.
In order ta evaluate the convective heat flux dissipated from the heated target
surface, it may be proposed that the total heat generated from the metal foil heater, Qt. is
converted into the foLIowing four heat transfer modes: (1) convective heat dissipated from
• the heated sutface, QI:. i.e. the main part: (2) intemal-energy change of the stainless steel
sheet, Qi. (3) conductive heat loss. Qk: and (4) radiative heat loss, Qr~ The heat balance is
thus as
l.e.
Qt = Qc -+- Qi ~ Qk ~ Qr
Qc = Qt - Qi - Qk - Qr
(3-1)
(3-2)
The net convective heat, Qc. dissipated trom the target surface ta the air jet. can then be
calculated from Eqn. (3-2). The total power input is Qt which equaIs eR., where 1 is the
electric CUITent of DC power supply and R is the electric resistance of the steel foiL Qi is
the intemaI-energy change of the stainIess steel foil. Qk is the conductive heat loss to the
insulation layer. Qr is the radiative heat loss from the stainIess steel sheet surface through
insulation layers. Figure 3.7 is a schematic of a part of the electrically heated foil which
consists of three equal-sized sections with dimensions of a breadth Band length L.
•28
InternaI conductive heat Q1. Q~
Bti--t-l
Electnc heat Qt 1-. t;:.;i-,;"l ~:+-'"=""!'~--_oo:===t.....:C....:o ...nvective heat Q,;ti
•L
Figure 3.7 Schematic ofheat tlux description ofthree consecutivesections of the heated thermofoil
In present study up to 32 type K thermocouples are used and insened in a ro\V on
the flat surface or cylinder surface extending outwards trom the stagnation line. Because
the highest temperature on the surface is below 85°C and the surtàce of the heating foil is
very smooth. the emissivity of the foil surface \vas estimated to be less than 0.072 at
1GODe. An estimation of the radiation heat loss \vas performed and found the radiation
heat amounted to less than 1.5% of the total and hence neglected here. The heat
conduction loss through the insulation were also examined here by inserting sorne
• thermocouples on the backside of the fiber glass layer across \'vhich the temperature
difference \vas measured ta be less than 12°C at the hottest point (the tàrthest point from
stagnation line). Correction ta the total po\ver dissipated from the impingement surtàce
\vas made for the heat conduction. \vhich amounted to less than -1-°0 of the total. The
detail of estimation of heat losses is described in Appendix 2.
3.5.2 Local Heac Transjèr Coefficient
The heat transtèr coefficient is defined as the surface heat flux divided by the
difference between the jet temperature and the surface temperature. According to
Newton' s law of cooling, the local convective heat transfer coefficient h" is defined by
(3-3 )
where Tw.x and Tj represent the local wall temperature and jet exit temperature.
respectively. q..: is the local convective heat flux. QcIA.
•
• The temperatures in the three sections were measured by three thermocouples
insened at their respective centers and assumed uniform (i. e. isothermal) on each surface
ofthem. Taking the middle section as the target. the heat balance for it is
(3 -4)
where Qi and Q~ are conductive heat outvlards and inwards the middle target section
from the t\\iO neighbor sections due to temperature differences. assuming tl-l > ti > ti-l. Qt
is the electrical input to section and Q;; is the convective heat to the jet flow with a
temperature T l - The four heat fluxes can be calculated. respectively. as
and
Q1 :: kfAx ll-t-:Q: = kt' Ax l'-L-{;
Qt = eR
(3-5)
(3-6)
(3-ï)
•13-8)
where kr is the thermal conductivity of the heating foil. Ax is the cross-sectional area of
the foil. l is the value of the electrical current and R is the electric resistance of section i.
Substituting Eqns. (3-5) - (3-8) into (3-4) and rearranging the both sides. the local
convective heat transfer coefficient of section i. hi. can be evaluated as
1 : ~ KrA.r ( . _ . _ Î )hi = .-t ( 1:- T; ) [I R L 1. - . II - , - li ] ( 3-9)
\vhere [ is read trom power supply, tl.l. tl • ti-l. and Tare read tram calibrated
thermocouples. As and A" are calculated from defined data of thickness (0). \vidth (B)
and length (L) of heating foil respectively. Here As and A"( are calculated
•
A-;=L· B
and
Ax = D· B
R is calculated from the resistivity of foi 1. Tl, and its size.
R = TlL/Ax
Substituting Eqn. (3-10), (3-11) and (3-12) into (3-9), we can get
.= _1_ r'7 -J- kfD . _ Îhl L _ T [ DB: . L: (tl. 1 T t1 - 1 - tl )]
30
(3-10)
(3-11)
(3-12)
(3-13)
• 3.5.3 Average Reat Transfer Coefficient and Nusselt Number
Practically, the area-average heat transfer coefficient is more useful than the local
heat transfer coefficient in the design of jet impingement application. For a 2-D siot-jet
flow the average heat transfer coefficient on the heated target surface can be evaluated
with the following equation., in the case shown in Figure 3.7
(3-14)
where the integration of equation (3-14) is approximated by summing up all products of
the reciprocal of local temperature difference and the corresponding length of segment
LlX along the impingement plate. Any other area-average heat transfer coefficients can
also be calculated by using the corresponding total length and the upper and lower limits
in equation (3-14) instead ofx.
The local Nusselt numbers can be evaluated based on the thennai conductivity of
the air at nozzle exit condition and the width of the jet nozzle. respectively as.
• Nux = ~v (3-15)
and
hsw XNus = T at = 0 (3-16)
where w is the nozzle width. k; is thennal conductivity of air at nozzle exit temperature.
Accordingly, the average Nusselt number based on the length of heated target
surface can be defined as
- lM'NII=-r;
3.6 Data Reproducibility
(3 -17)
•In order to test the reproducibility of the experimental results, several runs were
repeated at randomly selected conditions. Figure 3.8 shows three sets of temperature
distribution data at the specified experimental conditions and figure 3.9 shows the heat
transfer coefficients data calculated [rom the temperature data in figure 3.8. [t can be
31
• seen that the data correspond with each other quite weil. The standard deviation of heat
transfer coefficient is below 3% for ail runs tested, which can he considered satisfactory.
........'..- Run3
-:- Run 1
-=- Run 2
Flat surface
Heat nux: 3054 w/m l
60~---------------
Re.Z1~OOO =~ :~::/'55 _ -r...- Ru" 3
g~ t~ ,..J!,250- ~:!t~- f'" ~.~. .. Flat surface
~ Heat nux: 3054 w/m2
~ HIW=6 for nozz1e SMSN01
40~
350 -----------------
ij,.,- 300 - Î:iN~ '\.
~ 250 - \ë .~ -; HJW=6 tor noule SMSN01
~ 200 _ ~ Re.=1~OOO
o , ..~ ~ ...!l/lI 150-~ ,~ ...........
~ 100- ~
35----------------
Figure 3.8 Temperature measurement repeatability•a 4 8 xJW 12 16 20
so -----------------o 4 8 xJW 12 '6 20
Figure 3.9 Heat transfer measurement reproducibility
Heat flux (W/m 2)
~ 4950
- 41813054
2457
Flatsurface
Curvedsurface
- 300 --------------------~
N
E~ 250 :t-.i \i 200-\~ -~.! 150 - ._~
en 1:-
; ~~ 100 - ,.~- ~: A.'.;oC
~ 50 ë
o..J
o 4~.3 ~- --_-_
1614121042o 6xJW 8
Figure 3.10 Heat transfer coefficients measuredat various heat fluxes•
32
•
•
•
Figure 3. 11 shows the heat transfer coefficient data measured at vanous heat
t1uxes for non-planar surface. AlI the data correspond to each other quite weIl.
Therefore different values of heat flux do not affect the heat transfer coefficients
obtained, thus the effect can be neglected.
The uncertainty analysis of experimental results is described in Appendix L_ The
estimated relative uncertainty of heat transfer coefficient is generally below 6% .
33
• ICHAPTER4
REAT TRANSFER CHARACTERISTICS PLANAR SURFACE
4.1 Introduction
This chapter presents and discusses the experimentally measured heat transfer
results obtained using a plane (flat) impingement surface. The parametric effects of jet
Reynolds numbers and dimensionless nozzle-to-plate spacing on heat transfer
characteristics of the heated target surface are explored using bath smooth and "jawed"
slot nozzles. The parameters studied are:
(1) Jet Reynolds number~
• (2) Dimensionless nozzle-to-plate spacing, HJW~
(3) Nozzle type, i.e. two smooth nozzJes and three rough nozzles (see Chapter 3
for details)~
(4) Effect of inserting three types of turbulators at various locations in jet flow.
4.2 Heat Transfer on the Planar Surface
-J.2.1 Temperature Distribution along the Surface
Figures 4.1 and 4.2 display the measured impingement surface temperature
distributions obtained for constant heat flux of 4950W/m2 for smooth nozzle and rough
nozzles, respectively. The higher temperatures (farther locations from the stagnation
line) correspond to lower heat transfer rates The temperatures tested are low enough to
consider the heat transfer coefficients to be those for constant (temperature-independent)
physical properties .
•
• 9l------------~ 100 -------------
-70 (J
a .,.
-/.2.2 Heat Transfer Coefficients ofTwo Smooth /v'o==les•
~-------------
a 5 10 15 3J 25xIW
RglI'e4.1 TEf'Tl:Sët1J'e cistrituiCIl CIl ftà SLIfa:e(smxitll"lJZ2fe)
20-------------a 5 10 15 3l 25
'ANIFigse 42 Terrperatlre dstrituiCll CX1 flet sufaœ
(right-triargular rozzje)
•
Heat transfer rate is generally characterized as heat transfer coefficient. In present
study steady-state convective heat transfer plays the most important part in aIl heat
transfer existing between the hot flat surface and cooling air jet. plus the negligible
radiation and conduction heat losses. Iwo smooth nozzles of W=5mm and W=7.5mm
will be studied hereafter. Both nozzles have the same length of 100mm but have
different width. The other difference is that the cross-section area of air inlet channel is
the same as the wider nozzle, 7.5 x 100 mm. so the turbulence level of the air jet can be
considered as relatively low. While for the narrower nozzle of W=5mm, due to the
contraction of air flow caused by the change of width from 7.Smm ta 5mm, more
turbulence is expected for the air jet impinging on flat surface, thus improving heat
transfer to sorne extent.
35
• -+.2.2.1 Smooth Nozzle ofW=5mm
2520
-:- RI.a12000.
-=- RI.agODO
---.-.J-- RI.-6000
~ RI.-JOOD
-:- RI.-1500
1510
H/W1I6
5
smoottl nozzl.: waSmm
ox/W
Figure 4.4 Local heat transfer coefficient distribution
(HiW=6)
500 ~----~------------
Q------------------
';QI~ 100iiuo~
20
-)- Re."2DlXI
-=- Ra.,-'JOOO
......:.-R..~
-':'- R..C30lXl
-:- R...,S)Q
50-----------------
o 10 15xIW
Figlft 4.3 Local heat transter cœtficient distribUion
(tWF2.2)
•
o ------------.,-------
600 -----------------
2520
-:- ~2000:
-=-~
-0-~
~. Ra..rmJ-:- ~!iOO
151050------------------
o'1NI
Figure 4.6 Local heat transfer coefficient distribution(HIW=18)
360 ------------------~
N
.É XlO ---~ ~~>
-:- Re.-12000
-~ Re.-9000
~ Re.~DOo
-r;>- Re.a3DOO
-;- Re.-1500
Smoottl "oale: Wa5mm
HI'Nz9
o 5 10 15 20 25xIW
Figure 4.5 Local heat transfer coefficient distribution(HIW=9)
•Figures 4.3 - 4.6 exhibit local heat transfer coefficient distributions for the
smooth nozzle (W=5mm), corresponding, respectively, to small and large nozzle-to-plate
spacings of2.2, 6, 9 and 18 at Reynolds numbers ranging from 1,500 to 12,000. It can be
36
•
•
seen that in general the local heat transfer coefficient decreases as xIW (distance from
stagnation point) increases at ail IDW and Reynolds numbers with few exceptions where
secondaI]' maxima occur. It is worth mentioning that the second maximum in Figure 4.1
corresponds to the minimum in Figure 4.4, which rneans higher local temperatures
resulting in smaller local heat transfer coefficients.
The shape of the heat transfer profiles over spacings, HfW, of 2.2, 6 and 9 are
more or less aIike, with a second maxima away from the stagnation line for higher
Reynolds numbers. The explanation for this phenomenon was first given by Gardon and
Akfirat (1965)~ they proposed that the minimum marks the onset of boundary layer
transition from laminar to turbulent, while the maximum marks the completion of this
transition. The onset of transition can be related ta the end of the region of steep pressure
gradient at the impingement surface~ 5uch gradient provides the mechanism for
maintaining a laminar-like boundary layer under fully turbulent jets. \Vhile for the case of
H/W= 18 (Figure 4.6), which is much farther than the previous three spacings. there is
not any second maximum at any Reynolds number.
Figures 4.7 - 4.10 show local and average heat transfer coefficients,
corresponding ta the same conditions as those for Figures 4.3 - 4.6. respectively.
o-----~--~----- 0--------------
500 --------------Rew hloc havg
12000 -.-6000 __
1500 -..-
o 5 10 xJW 15 20 25
Figure 4.8 Comparison between local and averageheat transfer coemcients (HJW=6)
havgRew1200060001500
Smooth noule: W=5mm
500 --------------
450-
o 5 10 xIW 15 20 25
Figure 4.7 Comparison between local and average
heat transfer coemcients (HJW=2.2)
•.,~
JI
0----------------
havgRew12000600015'10
Smooth nome: W=5mm
0----------------
400 ------------------,
o 5 10 xJW 15 20 25
Figure 4.10 Comparfson between local and averageheat transfer coemclents (H/w=18)
Re..., hloc havg12000 -.- _.:r-
15QOO _ -w-1500 -.- -7
Smooth noz2te: W=5mm
600 ~--------------.....,
•300 -::. •
":"'r"-;c.-. .......-=_ e _.~ -..200 - . ........---r........-r-'CGr, ~_
........~-'-Gce-~r _
100 -7 ~_ ~--....:-.~ ;
~ ~., ~-'7-'7-"';'-<?-~-?-'7'.............. 1
a 5 10 xIW 15 20 25
Figure 4.9 Comparison between local and averageheat transfer coefficients (H/W=9)
~N 500-e "'- ~-~ •.:.,..ë 400 - • ~-:.:.u •UE8u..~1ftCl!--IlsU%
•
•
The average heat transfer coefficient at a specified X/VYT represents the average
value of the local heat transfer coefficient from stagnation point ta the position
concemed. It cao be seen from the graphs that average values are always greater than
local values and the curves of average values are much smoother due to the integration
results .
.f.2.2.2 Smooth j\'o=;:Ie with "VV=7.5rnrn
350 -----------------
o~----------------2016
-:;- Re.,,=12000
-=- Re.=9oo0
---ù- Re.=6oo0
-":'- Re.=3000
-:- Re.=1500
84
Smooth noule: W=7.5mm
o0------------------
300 ------------------
5Z :c"'e -~ 250 := :_
12xJW
Figure 4.12 Local heat transfer coefficient distribution
(HIVV=4)
2016
-:- Re.=12COO
-:;- Re.,,=900a--.,..- Re.=600a
-";"- Re.=300a-:- Re.=150a
1284
Smooth noule: W=7.5mm
HJW=1.S
'- -
oxIW
Figure 4.11 Local heat transfer coefficient distribution
(HIW=1.6)•38
• 3Sa -------------------
Q------------------2016
-)- Re.,.=12000
-=- Re.,.=9000~ Re.,.=6000---":'- Re..=3000-:- Re..=1500
128
HJW=8
4
Smooth nouJe: W=7.5mm
o0--------------------
300 -------------------
xmFigure 4.14 Local heat transfer coefficient distribution
(HIW=8)
2016
- Re..=1:000
-=- Re..=9000
-ù- Re..~OO
-7- Re..z30OO
Re..=1500
128
HJW=6
4
Smooth nouJe: W=7.5mm
ox/W
Figure 4.13 Local heat transfer coefficient distribution(HIW=6)
• 350 -------------------
0--------------------
-:- Re.,.:12000
-=- Re.,.=9000-w- Re..=6000-."'- Re.,.=3000
-:- Re.,.=1500
H/W=12
Smooth nozzle: W=7.Smm
0--------------------
350 -------------------
a 4 8 xm 12 16 20
Figure 4.16 Local heat transfer coefficient distribution(HJW=12)
Re.,.=12000Re...=9000Re.,.=6000Re.,.=3000Re.=1500
HJW=10
Smooth nouJe: W=7.Smm
o 4 8 xJW 12 16 20
Figure 4.15 Local heat transfer coefficient distribution
(HIW=10)
•39
• Figures 4.11 - 4.16 show local heat transfer coefficient distribution for smooth
nozzle (W=7.5mm), corresponding, respectively, to nozzle-to-plate spacings of 1.6, 4, 6,
8, 10, and 12 at Reynolds numbers ranging trom 1,500 to 12,000. Compared with those
for smooth nozzle of W=5mrn (Figures 4.3 - 4.6), it can be seen that in general the local
heat transfer coefficient displays the same trend along the impingement surface, i.e.
decreasing with xJW (distance trom stagnation point) at all HJW and Reynolds number.
While for this wider nozzle (W=7.5mrn), ooly at H/W=1.6 there exists a second
maximum at the higher Reynolds number, uolike the narrower nozzle (W=5mm), which
shows a second maximum at HIW=2.2, 6 and 9. This phenomenon implies that the
turbulence levels of the jets and the mean velocity profiles of the two nozzles must be
different, thus causing different behavior of the local heat transfer distribution.
-+.2.3 j'lusse!t iVumber Distribution
100 -------------
0-------------o 5 10~ 15 20 25
Ftgl.Ie 4.18 LocaJ and average rwssett rumer(~)
...-- ~-
_.....:.........._.. -:. -,. ....
~~~12m-'- -.:-
fOX) --- _.:
1500 -+- -:;-
Smx1h nazzfe: 'N=5nm80-
~~~13D> -.- -
8XX) --- -1!iD -+-
Sm:cth nazzIe: 'N=&rm
100 -------------•
•~o
• Figures 4. 17 and 4. 18 show the Nusselt number distribution, corresponding to the
conditions as those for Figures 4.7 and 4.9, respectively. Nusselt number, a
dimensionless parameter that reflects the extent of convection heat transfer is calculated
trom the heat transfer coefficient, characteristic length and thermal conductivity of fluid,
viz. Nu=hW!kj . In present study, the thermal conductivity of fluid (air) is almost
unchanged since the temperature of impinging air from nozzle changes only within a few
degrees (5°). If the width (or relevant width for a rough nozzle) of the nozzle is chosen as
the characteristic length, then Nusselt number only depends on heat transfer coefficient
for a given nozzle. Therefore for a given nozzie. the Nusselt number profile is exactly
the same as that of heat transfer coefficient.
100 ----------------
•90-.
40 ...:
~ Present-..:.- Siripon
-=- Van Heinlngen-=-- Gardon
2-0 slot nozzleH1W=6, Re
J=12,OOO
•
30 ----------------
o 4 8 12 16 20 24
xIWFigure 4.19 Comparison with previous work
Figure 4.19 shows a companson of the Nusselt number distribution between
present results for a smooth slot nozzle of W=7.5mm and the results of Siripon (2000),
Gardon and Akfirat (1965) and van Heiningen (1982), where the Nusselt number uses
nozzle width as the characteristic length. It can be seen from the graph that present data
and Siripon are close to each other, while bath are greater than those of Gardon and
Alcfirat and of Van Heiningen. The discrepancy may results from different boundary
·H
• conditions, nozzle designs and the turbulence intensity of the jet. Table 4.1 shows the
differences between the experimental conditions used.
Generally, impinging jet air with high turbulence level contributes most to heat
transfer and thus can produce higher heat transfer coefficients. Thermal and
hydrodynamic boundary conditions, confinement surface and nozzle design can also
affect the heat transfer processes
Table 4.1 Comparison of experimental conditions of current study and pre\-ious work
Current studr i van Heiningen Gardon and Akflrat
•
Boundary condition of Constant dissipated Constant temperature Constant heat flu.'C
impingement surface power per unit of area distribution (Heating) (Cooling)
(Cooling)
Turbulence level of air ·01. 0.6% 2.5%J /0
Confinement '{ES YES NO
Nozzle dimension 7.5 x LOO mm 6.2 x 203 mm liS x 6 inch
Nozzle aspect ratio 13 33 .+8
120 ---------------- 120 ----------------Rew Expt. Empll'1C eqn
12000 -+-9000 __
6000 --....- -3000 -+- -
HJW=6. smooth nozzle (W=6mm)1
20 -;
!:,='0 100 _Z
Rew Expt. Empinc eqn.
12000 -+-6000 ____
3000 -6
1500 -+-
HIW=12, smooth nouJe (W=5mm)80'_
Cl>
_ta 100 ~
Z
Figures 4.20 and 4.21 show the comparison of present experimental results with
the following empirica1 equation given by Martin (1977) for a single slot nozzle:•
0----------------o • 8 12 16 20 24 28
xIWFigure 4.20 Comparison of experimental data
with empiric equation (HJW=12)
o 4 8 12 16 20 24 28
xJWFigure 4.21 Comparison of experimental data
with empiric equation (HJW=6)
• _N_ll_avs_ = 1.53 Re ~XI~W. Ht':W)
Pril.l: xJ2W + Hl2W + 1.39(4.1)
where
m = 0.695 - (xJ2W -i:- (H/2W)l.J3 + 3.06)"1 (4.2)
This empirical equation is valid for 3000 < Re < 90,000,4< xJW < 50 and 4 < HIW < 20,
in which Re is based on the hydralllic diameter, i.e. twice the nozzle width as the
characteristic length. The exponent on Re, m, is dependent on the geometric variables
and it varies fram 0.56 to 0.68 in the given range of validity. Note that in above tvio
graphs the discrepancy arollnd the stagnation point (x!W=O) is large compared with that
for xIW greater than 4. The two sets of data correspond with each other reasonably weIl
'Nithin the 20~/0 range of validity~ arollnd stagnation point the difference may be over
400/0. It is not specified that Eqn. (4.1) is valid for confined or for unconfined jets.
•-+.2.-+ Comparison of'the Two Smooch iVo==les
500 ----------------
o .. 8 12 16 20 24 28xmFigure 4.22 Comparison of heat transfer between
two smooth nozzles (HIW=6)
Symmbol Rew Nozzle width........ 12000 S.Omm--- 12000 7.5mm--- 9000 5.0mm
- 9000 7.5mm-6-- 3000 S.Omm
--- 3000 7.5mm
•e.
500 ----------------
~Ne- 400 - •~ ...ë ••cg •tiSQJo(J....!lnC
~ëiiQJ~
jij(J
o-J 0 _
o 4 8 12 16 20 24 28xJW
Figure 4.23 Comparison of heat transfer betweentwo smooth nozzles (HIW=8 and 9)
Rew Nozzle width12000 S.Omm12000 7.Smm9000 S.Omm9000 7.Smm3000 S.Omm
3000 7.Smm
Symmbol......-
• From earlier sections it is known that the heat transfer coefficient profile of the
two smooth nozzles are generally sunilar except that the second maxima exist at greater
~3
• nozzle-to-plate spacings for the narrower nozzle (up to H/\V=9) than for the wider one
(below H/W=4). Figures 4.22 and 4.23 show a comparison of the local heat transfer
coefficient distributions between the two slot nOzzles with widths of 5mm and 7.5mm.
respectively.
It can be seen from Figures 4.22 and 4.23 that the heat transfer coefficients for the
narrower nozzle (W=5mm) are consistently much greater than those for the wider nozzle
(W=7.5mm): in most of the cases they are 500/0 greater at same nozzle Reynolds number
and nozzle-to-plate spacing. This great difference has been explained in tenns of the fact
that the jet from the narrower nozzle has not only a higher velocity at a given Re but aIso
possibly higher turbulence intensity due to the sudden contraction of the air flow at the
nozzle exit. both of which can improve the heat transfer on the impingement surface.
120 ---------------
0---------------
For W=5.0mm. HrN=9
For W=7.Smm, HrN=8
Symmbol Re.,.. Nozzle width~ 12000 S.Omm
- 12000 1.5mm_ 9000 S.Omm
- 9000 7.5mm--6- 3000 S.Omm
--- 3000 7.5mm
o----------~----
20 -
120 ---------------
o 4 8 12 ~6 20 :4 28xNY
Figure 4.25 Comparison of Nusselt number betweentwo smooth nozzles (HIW=8 and 9)
Symmbol Re... Nozzle width-.-.. 12000 5.0mm
- 12000 7.5mm
--- 9000 S.Omm- 9000 1.5mm
--Â- 3000 S.Omm
-.-.- 3000 1.5mm
HJW=6100 -
o 4 8 12 16 20 24 28xIW
Figure 4.24 Comparison of Nusselt number betweentwo smooth nozzles (HJW=6)
•
•Figures 4.24 and 4.25 compare the Nusselt number distributions which
correspond to Figures 4.22 and 4.23, respectively. Here it can be noted that the
difference between the Nusselt numbers for the two nozzles is much smaller than that of
heat transfer coefficient (less than 150/0). Although generally the Nusselt number for the
narrower nozzle is still greater than that for the wider nozzle in the high Reynolds
• number range tested, they are more or less the same at the lower Reynolds numbers. So
the non-dimensionalization of the heat transfer coefficient in terms of the Nusselt number
does reduce the difference and bring the convective heat transfer coefficients doser.
4.3 Effects ofNozzle-to-Plate Spacing
----- ----.--~:---:----=~--~-:-~
:-:--:-:----
4 5
HmFigure 4.27 Effect of nozzle-to-plate spacings
J--------------
JO -----------.----:- Re..=12000
Smooth noule: 'N=5mm -=- Re.=9000-..:.-- Re.=6000......r-r Re..=3000-:- Re..=1500
35 -
----:. .
-:- Re..=12000 1
Smooth "oule: 'N=5mm -=- Re..=9000-..:.-- Re.=6000~ Re.,=3000-:- Re.,=1500
IJ :2 4 S 10 12 1': '5 -S :0
HNV
Figure 4.26 Effect of nouJe-to-plate spacings
&. 200 -
=o~ 100-c:Cl~
in 0 --------------
~ 800 --------------('II
E
! 700 -
•
•
Figure 4.26 shows the effect of the nozzle-to-plate spacing on stagnation point
heat transfer coefficients. It can be seen that there exists a optimum nozzle-to-plate
spacing at a fixed Reynolds number, where the stagnation point heat transfer coefficient
achieves a maximum. The peak is more pronounced at higher Reynolds numbers. The
maximum is around 9 fer the highest Reynolds number of 12000 and reduces ta round 6
for lower Reynolds numbers. Gardon and Akfirat (1966) also reported the same trend of
shifting of the optimum H/\V.
Figure 4.27 describes the effect of nozzle-to-plate on the average Nusselt number
over the impingement plate. There aise exits an optimum where the average Nusselt
number has a maximum value at a given Reynolds number. It can be seen that the
optimum H/W is around 6 for higher Reynolds numbers (6000 - 12000) but increases to
around 9 for the lower Reynolds number (1500 -3000). This trend is opposite to that
displayed by the stagnation point heat transfer coefficient. The peaks in h (and Nu) are
45
• more pronounced at higher Reynolds numbers~ they are hardly discernible (and are within
the experimental uncertainty) for Reynolds numbers at and below 6000.
4.4 Effect of Reynolds J-.lumber
Figures 4.3 - 4.6 (for nozzle of W=5 mm) and Figures 4.11 - 4.16 (for nozzle of
W=i.5mm) clearly show that increasing Reynolds number increases the heat transfer
coefficient, as expected. Bath the Nusselt numbers at stagnation point and the average
values over the plate were correlated with jet Reynolds number and nozzle-to-plate
spacing in the most commonly used forros (r! = 0.96):
Nu = a ·Re b.(HIW) C (4.3)
Table 4.2 Correlation parameters of present and previous work
where parameter a, band c are listed in Table -+.2. which aIso include results from
literature.
• Stagnation point for
nozzle of W=5mm
Average value for
nozzle of W=5mm
a
0.10
0.0273
b
0.71
0.70
c
0.0050
0.035
Experimental conditions
HIW': 2.2 - 18
Re: 1500 - 12.000
ditto
•
! Stagnation point for ! 0.13 i 0.56 1 -0.047 WW: 1.6 - 12 1
1
1
1
1
nozzle ofW=7.5mm Re: 1500 - 12,000
1 Average value for 0.059 0.55 0.0311
ditto
1 nozzIe ofW=7.5mm11
Stagnation point 0.6 0.53 -0.11 H1W: 4 -12
by Siripon (2000) Re: 1700 - 13,600
Stagnation point by 1.2 0.58 -0.62 HJW: 14 - 60
Gardon and Akfirat Re: 2000 - 50,000
( 1966)
• From the above table it can be observed that there exists a discord between the
present results and previous work. The difference results from ditferent ranges of
experimental parameters used. The exponent on HIW is very small. Examining Figure
4.27 shows that HJW can only increase or decrease heat transfer on the two side of
optimum value of H/W. So it refers that correlating the data over different ranges of
HIW should be more reasonable. These results are shown in Table 4.3. The values of the
exponent, c, in Table 4.3 are much greater than those in Table 4.4.
•
Table 4.3 Correlation parameters for ditferent range ofHJW
a b c Range ofHIW Range of Re
Stagnation point for 0.13 0.63 0.37 2.2 and 6 1500 - 12,000
nozzle of\V=5mm
Average value for 0.31 0.78 -0.70 9, 12 and 18 1500 - 12,000
nazzle ofW=5mm
-
•
N 0 Z Z 1e W =5 mm, H IW =1 2:: Measured~ Corrlllated
80 - Nozzle W=7.Smm, H/W=1.6..:.. M easu red
----4-- Cor rI la te d70
50
20 -
40 -
30 -
90 -
60
10
100
o 2000 4000 &000 8000 10000 12000 14000Rew
Figure 4.28 Comparison between measured andcorrelated Nusseltnumber
•Figure 4.28 shows a comparison between the two sets of experimental data and
thase calculated from the above correlation equations. The maximum discrepancy is
within 150/0.
• 4.5 Heat Transfer Augmentation By Rough iVozzles
Attempts were made to increase the turbulence intensity of jet flow in order to
enhance the heat transfer rates. Three configurations of the rough-fin nozzle were tested,
which have the same open area as that of the smooth nozzle of W=5mm. Ali the tested
nozzles have the same open area, hence they can be compared at same mean jet velocity.
The Nusselt number was based on the effective width of the nozzle, which is defined by
w= (open area of the nozzle) l(length of the nozzle).
-1.5.1 Right- Triangzdar Rough iVo==/e (Jaws 1)
HIW=6
600 ----------------.- Re...=12000. smooth-:- Re...=12000, rough--6- Re",=6000, smooth-..:.- Re..,=6000. rough
- Re",=1500,smooth-=- Re..,.=1500, rough
•~ ..Gl 400 - -Ü .:... -~ -:.., .u
i""'e 500-
!
-300 ï.. .:. - .....:":
Gi ":... .-='7ii ... - ::....c .. -::.. -..;la 4 ................:: zao - .. -" ~ " ~ .....~ "~'" e:~.cu ~". .~~_~ - - -....~._. --.ii 100"~ -~3 "'~fr-.-r-ry- .
---~.~ - - -.... . . .0---------------
a 4 8 12 16 20 24 28xJW
Figure 4.30 Comparison between smooth nozzle andright-triangular rough nozzle (HIW=6)
HJW=2.2
600 --------------~-4- Re...,=12000. smooth-:- Re...,=12000. rough--4- Re.=6000. srT100th-..:.- Re",=6000. rough~ Re..,=1500.smooth
-=- Re...,=1500. nough
300 - --Gi '. --.....'7ii - .••• - .f'"
c -:...~ .....~ zoo ~o4 .:.!.!. ~.... _~ 4 -=--- ~wwGl .....~ .~~ - - ~~.-. ~ .....~ 100 -- ~~ _ ~'" ~.3 ~'r-._ ~
~ . -~-:---; ;; ~=:
o------------~--
o 4 8 12 16 ZO 24 28xJW
Figure 4.29 Comparison between smooth nozzle andright-triangular rough nozzle (HIW=2.2)
~....e 500 -
!i 400 U::8u
•
•
Figures 4.29 and 4.30 compare the local heat transfer coefficient distributions
between the smooth nozzle and the right-triangular rough nozzle with the same
equivalent relevant width (W=5mm) at HJW=2.2 and 6 and various Reynolds numbers.
From Figures 4.29 and 4.30, it can be seen that the enhar.cement of heat transfer
rate induced by the right-triangular rough nozzle i5 evident, especially around stagnation
point. For the highest Reynolds number used, i. e. 12000, the heat transfer can increase
• more than 30% for HIW=2.2 and 20% for IDW==6, respectively, while at the lowest
Reynolds number (1500), the heat transfer rate is improved even more, i. e. around 40°-10
for HIW=6. This enhancement phenomenon can be attributed to the higher turbulence
level generated by the large fins of the nozzle. Unlike the data for the smooth nozzle, a
second maximum in Nu does not exist for the case of the rough nozzle. It can be further
noticed that the enhancement effect gradually fades away with distance from the
stagnation line. The fading result may be explained from the tàct that the fins mainly
affect the turbulence level of air jet in the impingement zone. As the air spreads out from
stagnation point the turbulence intensity likely approaches the same level for the both
smooth and the rough nozzles.
400 -
0---------------o 4 8 12 16 20 24 28
xJWFigure 4.32 Comparison between smooth noozzle and
right-triangular rough nozzle (HNV=6)
ïoo ----------------- Re.,.,=12000,smooth
~ HJW=6 -:- Re..,=12000, rough"'e 600 :. -...- Re.,.,=6000, smooth~ .........r:.:,:::.- -w- Re.,.,=6000, rough
~ 500 _ --.--......~_ --- Re.,.,=1500, smoothëj _ . -=- Re...=1500, rough-= ..... ---.......-...............~8 400 -: ••• -" ..u _~. __-=--._! ---...:......:.....:-~~ :300..... -::...:...:....~ ........ ""'""-.:....:...:....2-.:...~.cu 200- ~..~ ......--....-.:.~ :'J-..---- --C-:r_l! 100 ..... -'~'J-.-~ _ ...:J:,::tD::-:.r---- _ _ _
~ ---- .2824208
HIW=2.2
4
iOO ----------------- Re.,.,=120oo, smooth-:- Re =12000 rough-...- Re:=6000, ~mooth-w- Re.,.,=6000, rough
---- Re.,.,=1500, smooth-:- Re.,.,=1500, rough
12 16xJW
Figure 4.31 Comparison between smooth noozzle andright-triangular rough nozzle (HJW=2.2)
•
•
Figures 4.31 and 4.32 show a companson between the average heat transfer
coefficients for the two nozzles corresponding to Figures 4.29 and 4.30, respectively. It
can been that in terms of the average heat transfer rate the enhancement effect is more
pronounced and exists over the whole surface. We introduce the following definition of
the enhancement factor, i.e., the ratio of average heat transfer coefficient for the rough
nozzle to that of the smooth nozzle under otherwise identical conditions
e· (4.4)
Figures 4.33 and 4.34 show the enhancement factor distribution profile along the
impingement surface corresponding to Figures 4.29 and 4.30, respectively. Both graphs
exhibit shifts in the peak values of e from the stagnation line, which indicate that the
highest enhancement effect of heat transfer occurs at around xIW=4-8. The only
exception is the case of Reynolds number of 12000 for HIW=2.2. for which the peak is
not marked. As mentiened earlier the case of the lowest Reynolds number, i. e.,1500,
exhibits the highest enhancement effect. The shifting phenomena is worth neting in that
it tells \vhich part of the impingement surface displays the most enhancement etIect on
heat transfer. lt can provide useful infonnation for practical applications when deciding
on the dimensions of the impingement surface. for example.
1.35 ----------------1.55 ----------------
---=-=-=
HlW=6
-:- Re..,=12000
-.:..- Re =6000
-:- Re =1500
, --,. --- --
Ga
~ 150 _ _=.:r-----'".ncIII _
.: 1.45 -i;iQI~
ë5 1.40 -
o -----U ~~~~:_.:! 135 - .._c ""~- -Ë :""I" _::
QI 1 .30 - -..:;u ~c -,-III ~
~
~ 1.25 -
-:- Rew=12000
~ Re...=6000
-=- Rew=1500
HJW=2.2
~ .......-'"-.-
---:--:-:-::-:
-_ -1'- - ...... ~--..'~...:.r-~-..:..
.:QI
'".n 1.30 ~- _-:; .:e:-.!:l ~-z - -:.~
J:: 1 25 _ ....::.-.;.. - - ...o ,~.. .:..::,- -'~
SUIII
:: 120 clUEQIuC~ 1.15 -cW
•
1.10 ----------------
a 4 8 12 xm 16 20 24 28
Figure 4.33 Enhancement effect of right-triangularrough noule (HI'N=2.2)
120 ----------------
o 4 8 12 16 20 24 2Bxm
Figure 4.34 Enhancement effect of right-triangularrough nozzle (Hm=6)
•Figures 4.35 and 4.36 compare the local heat transfer coefficient distribution for a
smooth nozzle (W=5mrn) and the equilateral-triangular rough nozzle with same effective
width at H1W=2.2 and 6, respectively.
50
• -1.5.2 Equilateral-Triangular Rough iVozzle (Jaws 2)
350 ~---------:---------~ Re..=12000. smooth-:- Re.212000. rough----.- Re..=6000, smooth~ Re.=6000, rough
- Re..=1500. smooth--~ Re..=1500. rough
::l~----------------
600 ---------=---------.-.- Re.=12DOO, smooth-:- Re..=12000. rough--6-- Re.=6000, smooth-ù- Re..=6000. rough
- Re..=1500, smooth-=- Re.=1500, rough
26::0~ 6xfW
Comparison between smooth noule andequilateral-triangular rough noule(HJW=6)
Figure 4.36
::8:::0~2 ~6
xIWComparison between smooth nozzle andequilateral-triangular rough nozzle(HNJ=2.2)
Figure 4.35
•600 -----------------
-..... Re.=12000. smooth':1:. HIW=6 -:- Re.=12000, rough
"'E seo .:.,..... -6- Re.=6000, smooth~ ~-r~ ~ Re.=6000. rough
i -...;--:...-~ - Re.=1500, smooth,,_, 4U0 _ •• ...r....> -=- Re.=1500. rough
;g • "'-:-:-:.:.- -'; .... --r
~ 300 :-.:..:...:.::.~... ...-..:-..:-.:-...:-..=--: _';i ....... ~._
~ A...~~.:...:-.:...~•.200- ~
i ~QI ::.-
~ 100~••*••.•------_- _~ .. .-.-.-
H/W=2.2
-:co ----------=------------+- Re..=1200a. smooth-:- Re..=12000. rough----.- Re.=6000, smooth---&-- Re.=6000. rough
- Re.=1500. smooth-=- Re.=1500. rough
ë:.!!!.g JGO -
âio(J
",':1:.
.ê 5eo -~
0-----------------
•
o 4 8 12 16 20 24 28x1W
Figure 4.37 Comparison between smooth noozzle andequllateral-trlangular rough noule(HIW=2.2)
a 4 8 12 16 20 24 28xNI
Figure 4.38 Comparison between smooth noozzle andequilateral-triangular rough nozzle(HIW=6)
5L
• It is shown that Jaws 2 nozzle produces similar trends for the heat transfer rates
and their distribution as that of Jaws 1 nozzle, although the magnitude of enhancement is
not as great as that shawn by Jaws 1. Figures 4.37 and 4.38 show the corresponding
comparison of the average heat transfer rate between the (Wo nozzles. These two show
the enhancement effect of equilateral-triangular rough nozzle is significantly smaller than
that of right-triangular nozzle (Figures 4.31 and 4.32). This fact suggests that turbulence
level of the air jet produced by Jaws 2 is smaller than that of Jaws l, thus resu1ting in
lower heat transfer rates.
Accordingly Figures 4.39 and 4.40 show the enhancement factor profile along
impingement surface, corresponding to Figure 4.31 and 4.32. respectively.
~:: --------------- ':2 ---------------
HJW=2.2
-:- Rew=12000
-..:,.- Re..=6000
-::- Re.=1500
; , 12 - ... :.:.,....
~ -~ 1 ~o .;-=r~-:.:~
c:w , ca -
-:- Rew =12000
---ü- Re.=6000
-=- Re.=1S00
-; =-"-r-~_
~ , '6 - -
a.9 , '4 U~
~ ~ ~2 -:: ::.,::":'.:.: _
E ,-:::--': - , :QI ............ - ... -u 1 • 0,-- - -, --
l ,ca ~- -' -'-~~~cp-~
•~ :'5 ---------------
Figure 4.39 Enhancement etfect of equllateral-trlangularrough nozzle (HJW=2.2)
a 4 '2xIW 16 20 ::4 28
1 Q€ ---------------
a 4 1::'x/W'6::0 24 28
Figure 4.40 Enhancement effect of equilateral-trfangularrough nozzle (HJW=6)
•
Similar to Figures 4.33 and 4.34 for right-triangular nozzle. Figures 4.39 and 4.40
show without exception shift of the peaks tram stagnation point. Figure 4.39 shows the
three peaks for three Reynolds number occurs nearly at the same location, i. e. H/W==6.
The case of the lowest Reynolds number, i. e.,1500, also exhibits the highest
enhancement effect, i.e.17% for H/W=2.2 and 20% for HIW==6. For the case of H;W==6,
the higher the Reynolds number, the doser the peaks are to the stagnation point.
52
• Compared with Jaws 1 it is clear that the enhancement factors of Jaws 1 are generally
only about half of the former or even less under the same conditions.
Jaws 2 nozzle was also tested by shifting one side of the nozzle plate one half of
the spacing between two consecutive fins, i. e. 2mrn. This was done in order ta see the
effect of changing the flow structure (velocity profile) of the air jet out of the nozz1e on
heat transfer. Figures 4.41 shows the comparison of the two jaw-arrangement.
(a) Normal jaw-arrangement(tips to tips)
(b) Shifted jal"-arrangement(tips to \'aJle)'s)
Figure 4.41 Sketches of nozzle jaw arrangements
0---------------:8242012 16
xJW
(b) HIW=6
oo--------~------
EOO --------=------=-~----:-~--:-------- Re.=12000. shifted
~ HIW=6 -:- Re.=12000. rough"'e ---6- Re.=6000.shifted- 5ca .~ ..r........ --..:.- Re.=6000. rough~ ~'-, --- Rp =1500.shiftedë .-- -.q) • .:...-. -=- Re =1500 rough1j 400 - •• ...,.~ .'
~ ..~.--rQ -..;:.-
~ 3eo _. ....~-.--..r-~ -~=~';j :00 -q)
::
~ ~.... .-!lS '00 - .- .iü . . . __~ ...._---.:-.:----=-=
2820 :;4,:; ~ 6xJW
(a) HIW=2.2
a<1a
seo -------~-------~ Re.=12000. shifted
HIW=2.2 -:- Re.=12000. rough-..- Re.=6000. shifted-.:.- Re.=6000. rough
--- Re.=1500. shifted-=- Re.=1500. rough
•
Figure 4.42 Comparisons between shifted nozzIe and normal nozzle (for Jaws 2)
•Figure 4.42 show the comparisons between the shifted nozzle and normal nozzle
(for Jaws 2). It can been seen that heat transfer rates for shifted Jaws 2 nozzle are slightly
lower than those for Jaws 2. The maximum difference is below 80/0 for the cases studied.
The decrease of heat transfer rates should also be explained as that the turbulence levei is
53
•
•
probably lower for shifted nozzle since they both have the same open area, hence the
same mean velocity at the exits of the nozzles.
.J. 5. 3 Rectangular Rough Nozzle (Jaws 3)
Figures 4.43 and 4.44 show the comparison of local heat transfer coefficient
distribution between smooth (W=5mm) and rectangular rough nozzle with same
equivalent width at H/W=1.2 and 6, respectively. It is shown that unlike Jaws 1 and 2,
Jaws 3 produced negative effect on the heat transfer rates compared with the smooth
nozzle, though the negative effect is only pronounced at higher Reynolds number. The
diITÜnishing effect is up to 12~'O at a Reynolds nurnber of 12,000 and below 7~o at a
Reynolds number of 1500. This different effect is attributed to lower turbulence level
induced by the rectangular teeth at the nozzle exit, which needs verified by measuring
flaw structure of the jet. The three dimensionality of the jet turbulence structure may also
be a contributory factor.
H1W=2.2
•
:c-a ----------------+-- Re.=12000.smooth-:- Re.=12000. rough.-.- Re.=6000, smooth-.:.- Re.=6000, rough- Re..=1500, smooth-=- Re..=1500. rough
o----.....,.....------~----
o 4 8 12 16 20 24 28xm
Figure 4.43 Comparison between smooth nozzle andrectangular rough nozzle (HI'N=2.2)
5.+
:ca ~---------------+-- Re =12000. smooth
..~ • HIW=6 -:- Re:=12000. rough.E ~ ~ Re.=6000. smooth~ olCO _ ---:;.- • ---..r- Re.=6000. roughë .---~~. -- Re.=1S00. smoothal -:.!... -:- Re.=1500. rough~ .._.~~<ü 300 - --~8 ,... ~~... -'t-.'&'al ~ - __
~ ~~ -~ 200 -
~ ~al
:;100 .~Qi -'~O"'~~ - ~n:~=t:.=/"'....-........J"__
- - - - -0---------------
o 4 8 12 16 20 24 28x/W
Figure 4.44 Comparison between smooth nozzle andrectangular rough noule (HJW=6)
• 4.6 Effects ofInserting Turbulators in Jet Flolv
•
Three types of turbulence generators, square rad, strip, and a perforated plate
were used to study the effect of inserting turbulence generator in the jet tlow field. The
turbulence generator (only ane is used each time), fixed rigidly between the nozzle and
the impingement plate at variaus locations, is expected ta affect the jet tlow, producing
large-scale vortex wakes with high turbulence kinetic energy, and thus influence the
impingement heat transfer rate. Although heat transfer rates were measured main1y on
one side of the stagnation line due to symmetry, single and multiple insertions were
employed in each experiment to ensure symmetry. Insertions were placed symmetrically
above the jet axis although heat transfer measurements were made on one side ooly. The
location of the turbulator is represented by L/\V and x!W, the ratio of its distance from
impingement plate to nozzle width and of the distance tram the stagnation line to the
nozzIe width, respectively (see Figure 3.4). Ooly the smooth ofwidth 7.5mm was used
in trus study of the effect of turbulator on Nusselt number.
.J. 6.1 Single Insertion ofSquare Rod or Strip
9)---------- 8)----------- 4)-----------....- ~ ldUiItr-:- l.N*1. JIN(J~.r IINQ, JIN(J
-+ 1JW:1.1fH!l
o5-----------
....- N:llLltUim'-:- 1.JN=1. '1fN:(J
-..:.- l.JN:!3. '1fN:(J
-+ 1.JN=1.1ftN:!.2.
15----~------
o7l~---------
o
• Figure 4.45 Effect of inserting a rod turbulator in jet flow at ditferent Reynolds numbers
55
• Figure 4.45 shows the etfect of inserting the rod turbulator in the jet f10w at
Reynolds numbers of 12,000, 6000 and 3000. It can be seen that the insertion of single
rcd significantly changes the heat transfer profile. The insertion does not enhance the
heat transfer rate as expected but in fact decreases heat transfer. The insertion aIso shifts
the Nu maxima in most cases. In the case of LJW=l and xJW=2, the heat transfer was
decreased by nearly 500/0 at locations far tram the stagnation line. The shifted maxima
occur at xJW ranging from 2 to 6 depending on the Reynolds numbers and insertion
locations.
•
90----------------.- r..D 1LItJJatr-:- LNI=1.~
~ l.fN:2.~
-.- LW:1. JNtI:!2.
3)-
90----------------.- No turtUatrr-:- LJ\N=1, JN.I::(J
-.:.- l.NI=2., JN.I::(J
-.- LJ\N=1, 7f\N=2
~Fè=lml
164 8 JNJ 12
(b)~=6DJ
10--------~-----
o8 ~ 12 16
(a)~=12,aD
4
10 --------------
o
Figure 4.46 Effect of strip turbulator and its inserting locations
•Figure 4.46 shows the effeet of a strip turbulator in the jet flow at Reynolds
numbers of 12,000 and 6000. It can be seen that the shapes of the Nu curves are similar
ta those obtained for a rod turbulator.
56
• In order to see the difference clearly, the effects of the !Wo types of turbulators are
plotted in one graph as shown in Figure 4.47. It can be seen that the difference is very
minor, almost indistinguishable. This phenomena suggests that the two turbulators
employed affect the flow structure in nearly the same way, which therefore affect the heat
transfer nearly identically.
10----------a
JJ~----------
aJJ-----------
a•Figure 4.47 Effect of turbulator types at different locations
.J. 6.2 A1ultiple Insertion ofSquare Rods or Stnps
•
In this case the twc identical rcds or strips were placed symmetrically on both
sides of the stagnation line to ensure the symmetry of the flow structure generated.
Figure 4.48 shows the effects of turbulator types and insertion locations on the
average Nusselt number at a Reynolds number of 12,000 and HJW=4. It can be seen that
bath rcds and strips shift the heat transfer maxima away from the stagnation line to x/W
of 3-4. The shifted peaks induced by turbulator insenion are all higher than those without
Si
• insertion. There are actually no significant change of heat transfer rate at locations far
more than xIW =7.
~Zl---------~ 110---------- 1:;D----------
'6
-+- ~ tl6t:Uâ:I"-)- Fba~ 9rips
h..rtilll k:1c;6n L.!N=1.~!
~----------J
"0 -
1(l) -
'6:tl----------
(]
~ '''.iBIBI--:
Figure 4.48 Effect ofturbulator types and inserting locations (Rew=12,OOO)
"0 -
•
•
From Figure 4.48 it can he estimated that in case Ca), i.e. LIW=l and x/W=ü, the
peak enhancement for rods insertion is about l :2~/o and for strips insertion is only about
6~/0. While for case Cb), i.e. LAV=:2 and xfW=O, where the insertion is farther away from
the impingement plate, the peak enhancement for rods is over 27% and for strips
insertion is up to 230/0. For the case of insertion shifted away trom the stagnation line,
(c), i. e. LIW==l and x1W=2, the peak enhancement for rads and strip is negligible. So it
shows that case (b), i. e. LJW=2 and xJ\V=O, produced the highest heat transfer
enhancement (around 20%-300/0). It is better than the case of (a), i.e. LIW=l and x/W=O,
the doser bcation of the insertion to the irnpingement plate, in which heat transfer
enhancement is only around 100/0. These results suggest that insertion of a turbulator
doser to the nozzle would help enhance the heat transfer to a greater extent.
Comparing single rod insertion and multiple rod insertions it was seen that the
resultant heat transfer effects are very different: single rad insertion produces negative
heat transfer etfect while multiple insertion promotes heat transfer significantly if the
58
• insertion location is appropriate. This contrary effect of insertion suggests the
importance of location of the turbulator for optimum heat transfer enhancement.
It should be pointed out that the stagnation point heat transfer rate is deteriorated
in all cases of turbulator insertion studied. This is attributed to the "shadow" effect of the
turbulator. The jet flow appears to be detlated by the turbulator away from the stagnation
line. From the above graphs it can be conc1uded that turbulator types, whether rcd or
strip make no much difference concerning the enhancement of heat transfer.
100 -
~20 -----------------.- Re..=12,OOO, no turbulator-:- Re.=12.000, with rods-.- Re=GOOO. no turbulator-..:,...... Re.=6000, with rads
---- Re.=1500. no turbulator~~ 30' : _ -=- Re.=1500, with rods
~ :- .;;':':' HIW--4, UW=2, xJw--O
11160-.:..:.~Z~ ~&. --"'-..r---.--. - ----t40~ ...~:=ct...,::.... ~ -
:0 ~ @;o:CD:Ij~: enh:n:ement
----------~-:-~
~co -----.-------R.-e..,-=......11r'"o=oo=-.n-o--"t-uf'b""---'ul""""at-or--_ -:- Re..,=12,OOO. with rads
_ ---6- Re=6000, no turbulatorBQ'" ~_ ........:.- Re..,=6000, with radsCi -.. -:_ ---- Re..,=1500. no turbulator~ e. -_.;:- Re..,=1500, with rods
Ë '50- • .:
~ - ~..:..~ " .~. .......-~ - --.;... --:Z .:0- ~
i -------~.:.-.:.Ci:-<•
axIW
:0Q---------------
:0
(a) HJW=4, LJW=1, xfN=O
Figure 4.49 Effeet of Reynolds number on heat transfer with rcds (H;W=4)
Figure 4.49 shows the effect of jet Reynolds number on the average Nusselt
number. The data show that the peak enhancement occuring at lower Reynolds numbers
of 6000 and 1500 are quite similar ta those at the higher Reynolds number of 12,000. It
suggests that insertion can enhance heat transfer even at lower Reynolds number.
•59
•
•
-1.6.3. Insertion ofa Perforated Plate
-1.6.3.1 Effect ofNormal Distance
The etfects of insertion of a finite-size perforated plate between the slot nozzle
and impingement plate were investigated by inserting the plate at pre-selected locations
between the nozzle and the impingement plate. These results are shown in Figure 4.50
where LIW is the ratio of the distance between the plate and the impingement plate ta the
width of the sIal nozzle, and xf\V, which is the ratio of the horizontal distance between
the left edge of the perforated plate and the stagnation point ta the width of the slot
nozzle xJW=-5.2 is calculated from the perforated plate is fixed symmetrically across the
nozzle, 50 half the width of the perforated plate (39 mm) divided by 7.5 mm produces
5.2. where negative sign means the left edge of the perforated plate is on the other side of
the stagnation line.
350 -----------------e-- Without Insertion
--- LJIN=8,x1W=-5.2-...- LJIN=5, x1W=-5.2.CC--......- LJIN=2,xJW=-5.2.CC
-c:::,S!.~ 250 =oQIU~
.; 200 -c:::~~--~QI
.J::.
'ic.Jo.J
• ,......•
HIW=10, Re=12,592
•
100 ----------------
a 1 2 X1W3 4 5 6
Figure 4.50 Effects of insertion abject and its locations
It can be seen clearly from the figure that sorne of the curves deviate significantly
near the stagnation plane (XJW=O), while there is no much difference at locations far
60
• fram the stagnation point (XIW > 3-4). The phenomenon that the etfects rnainly Decur
near the stagnation point is rnainly because that the characteristics of jet tlow are affected
greatly near SP by insertion of the plate. AIso the perforated plate is located
symmetrically on both side of the stagnation point, then its effect is restricted aIso
therein. Figure 4.51 shows a schematie of the jet flow affeeted by the insertion of the
plate.
Confinement plate Nazz!e
•
•
Imptngement surface
Figure 4.51 Schematic ofjet flow affected by the insertion
It is ev;dent that the insertion affects h more when it is doser ta the impingement
plate, i. e., LJW=2, while those of LIW=5 and L/\V=8 are very dose to each ather and
both are even smaller than thase without insertion near the stagnation point. This cao be
explained possibility by that the etfect of insertion of the perforated plate diminishes after
the tlow leaves the perforated plate for the impingement plate. The decrease of h of
LIW=5 and LIW=8 campared with those without insertion is probably caused by the
decrease of the normal impinging velocity of the flow on the impinging plate due to
insertion. However, in the case of LIW=2, where the perforated plate is located very
close to the impingernent plate, the insertion intensifies the turbulence level of flow and
increases h compared with those without insertion. More experiments are expected to be
condueted in order to validate the above speculated explanation. The open area and
dimensions of the perforated plate are clearly significant parameters affecting the flow
and heat transfer rates.
61
•
•
-+.6.3.2 Effect ofHorizontal Distance
Figure 4.52 shows the effect of horizontal location of the perfcrated plate on heat
transfer coefficients h when the impingement plate is placed doser ta the nozzle, Î. e
H!W=4 and Re=12,OOO. Comparing the two curves for x1W=O and x/W=-5.2 (both at
LIW=2), it can be seen that near the stagnation point, the tirst set of h is higher than the
second one, i. e. the horizontal shift of the plate from the stagnation point enhances heat
transfer. The difference between the !Wo curves of L/\V= 1 and LIW=2 (bath correspond
ta CC, i. e., the plate is lacated symmetrically on the bath side of the stagnation point),
further confinns the above results that the doser the normal distance of the perforated
plate ta the impingement plate, the higher the heat transfer coefficients near the
stagnation point. Since this study was rather limited in scope no graund conclusions can
be drawn regarding the effect of the perforated plate as turbulator on impingement heat
transfer.
400-----------------
~ 350 ....:..N
E~ 300 ,- "
-:- Normal (without plate!-..:.- UW=1,CC-=- LlW=2,xJW=-5.2{CC)
--- LlW=2,xlW=O
HIW:4, Re=12,OOO
•50 -----------------
02345 6XNI
Figure 4.52 Effects of the plate and its locations
62
• 4.7 Comparison with Simulation Results
The present experimental data were compared with bath the RSM and standard k
g model the turbulence model simulation results, bath obtained by Shi Yuling (from
Dept. of Chem. Eng., National University of Singapore) using the commercial software
for computational fluid dynamics, FLUE~l (version 5.0). Appropriate grid
independence tests were made ta ensure numerical validity of the simulations.
Figures 4.53 - 4.57 show the comparison between the CFD model results and the
present data for the smooth nozzle with W=7.5mm and HIW = 12, for 5 jet Reynolds
numbers ranging from 1500 - 12,000. It is noteworthy that the predicted heat transfer
coefficients deviate very significantly from the measured values in this work even for the
smooth slot nozzle. In general the deviations are greater in the impingement zone and
they increase \\'ith the jet Reynolds number. .~ide from the weIl recognized limitations
of the turbulence models tested for such a flow. one possible problem with the large
deviations in the stagnation region may be due to the physical model used to calculate the
• heat transfer coefficient in the region where the mean tlow is almost normal to the
surface rather than paraIlel to it. Clearly, better and more advanced turbulence models
need to be developed to predict the complex impingement flow region weB. Further, the
heat transfer coefficient is calculated from the temperature gradient at the surface- the
process of differentiation amplifies minor errors resulting in poor predictions. It is
expected that the comparisons with statie pressure distributions, mean velocity profiles
and temperature distributions in the flow-field will be much more favorable .
•63
••
Figure 4.53 Comparison with RS~{ model and standard k-E model(Rew=11.000, heat flux = 3490 W/m:)
25201510
Xl\V
5a
-~ 300... 1
e -- standard k-E model~- 250 .. RSM model~ • • experimenu! data-~'ë:j •t:: 200-:~..
...:!:.Il 150-~--:1:.a 100~ÇJ
~~
50
•
..... -;-:..~.~._-
-- standard k-E modelRS~1 mode!
• experimental data
-:.c:: 220... 1- ..- 200~ •\ri 180--~
'ü 160t::-~ 1..0Q
ÇJ
~
..;: 120fi:!-~ 100-= 80.a~
60e."=-~== ..0iJJ
Figure 4.54 Comparison with RSM model and standard k-E model(Rew =9000, heat flux = 2827 W/m2
)•o 5 10
XJW
15 20
•
~o ~--~----:----------r----"----~
25201510
-- standard k-E modelR5M model
e experimental data
5
•••
o
..e.
60
80
100
-~ 160 ~-------------------.,
MI....::::~ 140
•XJW
Figure 4.55 Compar1son \Vith RSN( mode! and standard k-e model(Rew=6000. heat flux := 1963 W/m2
)
Figure 4.56 Comparison with RSM model and standard k-e model(Rew=3000. heat flux = 1256 W/m1
)
201510
XJW
-- standard k·E mode1RS~t modd
e experimental data
5o
...
•65
Figure 4.57 Comparison with RSM model and standard k-E model(Rew =1500. heat flux = 872 W/m':.)
15201510
XJW
5o
• -~ 70.... 1,.= -- standard k·E model-~ 60 RSMmodel-~
experimental data- •=a...~ 50tE~=!;oj 40~
=~ 30..--~~ 20a..~
~ ID~
•
•
It can be seen that for most of jet Reynolds numbers the present data are about
20-30% higher than the other t\vo sets around the stagnation point except for the loVY'est
Reynolds number of 1500. The difference decreases with increasing the distance from
the stagnation point (x/W) and almast disappears far from the stagnation line i.e. in the
wall jet region. At the Iowest Reynolds number (1500) tested, the experimental and
simulation results are in close agreement. There is not much difference between the
predictions of the two turbulence models tested either. This agreement is much closer
than what several authors have reported in earlier literature when comparing different
turbulence madels with experimental data. There is much disagreement between the data
of different experimenters as weil. Also, the boundary conditions are often not fully
specified in the published papers 50 that the CFD model predictions must use
approximate values for such boundary conditions as the nazzle exit velocity profile and
turbulence kinetic energy distribution. It is interesting ta note that Shi Yuling has alsa
66
•
confirmed via numerical experiments with Fluent 5.0 that the local Nusselt numbers
ca1culated for isothermal impingement surface are almost the same as those for a constant
wall heat flux boundary condition provided the physical properties of the fluid are
constant. This is an important result since ditferent researchers have used difTerent
thermal boundary conditions (ranging from isotherrnal to constant heat flux ta variable
temperature along the impingement surface). Thus, the differences reported 10 the
experimental literature on impinging jets cannot be attributed to differences 10 the
thermal boundary condition.
4.8 POlver Consumptionfor Nozzles
As discussed earlier Jaws 1 (the right-triangular rough-edged nozzle) was found
to have the mast enhancement etTect on impingement heat transfer of ail nozzles tested.
ln view of the fact that power consumption related to the pressure drop across the nozzle,
however, it is useful to compare the pressure drop and nozzle discharge coefficients for
the nozzles.
According to. White (1986), the discharge coefficient (Cd) of a nozzle can be
ca1culated from the following equation:
The nozzle discharge coefficient can vary from 0 ta 1. A value close ta 1
indicates the least tlow contraction of the exit jet due ta the nozzle. which is most
desirable. More tlow contraction (Iower Cd) consumes more energy (pumping power) at
a gi ven flow rate.
Pumping power is defined as the power needed by the blower ta overcome the
flow resistance of the whole system. In order to compare the difTerence caused by the
different configurations of the nozzles. Ta better describe the heat transfer behavior, a
parameter, o., called unit pumping power is introduced:
•cr = ~P/h
67
(4.5)
(4.6)
• where LlP stands for the pressure drop across the nozzle and h is the average heat transfer
coefficient measured under the same operating condition. Physically a represents the
power consumed on the nozzle for one unit of heat transfer under unit temperature
difference, 50 it can be used to compare different nozzles.
0.5 ----------------
-•
12000
HJW=6. W=5mm
9000
Smooth noule
• Jaws 1- Jaws 2
- Jaws 3
6000Rew
Figure 4.59 Comparison of pumping power
o 2000 4000 6000 8000 10000 12000 140C
Rew
Figure 4.58 Nozzle discharge coefficients
0.95 2.5-.- Smooth "ozzle •- Jaws 1 /'/
~- ~ Jaws 2 ./'C 0.90 -2- Jaws 3 -~.. :li::
----- N- - EC .../
~ 2.0 -cu ~'ëj 0.85 - E::: ...cu
.~cu
0 • ~(,J0.80 - --- 0
cu ~ 1.5-C) • .~ C)...!! cÜ
/' ·a.~
O.iS - .::
.-----------= §"cu _. ------- .= 1.0-N .------ -----N 0.70 - .,.-
~--------'c
0 'r ~Z --
0.65•
•
Figure 4.58 shows the discharge coefficients for three rough nozzles and the
smooth nozzle (with the same equivalent width) for Reynolds numbers from 1500 to
12,000. It can be seen that aU the four nozzle discharge coefficients increase as Reynolds
numbers increase, indicating that flow contraction decreases with increasing flow rate
It can be seen from Figure 4,58 that of aH four nozzles studied, the smooth nozzle
has the highest values (the closest to 1) of Cd for any given Reynolds number, Jaws 2 the
second, Jaws 3 the third and Jaws 1 the lowest. This result implies that although Jaws 1
can produce highest heat transfer coefficient for a given condition, it causes more flow
contraction and thus consumes more energy for the same heat transfer performance.
Figure 4.59 shows a comparison between the pumping power for the four nozzles
tested at Reynolds numbers of 6000, 9000 and 12,000. It can be seen that compared with
the smooth nozzle, Jaws 1 cornes with a little higher (Jess than 1O~/o) unit pumping power,
•
•
•
Jaws 2 has more or less pumping power requirement depending on the jet Reynolds
numbers, and Jaws 3 has evidently higher unit pumping power. Considering that Jaws 1
enhances heat transfer the most (over 40°,/0), Jaws 2 the second (around 15~/o), and Jaws 3
worsens the heat transfer, it can be concluded that in view of heat transfer enhancement,
as weil as power consumption, Jaws 3 is not a recommended design for a rough nozzle,
Jaws l is recommended, while Jaws 2 may be adopted in sorne cases.
4.9 Closure
An extensive experimentai study of the heat transfer behavior of confined slot jet
impingement on a planar (flat) surface were conducted. using both smooth and rough
nozzles and with insertion of various turbulators. The parametric etfects of jet Reynolds
numbers and dimensionless nozzle-to-plate spacing on the heat transfer characteristics of
the heated target surface were explored using bath smooth and tV.lo-dimensional jawed
slot nozzles of novel design proposed by Nlujumdar (1998).
The results show that the two triangular rough nozzles, 1. e. right-triangular and
equilateral (Jaws land Jaws2) can enhance impingement heat transfer up ta 40~·~ while
the rectangular (Jaws3) rough nozzle actually decreases the heat transfer rates. The
enhancement of heat transfer induced by insertion of turbulators depends on the type and
location of the turbulators employed.
69
• IICHAPTER 511
HEAT TRANSFER CHARACTERISTICSNON-PLANAR SURFACE
5.1 Introduction
•
•
This chapter includes results of an experimental study on the heat transfer
behavior of an impingement surface with a semi-cylindrical pedestal. Here "non-planar""
refers ta the impingement surface with a semi-cylinder attached on top of the flat plate.
\Vhen jet impingement cooling is applied to a curved surface such as the outer
surface of a cylinder, the curvature effect needs to be taken into consideration. Gau and
Chung (1991) reported that for impingement on a convex surface, three-dimensional
counterrotating vortices in the stagnation region can affect the local heat transfer. It is
expected that the appearance of counterrotating 'lortices can increase the momentum and
energy exchange in the flow, which can enhance the heat transfer along the wall.
5.2 Thermocouple Arrangement
Figure 5. 1 shows the schematic of the non-planar surface used with the
thermocouple arrangement on a semi-cylinder surface and the downstream flat plate ta
farm the test surface.
A semi-cylinder with a diameter of 64cm was attached on a flat plate to form the
test surface. It is located at the geometric center of the flat base plate. The cylinder is
attached firmly to the flat plate with a double-sided adhesive tape. About 30
thermocouples were inserted between the stainless steel foil layer and the double-sided
adhesive tape layer.
Most of the thermocouples were installed on one side (the right half side) of the
stagnation line due to the symmetry of the flow pattern, except that a few were also
placed on the left half side of the stagnation line (such as position C-l) that were used to
validate the flow symmetry.
70
• +Jet
--""'1~ lil __
1B
Semi-Cylinder
/'Flat Plate
o
Confinement Surface
Angle between adjacento
thermocouples: 6
6 i~.~ 6.5 xJ\V
•
•
Figure 5.1 Schematic of the thennocouple arrangement on a semi-cylinder surfaceand the flat plate region
On the right haIf side of the semi-cylinder. 14 thermocouples were instaIIed at
equaI spacing, dividing the right half side into 15 identical sections. The central angle
corresponding to any two consecutive points is 6°, i.e. 1115 of900. The arc length of each
section is about 3.35mm. The thermocouple positions on the semi-cylindricaI surface are
designated as Ci (i = -1,0,1,2.3 ···13,14). The exact temperature at thejunction point
of the semi-cylinder and the flat plate was difficult to measure accurately and hence was
not used in the calculation of the local heat transfer coefficients. Additional
thermocouples were instaIIed on the flat plate region, in the range of xJW=5.5 to 16,
designated as Pi (i = 5.5, 6, 6.5, 7, ···15, 16), as shown in Figure 5.1. ln the figure Hl is
the distance between the nozzIe and the top of the semi-cylinder while H is the spacing
between the flat confinement surface and the flat portion of the test surface.
The values on the x.JW axis in Figure 5.1 are for a nozzle width of 7. 5mm. In
order to plot the heat transfer coefficient distribution for both the curved and flat surfaces
71
•
•
•
on one graph~ we define x!W for the curved surface in the same way as that for flat
surface. The radius of the cylinder is 32mrn. The xJW locations of the thennocouples on
the curved surface are 0, 0.284, 0.568, 0.852 .... .4.27 (the junction point), and for flat
surface are 5.5, 6, 7, .... 15 and 16. Here x!W for points on the semi-cylindrical surface is
the distance between Ci and stagnation line (e.g. the distance x4 for point C4, as shawn in
Figure 5. 1). which is the same as those for flat surface.
It is worth mentioning that the nozzle-to-plate spacmgs defined here for the
curved surface study remain the same as those for IIF flat surface~ i. e. the distance
between the nozzIe and the flat surface. Actually the HJW for the stagnation line. i.e. the
top line of the cylinder. is much smaller than that for the flat portion of the impingernent
plate.
5.3. Temperature Distribution along the Surfaces
Figures 5.2 and 5.3 display the measured irnpingement surface (of both cylinder
surface and flat surface) temperature distributions obtained for a constant heat flux of
4181 \V/m2 for the smooth nozzle (W=7.5mm) and for both smooth and rough nozzles of
W=5mm. respectively. Higher wall temperatures correspond to lower heat transfer rates,
as discussed in the previous chapter. Both figures show similar temperature distributions.
On the curv"ed surface the stagnation point (the top of the semi-cylinder where x/W=O)
temperature is the lowest. The temperature increases with x/W, reaching a peak in the
low half of the semi-cylindrical surface (near the junction point), and then falls. While
on the flat surface. the temperature generally increases with xJW, except for sorne smail
bumps. In both figures the temperature profiles on the curved surface are difficult to
distinguish. while those on the flat surface are distinct. Figure 5.2 shows that the
temperature on the flat surface increases with HIW, while Figure 5.3 shows that the
temperatures for the smoath nozzle is lawer than thase for the two rough nozzIes,
indicating a higher heat transfer coefficient for the smooth nozzle which was a counter
intuitive observation
72
•
HIW=6Tj=26.1 0C
HIW=8T.=26.30C
J
HIW=12Tj=22.60C
Flatsurface
.. ' 1..:. ~=.-~.
:ci••
4.3
Curvedsurface
-------:.~
~- ."::&:~-~-.:=--
CIft40 -
70 -
eu~
~ 60 -...ca~
eu~
E 50 eu~
Smooth nozzle: W=7.5mm80 ~ Heat Flux: 4181W/m2
Rew=12,OOO
-(Jo-•
30 -------------------
-2 0 2 4 6 8 10 12 14 16 18xJW
Figure 5.2 Temperature distribution for smooth nozzle (W=7.5mm)
Heat Flux: 4181W/m2
80 -Rew=12,OOO
HIW=9, W=5mm
•70 -
l)o-Q,)~
~ 60 -...ca~
Q,)c.E 50 Q,)
~
40 ~
Curvedsurface
Flatsurface
RectangularT.= 28.0oC
JEqui-triangularT=25.4oC
JSmoothT.=26.20C .../e
J /'- -"""'" -
•",'
306.4
•-2 0 2 4 6 8 10 12 14 16 18 20 22 24 26
xlWFigure 5.3 Temperature distribution for smooth and
rough nozzles (W=5mm)
73
•
•
•
70 -------------------Smooth nozzle: W=7.5mmHeat Flux: 4181W/m2
60 - Rew=12,OOO~--
Curved Flat ".' -50 - ::: .
surface surface -- .......(J - ><::::: -0- "' •- 40 - -41
~- - •1 - ~
~~~.. - : .. - - -...---.-.~.
.,HIW=630 - - ..T.= 26.1 0C- i~.~ - J
.r-
~ .=. HIW=8-.-. T.=26.30C20 - ,"T',.O;. -; JQ - HIW=12• 4.3 T.=22.60C
J10
-2 0 2 4 6 8 10 12 14 16 18xJW
Figure 5.4 Temperature difference profilefor smooth nozzle (W=7.5mm)
Figure 5.4, which corresponds ta Figure 5.2. shows the profiles of driving forces
I.e. temperature difference between the impingement surface and the jet, of the
impingement convective heat transfer. Similar ta Figure 5.2, the three curves are
distinctly different on the flat surface but not on the curved surface. It is expected that
the thermal history affects the heat transfer on the flat portion of the impingement surface
and a recirculation region is expected between the curved surface and the flat plate which
can affect the heat transfer for the flat portion of the impingement surface.
7~
• 5.4. Heat Transfer for Smooth Nozzles
5. -1.1 Heat Transfer Coefficient Distribution
161412
-:- fèK=120001-=- fè.y-""9lOO-..:r- fè.v=6OOO---':'--- Fè.N=3000-:- Aa..f=1500
106 8'JN.J
Rasurface
42
Q.arvedsurface
Figure 5.6 Local heat transfer cœfticient distribution(t-WP8)
161412
-;- JèK=12DOO1
-=- AIwI9000--ti- AIwI9000--7-~
-:-~sœ
10
----....:_-........... ""' .....
42
Q.!M!Idsurface
6~ 8
Figure 5.5 L..ccaI heat transfer ooeffident disbibution
(~)
-...;: 4..3o----..:.--=~-----------
o
i"DJ ,-N
.Ë '!250 -:
•
161412
-:- R2yr12000
-:- Reyr9000
--.:r- ~ooo
--;- Reyf=3000
-: - fèw=1500
10
Ratsurface
4
5mooth nome: W=7.SrTIT1
HJW=12
2
eurvedsurface
6xm 8
Figure 5.8 Local heat transfer coefficient distribution(HI'N=12)
14 1612
-:- R2yr12000.-=- R2w--gooo~~
-":'- Reyr3000
-:- Pew=1500
10
HIW=10
Ratsurface
4
5m::loth nozzfe: VF7.5mn
2
Curvedsurface
o 6xN1
B
FlQure 5.7 Local heat transfer coefficient distribution
(J-UW=10)
•75
•
•
Figùres 5.5 - 5.8 show local heat transfer coefficient distribution curves for the
smooth nozzle (W=7.5mm), respectively, corresponding ta nozzle-ta-plate spacings of 6,
8, 10, and 12 at Reynolds numbers ranging from 1,500 ta 12,000. The discontinuity of
the curves is due to the junction between the semi-cylinder and the flat plate.
It can be seen that in general the local heat transfer coefficient displays the same
trend aIong the impingement surface, i.e. decreasing with xIW at all H/W values and
Reynolds numbers. The curves around the stagnation line are aIl very sharp, meaning
that the heat transfer rate at the stagnation line are is much higher than those in the
surrounding area. It can aiso be seen that at any specified condition, the heat transfer rate
at the point on the t1at surface c10sest to the stagnation line is higher than that at sorne of
the points on the lower part of the semi-cylinder surface.
Figures 5.9 - 5. Il show the local heat transfer coefficient distribution for the
narro\ver smooth nozzle (W=5mm), corresponding, respectively, ta nozzle-to-plate
spacings of9. 12, and 18 at Reynolds numbers ranging from 1,500 ta 12.000.
It can be seen that the curves for the narrower nozzle of \V=5mm sho\v similar
trends as those for the wider nozzle of W=7.5mm. For the flat portion, the curves for
W=5mm are generally flatter than those for W=7.5mm. It is noted that there does not
exist a distinct second maximum on the curves.
Ratsurface
-:- Reyr12000-.:.- Rew=6OOO--;'- Re.,.,=3000- :- R&w-=1500
--":""':---.-
--~.--....-----.................. -7, ~'I ---:--~:
,',~-=;~
8 ~ 12 14 16 18 20 22 24
Flatsurface
Smoolh nozzJe: 'W:=5nm
HN/=12
246
Figure 5.10 Local heat transfer coefficient distribution
at various Reynolds numbers (HJW=12)
_600 ----------------~
NE~500 -
- ~- Rew=12000-v- Rew=6OOO-':'- Rew=3000-:- Rew=1500
SfTI)()th nome: W=5mn
K'W=9
o 2 .. 6 8 JN 12 14 16 18 20 22 24
Figure 5.9 Local heat trans1er coefficient distributionat various Reynolds numbers (HI'N=9)
_ 600 ---------------::c:: T
1 1~ 500 _!- :
•i6
Figure 5.11 Local heat transfer coefficient distributionat vanous Reynolds numbers (HJ'N=18)
Rew hloc havg12000 -.- -6000 -.-1500 ~ -<"'"'-
Srnooth nome: 'N=5mm .
HJW=9
600 --------------
o 2 4 6 8 10 JW14 16 18 20 22 24
Figure 5.12 Comparison between local and averageheat transfer coefficients at variousReynolds nurriJers (HI'N=9)
~ lN 5OO-~E .\~ '.
- 1. C ed.. 400 _.~ UrY Aatai .-... surface surface'ü .'-1- -~ 300 ~ ••~-r_o ~I ~.-...ur'" -.-.! ~.~~ --- - .~~ 2IJO ~.;;.~~~ .~
i 100 :-~~~ ~~.~:1::~-.-~-.-;-.-.~o ~~6.:..:..4 _
-:)- Ra,y=12000 ;-&- Ra,y=l6OOO~ Rey,=3000-:- Rey..=1500
5mood1 nozzJe: W=5mn
HIW=18
-ClU;g 400 _ CuM!d. Rat~ surface surfacelUo ~
u ,~ 300 -._li)c('Q
.= 200 ~~ ~,- ~-
; -~~.:~- ~-::-= ......:_=-~... ,_ -L..". --::-:-:-:_:_:
ë; 100 - .-...~~~ ~~,
~ /:~./~.~;~~ 7~~_:_:~~_:~ ......- 6.40----------------
o 2 4 6 8 iJlN 12 14 16 18 20 22 24
_ 600 ---------------~:.::
N
E!SOO -•
•Figure 5. 12 shows both local and average heat transfer coefficients. corresponding
ta the same conditions as those for Figures 5.9. Because there is a unmeasured region of
heat transfer around the junction line and the flow pattern there is considered complex
(unidentified vortex etc.), it is not accurate if we interpolate the values. Sa the average
values are calculated separately for cylindrical surface part and the flat surface part. For
the semi-cylindrical part, an average heat transfer coefficient at a specified x/W
represents the average value of the local heat transfer coefficient from stagnation point ta
the position concemed. For the flat plate part a average value represents that of the local
value tram the first left point on the flat plate to that concemed.
5.-/.2 Conzparison between the Smooth iVo==les
•
Figures 5. 13 shows a comparison between the local heat transfer coefficient
distributions for the two smooth slot nozzles of widths 5mm and 7.Smm. It can be seen
that the heat transfer coefficients for the narrower nozzle (W=Smm) are consistently
greater than those for the wider nozzle (W=7.Srnm) at the same Reynolds number: At
sorne points they are over SO% greater. This great difference is explained in terms of the
77
• fact that the jet from the narrower nozzle has not only a higher velocity at a given Re but
aIso possibly higher turbulence intensity due to the sudden contraction of the air flow at
the nozzle exit, both of which can improve the heat transfer on the impingement surface.
500 --~------------
o :2 4 6 8 10 12 14 16 18 20 22 24xNI
Figure 5.13 Comparison of local heat transfer ratesbetween two smooth nozzles (HI'N=12)
90 ---------------Symmbol Re.. Nozzle width
--- 12000 5.0mm- 12000 7.5mm
___ 6000 5.0mm
- 6000 7.5mm-6-- 3000 S.Omm
-..:r-- 3000 7.5mm
10 -
0---------------a 2 4 6 8 10 12 14 16 18 :0 22 24
xJWFigure 5.14 Comparison of average Nusselt number
between two smooth nozzles
••80 - •
•70 -:. •
~ ,.E 60 -:::; ••
Ë .: •- ..:. ---i 50 -.~;.. "' HJW=12
~4°~R-_ .~
I::~~~::-~~
Re.. NozzJe width12000 S.Omm12000 7.Smm6000 S.Omm6000 7.Smm3000 S.Omm
3000 7.Smm
Symmbol
----HJW=12
•Figure 5.14 shows a comparison between the average Nusselt numbers for the t\vo
smooth nozzles at the same conditions as those in Figure 5.13. It can be found that
although generally the Nusselt number for the narrower nozzle is still greater than that for
the wider nozzle in the high Reynolds number range tested, the difference between them
is reduced significantly. 50 the non-dimensionalization of the heat transfer coefficient in
tenns of the Nusselt number does reduce the difference caused by the nozzle width,
however, there is still an effect due ta the scale of the nozzle
5.5 Heat Transfer for Planar and Non-Planar Surfaces
•Heat transfer for non-planar surface is as expected different from that for a planar
surface in that the flow pattern of the jet on the impingement surface is characterized by
configuration of the surface. It must be noted that HIW used here stands for the distance
between the nozzle and the flat surface for bath planar and non-planar impingement
i8
• surfaces studies. HtI\V denotes the dimensionless distance between the nozzle and the
top of the semi-cylinder.
100 ----------------
Re.v PIa1ar Non-Plalar12000 -... -:-6000 -6- -
1500 -.-
Smooth nozzIe: W:7.5nm
tWJ=10, t\JW=5.7
00...00 - •
••
100 ----------------
o---------------~-o 2 4 6 8 10 12 14 16
Figure 5.16 CorJ1)élrison of local Nu nurriJers bet'Neenplanar and non-planar surfaces
10 12 14 168
Re.., PIanar Non-PIanar12000 -.- -:-6000 ---6--- ---
1500 --.-
5mooth nozzIe: ~7.5nm
HW=6, HtN=1.7
62 4
~~0-------=-------------o
Figt.re 5.15 Co~son d local Nu numbers betv.Eenptanar and no~lanarsurfaces•
o 0 ----------------
o 2 4 6 8 1Q 12 14 16 0 2 4 6 8 10 12 14 16
Figure 5.17 CorT1Jarison of average Nu numbers bet\\een Figure 5.18 Comparison cl average Nu numbers betweenplanar and norrplanar surfaces planar and non-planar surfaces
100 ---------------Re..., Planar NolH'lanar12000 -... -:-6000 ~ ~
1500 -.-
Smooth nol2fe: ~7.Srrm
~O,~5.7
00....80 - •••e.
e•
100 ----------------
LQ)
.c ïO·"E '.~c: 60 _.:
R~ Planar Ncn-Plalar12000 -.- -:-6000 -6- -
1500 -.- ~""""'-
Srmo1h nozzJe: VF7.Snm
tiW=6,~1.7
......00 _1 ••
•••••••
00-
L-a,) -~ 70 - _e='c::
2)~~~1 ••7--
10 - '
•79
•
•
•
Figures 5. 15 and 5.16 show the comparison of local Nusselt number distribution
for planar and non-planar surfaces, for HJW=6 and 8, respectively, while Figures 5.17
and 5.18 show a comparison of corresponding average Nusselt number distributions. The
discontinuity in ail curves is in the consideration of comparing the two distinct regions of
small and large x1W, defined as Region 1 and II, respectively. Region 1 covers the
cylinder surface and its projection on the flat plate. while Region II covers the far region
on the flat plate which extends downstream of the cylindrical surface
It can be seen from Figures 5.15 and 5.16 that for both HJW=6 and 10, the
difference of local Nusselt number distributions between planar surface and non-planar
surface in Region 1 is significantly different from that in Region II. In Region L the local
Nusselt numbers for planar surface are almost ahvays higher than those for the noo
planar surface, except for locations near the stagnation point at the lo\vest Reynolds
number (1500) tested_ The difference is most significantly large at the right border of
Region L up ta nearly 90~/O. The difference at the stagnation point can be small or large
depending on the Reynolds number and H/\V. Generally the difference increases with
xlVI. In contrast with Region 1, the t\vo corresponding curv-es are very dose ta each other
in Region II.
The above phenomena can be attributed to the different flo\\i pattern of the jet on
the impingement surface. In Region I, the jet flow on the semi-cylinder surface is very
complex. It is expected that there exist three-dimensional counterrotating vonices around
the stagnation region. The flow pattern is affected by the jet Reynolds number, location
on the surface and HJW, etc. In Region Il, the fluid flow tends to be unaffected by the
vortices but should be in fact by possible generation of recirculation zones. It is expected
that computational fluid dynamic simulation will show more light on the complex flow
patterns expected.
5.6 Heat Transfer for Rough iVozzles
Three configurations of the rough-fin nozzle. which have the same open area as
that of the smooth nozzle of W=5mm, were tested for a plaoar surface in Chapter 4. The
turbulence intensity of the jet produced by rough nozzles is expected to be higher than
80
• that for a smooth nozzle. Nevertheless not ail the rough nozzles enhance the heat transfer
rates on the impingement surface, as found earlier. The three rough nozzles were tested
for non-planar impingement surface, as well.
5.6.1 Right-Triangular Rough iVo==le (Jaws 2)
HJW=9
o 4 8 12 16 20 24xlW
Figure 5.19 Comparison between smooth nozzle
and Jaws 1 (HfW=9)
~ Re.".=12000. smooth-:- Re...=12000, rough-6- Re...,=6000, smooth-..:r-- Re...=6000, roug h--.- Re.,=1500, smooth
--=- Re.",=1500, rough
1:30 -----------------
o 4 8 12 16 20 24xm
Figure 5.20 Comparison between smooth noozzleand Jaws 1 (HIW=9)
~
120 l H/W=9
1'0 ..
.. 1OO-~QI
Ê 9J -~~
È 80 - -- r-
! : ~ .~~-:::--z " .~
~ 50 -~.~ 4() - ~4
.( 30- -,
•20 -:'_
10--~
0-----------------
700 --------=--~--------.- Re.=12000, smooth-:- Rew=12000, rough~ Rew=6000,s~oth
--~ Re.=6000,rough
~ Rew=1500,s~oth
--:- Rew=1500,rough
~
:i 600 ~E •
lSOO -\1;:
'; 400 -,Cl ~
C.J c:..~ 300 ~..~
.=; \~-\ ea::..~ 200 -~ ~:! ~'." te
~.ii • ~'-
C.J 100 -=1- ~Cl ~ .......
-1 ~_~
--..:;;o---~-------------•
H/W=12
100 -----------------.- Re.=12000, smooth-:- Re.=12000, rough-Â- Re.=6000, smoolh-w-- Re.=6000, rough--.- Re.=1500, smooth
-=-- Re.=1500, rough
4 '2 16 20 24xJW
Figure 5.22 Cornparison between smooth noozzleand Jaws 1 (HJ'N=12)
Q~-~-------------
5CG -----------::---=--------.- Re.=12000.smooth-:- Re.=12000. rough--Â- Re,.=6000, smooth-ù- Re..=6000, rough
- Re.=1500, smooth--:- Re.=1500, rough
a 4 '2'6 20 24x!'N
Figure 5.21 Comparlson between smooth nozzle
and Jaws 1 (H1W=12)
o~---------------
•81
•
Figures 5.19, 5.21 and 5.23 compare the local heat transfer coefficient
distributions between the smooth nozzle and the right-triangular rough nozzle of the same
equivalent width (W=5mrn) at H/W=9, 12 and 18 and at various Reynolds numbers for
the non-planar surface
From the three local heat transfer distribution curves, it can be seen that generally
the difference of the heat transfer rates between the smooth nozzle and laws 1 is
insignificant. The exception i5 that at higher Reynolds numbers (above 6000) the
difference of between the two stagnation point heat transfer rates and locations far from
the stagnation point become somehow evident. For the highest Reynolds number used" i.
e. 12000, the stagnation point heat transfer rate for smooth nozzle is about 10% lower
than that of Jaws 1 at H/\V=9, but about 100/0 higher at HIW=12 and 20% higher at
HJW=18. At locations farthest from the stagnation point (where xJW=22), for Rew
= 12000 the heat transfer rates for smooth nozzle are over 500/0 higher than those for Jaws
1 at HJW=9, 12 and 18.
Figures 5.20, 5.22 and 5.24 compare the corresponding average Nusselt number
distributions, at the same conditions as Figures 5.19, 5.21 and 5.23, respectively. AlI the
HIW=18
~ Re.·12000,Sft'Ooth-:;- Re.=12000, rough-Â- Re.=6000. smooth-ù- Re.=6000, rough
~ Re.=1500.s~h
-=- Rew=1500, rough
:.-~-
~"O.::_:-~~
- --'--..."!-~~.. ~-<:"=-=--=-~
0---------------
70 ---------------
a 4 12'6 20 24xNI
Figure 5.24 Comparison between smooth noolzleand Jaws 1 (HIW=18)
0---------------
400 --------=---=-----~~ Re..,=12000.smooth--:- Rew=12000. rough.-......6- Rew=6000. smooth~ Re..,=6000. rough
---- Rew=1500. smooth-=- Rew=1500. rough
a 4 8 12 l1W 16 20 24
Figure 5.23 Comparison between smooth nozzleand Jaws 1 (HIW=18)
;;CJQ
...l
•
•82
• curves for average values are smoother than those for local values, and show the same
trend as the latter ones, as expected.
The above results for non-planar surface are not consistent with those obtained for
planar surface in previous chapter for Jaws 1, though the turbulence level produced by
Jaws 1 is also expectedly higher than that by the smooth nozzle. The inconsistency can
he attributed to the different flow patterns produced by the clIrvature of the impingement
surfaces.
5.6.2 Equilateral-Triangular Rough No==le (Jaws 2)
•
Figures 5.25, 5.26 and 5.27 compare the average ~usselt number distributions. for
a smooth nozzle (W=5rnm) and Jaws 2 with the same effective width at Hl\V=9, 12 and
18. respectively.
It can he found that nozzle Jaws 2 produced more negative effect, i.e. reduction in
the heat transfer rate than Jaws 1. At higher Reynolds numbers Caver 6000) and H/\V=9,
the heat transfer rates for Jaws 2 in both Region 1 and II are generally lower than those
for the smooth nozzle, and trus is more significant in Region L up to 40~/o. This result,
again, is caunter-intuitive and difficult ta explain in the absence of detailed flow
measurements. The results are reproducible within ± 5~/O.
~ Re =12000. smooth-:- Re:=12000, rough---..- Re =6000, smooth-~- Re·=6000 rough. '- Re.=1500. smooth-=- Re.=1500. rough
HIW=12
90 -.
130 ---------------
o~--------------
a 4 8 12 16 20 24x/W
Figure 5.26 Comparison between smooth noozzleand Jaws 2 (HJW=12)
80 - :••] 70 ~ ...!!! ...
z 60 1 ~~ 50~ ~~ 40 -~ --
~ 30": "- ""'-20 .:-,;.,-~
10 - -----.
--...- Re.=12000, smooth-:- Re.=12000, rough....-......- Re.=6000, smooth~ Re.=6000, rough
----- Re.=1500, smooth-=- Re.=1500. rough90 -.
1::0 ---------------
o 4 12 16 20 24xlW
Figure 5.25 Comparison between smooth noozzleand Jaws 2 (HIW=9)
120 - HIW=9
110 •1
10C -~.0gE:
•83
•80 ~--------------
"..
70 :~ H~18
~ Re.=12000.smooth-:- Re =12000 rough-..- Re:=6000, ~mooth--....- Re.,.,=6000. rough
-- Re.=1500.smooth-=- Re.=1500.rough
.~~~ _.......
-- ---------
•
•
o---~-----------
a 4 8 12 16 20 24xJW
Figure 5.27 Comparison between smooth noozzleand Jaws 2 (HIW=18)
It is shown that for HJW=9 nozzle Jaws 2 produces similar negative effects on
heat transfer rates as Jaws l, and the magnitude is even greater. This is consistent with
previous results obtained for a planar surface, that the heat transfer enhancement effect by
Jaws 2 is less than that by Jaws 1. This fact suggests that turbulence level of the air jet
produced by Jaws 2 may be lower than that of Jaws 1. However, there may be yet other
effects that cannat be identified at this time. For Hl\V= 12 and 18 there is not much
difference between the heat transfer under the smooth nozzle and Jaws 2.
5.6.3 Rectangular Rough Nozzle (Jaws 3)
Figures 5.28, 5.29 and 5.30 show a comparison between the average Nusselt
number distributions under the smooth (W=5mm) and rectangular rough nozzles with the
same equivalent width at HJW=9, 12 and 18, respectively. It is shown that similar to
Jaws 1 and 2, Jaws 3 also generally produced negative etTect on the heat transfer rates
compared with the smooth nozzle, except for a few points. The dirninishing effect is
• most pronounced at H1W=18 and Reynolds numbers of over 6000, and a reduction of up
to 500/0 is noted. Unlike Jaws 1 and Jaws 2, Jaws 3 behaves the same for the planar and
non-planar surfaces.
0------------------a 4 8 12 16 20 24
x/WFigure 5.28 Comparison between smooth noozzle
and Jaws 3 (HJW=9)
-.- Re..,=12000, smooth-:- Re.".=12000, rough....-......- Re..,=6000, smooth~ Re..,=6000,rough
- Re",=1500,smooth-:- Re..,=1500, rough
HIW=12
130 -----------------
120 -
o 4 8 12 16 20 24xJW
Figure 5.29 Comparison between smooth noozzleand Jaws :3 (HIW=12)
0-----------------
110 .~
\... 100-:Q,I~
E:=c
90 -.:.'80 --.- ....-= 70 - =-~ 50 - ~z . ~
~ 50 ~ ._-~
i 40-~,.ct 30 _ ~
•~:~
--- Rew=12000,smoottl-:- Re...=12000, rough---6- Re.=6000, smooth-..:.- Re.=6000, rough
-- Re..,=1500, smooth-=- Re..,=1500, rough
130 \
120 -\ HIW=9i
110 t... 100-7Q,I
Ê 9O-~
ê 80 - -- ..~ 70 ~ r-~ 50~~ "-
& 50 -~ "
j 40-'--30 :. .....-6
20 ::
10--~
•aD -----------------
HIW=8
'r-
a,.
• ••
Re, OIW-+- Present data 12,000 8.5
-:- Gau&Chung 11,000 4570 ~
30-
t":
g 40-....J
80-----------------~ Re..,=12000,smooth-:- Re",,=12000, rough...--..- Re""=6000, smooth-..:.r- Re....=6000, rough
- Re..,=1500,smooth
- Re""=1500, roug h
HJW=1870 ~
-10~~---
...... 60 - ....~ "-~ 50- ,_- -Qi :
~ 40~C:~Z -G.J :~= 30 - --=..ra --:..... -~ ~
ce: 20 - -::.
20 --~--------------
•0-----------------
o 4 8 12 16 20 24xNI
Figure 5.30 Comparison between smooth noozzleand Jaws 3 (HIW=18)
o 2 3xfW4 5
Fig ure 5.31 Comparison with previous work
6
85
•
•
5.7 Data Comparison and Correlation
5. 7.1 (-'omparison wi(h Previolls ~Vork
The publîshed heat transfer data for impingement on a cyiindrical surtàce is very
scarce. Figure 5.3 1 shows a comparison of the present experimental data with those of
Gau and Chung (1991). It can be seen that the present data are generally lower than those
of Gau and Chung. except for locations close ta the stagnation point: the maximum
discrepancy is less than 25~/o. The effect of convex CUf\tature might be one of the reasons
since the t'..vo curvatures (the ratio of the diameter of the impingement cylinder ta that of
the width of the slor nozzle. represented by D/\V) is different by a tàctor of 5 In addition
the present data were obtained \vith a continement surface. \vhile the Gau and Chung data
were not.
- ~ ) C 1J. / . _ orre Q(lOn
~usselt numbers at the stagnation point \vere correlated with the jet Rçvnolds
number and nozzle-to-plate spacing in the most commonly used forms (r: = 0 95)
~li = a . Re b. (MV) .: (5 1)
•
where parameter a. band c are listed in Table 5. L \vhich aiso include results for planar
surface from Table -t.2.
Table 5 1 Correlation parameters of experimental data of stagnation pointNu numbers for the t"'....o smooth nozzles
(Re numbcr range: 1500 - 12.0(0)
Nozzle-widthlsurface-type a b c Experimental conditions
5mm/planar 010 0.71 00050 HIW: 2.2 - 18
5mm/non-planar 0.63 0.71 i -070 MV: 9 - 18
7.5mm/planar 0.13 0.56 1 -0.047 HI\V: 1.6 - 12
7.5 mm/non-p lanar 0.69 0.54 -0.21 HIW: 6-12
From Table 5.1 it can be observed that the exponent on Re number. b. is not
affected by the surface types, but decreases with nozzle widths. suggesting dimensionless
86
•
•
•
parameter (Re number) is better than geometric factor when used in correlation. The
difference results from different ranges of experimental parameters used. The exponent
on HI'N is negative for bath smooth nozzles tested on non-planar surface. implying that
heat transfer rates decreases monotonically with H/W.
5.8 Closure
An extension of experimentai study on heat transfer behavior of contined siot jet
impingement. were conducted on a semi-cylinder attached on top of a flat surface (the
combination is called non-planar surface here). using bath smooth and rough nozzles.
The results sho\v that in comparison with planar surface aIl the three rough
nozzles. i.e. right-triangular (Jaws 1). equilateral (Jaws2) and rectangular Uav.:s 3)
generally produce negative effects on the non-planar impingement heat transtèr.
The comparison of the heat transfer rates bet\veen the two smooth slot nozzle for
non-planar surface shows the similar behavior as that for planar surface. i.e. that the
narrower nozzle produced higher heat transfer rates than the wider one.
Si
• 11CHAPTER 6~
CONCLUSIONS
This study included t\vo parts. (1) comparison of heat transfer rates between t\VO
smooth slot nozzles with different widths (Smm and 7.Smm). (2) the effects of using
rough-edged nozzles of various configurations. i. e. right-triangular. equilateral triangular
and rectangular. on convective heat transtèr rates benveen air jet and planar and non
planar impingement surfaces.
The main conclusions ofthis work are summarized as follo\\is
•
•
,/ The configuration of nozzle-exit-edge can affect heat transfer rate benveen the jet and
the impingement surface (either planar or non-planar) .
,. For planar surface. right-triangular and equilateral-triangualr nozzles can enhance
impingement heat transfer up to 40~'o \vhile the rectangular nozzle actually
decreases the heat transtèr rates;
,. For non-planar surface. right-triangular. equilateral-triangular and rectangular
nozzles generally produce negative effects on the non-planar impingement heat
transfer. up to 50'%.
,/ The enhancement of heat transfer induced by insenion of turbulators depends on the
type and location of the turbulators employed. The insenion of turbulators can be
used to raise local heat transfer rate as required.
,/ The smooth nozzle of 5mm width can produce higher heat transfer rates than that of
7.5mm width at same experimental conditions. This result is probably due to the
higher turbulence level of the jet from the narrower smooth nozzle
88
•
•
•
./ The different effects of three rough nozzles on heat transfer rate for planar and 000
planar surfaces are caused by the difference of gas tlow pattern on the t\\'"o surfaces.
RecommendatioDs
.:. In arder to elucidate the mechanism of heat transtèr behaviors for impinging jet. t10w
structure (velocity distribution and turbulence intensity) of the jet should be measured
\Vith the help of laser Doppler velocimeter (LDV).
•:. Other configurations of rough-edged nozzles and curved targets need to be tested.
•
•
•
REFERENCES
Bhansali A. P. and Black W. Z., "Local, [nstantaneous Heat Transfer Coefficientsfor Jet Impingement on a Phase Change Surface", Transaction of the ASME, Vol. IlS,334-341, May 1996
Brahma R. K., Padhy 1., and Pradhan B., "Experimental studies of heat transfer byslot jet and single/triple row of round jets impinging on semicylindrical concavesurfaces", Heat Transfer Engineering, Vol. 1S, No. 4, 1994
Chung Y. S., Lee D. H., and Lee 1. 5., "Heat transfer characteristics of anaxisymmetric jet impinging on the rib-roughened convex surface", Int. 1. Heat MassTransfer, Vo1.42, pp. 2101-2110,1999
Colucci D. W. and Viskanta R., Effect of nozzle geometry on local convectiveheat transfer to a confined impinging air jet, Experimental Thermal and Fluid Science 13(1996) 71-80
den Ouden C. & Hoogendoom C. J., '"Local convective heat transfer coefficientsfor jets impinging on a plate: Experiments Using a Liquid-crystal Technique", Proc. Sth[nt. Heat Transfer Conf., AIChE, New York, Vol. 5, pp.293-297, 1974
Gabour L. A. and Lienhard 1. H. V., "Wall Roughness Etfects on StagnationPoint Heat Transfer Beneath an [mpinging Liquid Jet", 1. of Heat Transfer, Transactionsof the ASME, Vol. 116, pp. 81-87, 1994
Gardon Robert and Cahit Ak.firat 1., '"The raie of turbulence in determining theheat-transfer characteristics of impinging jets", Int. J. Heat Mass Transfer, Vol. 8, pp.1261-1272, 1965
Gardon Robert and 1. C. Akfirat, ""Heat transfer characteristics of impinging two-dimensional air jets", Transactions of the ASMEJ. of Heat Transfer, ppl01-1OS, February( 1966)
Garimella, S. V., and R. A. Rice, "Conftned and submerged liquid jetimpingement heat transfer," ASME 1. Heat Transfer, 117, 871 (1995)
Gau C. and Chung C. M., ""Surface curvature effect of slot-air-jet impingementcooling flow and heat transfer process", 1. of Heat Transfer, Transactions of the ASME,Vol. 113, 858-864, November 1991
Geunyoung Yang, Mansoo Choi and Joan Sik Lee, '"An experimental study of slotjet impingement cooling on concave surface: etfects of nozzle configuration andcurvature", Int. 1. Heat Mass Transfer, Vol. 42, pp2199-2209, 1999
Gm. S. Azad, Yizhe Huang and Je-Chin Han, '"[mpingement heat transfer ondimpled surfaces using a transient liquid crystal technique", Journal of Thermophysicsand Heat Transfer, Vol. 14, No. 2, April-June 2000
Gutmark E., Wolfshtein M., and Wygnanski 1., ""The plane turbulent irnpingingjet", 1. Fluid Mechanics. Vo1.88, No. 4, pp737-756, 1978
•
•
•
Haneda Y., Tsuchiya Y., Nakabe K. and Suzuki K., "Enhancement of impingingjet heat transfer by making use of mechano-fluid interactive flow oscillation", [nt. 1. Int.1. Heat Mass Transfer, Vol. 19, pp. 115-124, 1998
Huang Bing, "'Heat transfer under an inclined slot jet impinging on a movingsurface", Ph. 0 thesis, McGili University, 1988
Janice A. Fitzgerald and Suresh V. Garimella, HA study of the flow field of aconfined and submerged impinging jet", Int. 1. Heat Mass Transfer, VolAI, Nos 5-9,pp 1025-1034 (1998)
Kline, S. 1., and F. A. McClintock, "Describing uncertainty ln single-sampleexperiments", Mechanical Engineering, Vol. 75, l, pp. 3-9, January 1953
Lee D. H., Chung Y. S., and Kim D. S., "'Turbulent flow and heat transfermeasurements on a curved surface with a fully developed round impinging jet", Int. 1.Heat and Fluid Flow, Vol. 18, pp160-169, 1997
Lin Z. H., YJ. Chou and Y.H. Hung, "Heat transfer behaviors of a confined slotjet impingement", Int. 1. Heat Mass Transfer, VoIAO, No.5, ppI095-II07(l997)
Luis A. Brignoni, Suresh V. Garimella, Effects of nozzle-inlet chamfering onpressure dro and heat transfer in confined air jet impingement, Int. 1. Heat and MassTransfer, 43 (2000) 1133-1139
Manseo Choi. Han Seoung Yoo, Geunyoung Yang, 100n Sik Lee and Dong KeeSohn, Measurements of impinging jet flow and heat transfer on a semi-circular concavesurface, Int. J. Heat and Mass Transfer, 43 (2000) 1811-1822
Martin H., "Heat and mass transfer between impinging gas jets and solid surface",Advances in Heat Transfer, 13, 1-60, 1977
Masood Mesbah, James W. Baughn and Christopher R. Yap, '·The effect ofcurvature on the local heat transfer to an impinging jet on a hemispherically concavesurface", The Ninth International Symposium on transport phenomena in Thermal-FluidsEngineering, 795-800, Singapore, June 1996
McCleave R. W. Hlmpinging jet heat transfer with turbulence enhancement at thenozzle", M. Eng. Thesis, McGiIl University, 1993
McDaniel C. S. and Webb B. W., Siot jet impingement heat transfer from circularcylinders, Int. 1. Heat and Mass Transfer, 43 (2000) 1975-1985
Metzger D. E., Yamashita T. and Jenkins C. W., Hlmpingement cooling ofconcave surfaces with lines of circular air jets", J. of Engineering for Power, 149-158,Transactions of the ASME, July 1969
Miyake G., Hirata M. and Kasagi N., '"Two-dimensional jet impinging on a wallwith roughness elements", Experimental Heat Transfer, 7: 1-17, 1994
Mohanty A. K. and Tawfek A. A., "'Heat transfer due to a round jet impingingnormal to a flat surface", lnt. 1. Heat Mass Transfer, Vol. 36, NO.6 pp. 1639-1647, 1993
Mujumdar A. S. and Huang B., "Impingement Drying", Chapter 14 in Handbookof Industrial Drying edited by Mujumdar A. S., 2nd Edition, Marcel Dekker, Inc., pp489 501, 1995
Il
•
•
•
Mujumdar,A. S. "Innovation in Drying", in Trends in Heat and Mass Transfer,Vol. 4, pp. 115-126,1998 Trivandrum, [ndia (ISBN 81-258-0047-6)
Pekdemir T. and Davies T. W., "A study of convective heat and mass transferfrom circular cylinders exposed to a jet flow", Proceedings of the Tenth InternationalHeat Transfer Conference, Vol.3, 4-EC-17, 83-88, 1994
Polat S., Van Heiningen A. R. P. and Douglas W. 1. M., "Sensor for transient heatflux at a surface with throughtlow", lnt. 1. Heat Mass Transfer, Vol. 34, No. 6, pp. 15151523, 1979
Roy A. Rice and Suresh V. Garimella, "Heat transfer from discrete heat sourcesusing an axisymmetric. submerged and confined liquid jet", Proceedings of the TenthInternational Heat Transfer Conference, Vol. 3, 4-EC-18, 89-94, Brighton, UK, 1994
Saad N. R., A. S. Mujurndar, W. Abdel Messeh and W. 1. M. Douglas, "Localheat transfer characteristics for staggered arrays of circular impinging jets with crossflow of spent air, ASMP papers, pp. 1-8, 1980
Sailor 1. David, Daniel 1. Rohli and Qianli Fu. "Effect of variable duty cycle flowpulsations on heat transfer enhancement for an impinging air jet", Int. 1. of Heat andFluid Flow, Vol. 20, pp. 574-580, 1999
Schuh H. and Persson B., "Heat transfer on circular cylinders exposed ta free-jetflow", Int. J. Heat Mass Transfer, Vol. 7, pp. 1257-1271, 1964
Seyed-Yagoobi 1., Narayanan V. and Page R. H., "Comparison of heat transfercharacteristics of radial jet reattachment nozzle ta in-line impinging jet nozzle", J. ofHeat Transfer, Vol. 120, pp335-341, May 1998
Sheriff H. S. and Zumbrunnen D. A.. "Local and instantaneous heat transfercharacteristics of arrays of pulsating jets, Trans. Of the ASME, 1. of Heat Transfer,Vol. 121, 341-348, May 1999
Shi Yuling, M. Eng. thesis, Dept. of Chem. Eng., National University ofSingapore, 2001
Sing-Min Kum, Yasuo Kawaguchi and Jeong-Yun Seo, "A study on heat-transferenhancement by a square-rod array in an impinging jet system", Heat Transfer Japanese Research, V.24, No.2, P 193-207, 1995
Siripon Anantawaraskul, "Heat transfer enhancement under a turbulent impingingslot jet", M. Eng thesis, McGill University, 2000
Siripon Anantawaraskul and Arun S. Mujurndar, "Techniques for enhancement ofimpingement heat and mass transfer rates", Proceedings of the First Asian-AustralianDrying Conference (ADC'99), 117-135, Bali, Indonesia, 1999
Sparrow E. M. and Wong T. C. , "Impingement transfer coefficients due toinitiaUy laminar slot jets", Int. 1. Heat Mass Transfer, Vol. 18, pp. 597-605. 1975
Suna Polat, "Transport phenomena under jets impinging on a moving surface withthroughflow", Ph. D thesis, McGiIl University, 1988
Van Heiningen, "'Heat transfer under an impinging slot jet", Ph. D thesis. McGillUniversity, 1982
III
•
•
•
White Frank M., "FLUID rvŒCHNAMICS", 2nd edition, McGraw-Hill BookCompany, 1986
Williams F. and Watts J., "The development of rough surfaces with improvedheat transfer performance and a study of the mechanisms involved, Proc. 4th lnt. HeatTransfer Conf, Versai les, 1974
Xiaojun Yan and Nader Saniei, "Heat transfer from an obliquely impingingcircular air jet to a flat plate", lnt. 1. Heat Mass Transfer Vol.lS, pp. 591-599, 1997
Yoshida H., Suenaga K. and Echigo R., "Turbulence structure and heat transfer ofa two-dimensional impinging jet with gas-solid suspensions", lnt. J. Heat Mass Transfer,Vol. 33, No. 5. pp. 859-867, 1990
IV
•
•
Appendix 1 Uncertainty Analysis
(A). Standard A,;fethod ofSimple Sample Anaf:ysis
The objective of uncertainty analysis is to estimate the probable random error in
experimemaI results. Analysis of uncertainty is very important since experimental data
are often used to supplement engineering analysis as a basis for design and errors always
are present when experimental measurements are made. It is aIso useful to analyze the
uncertainty of heat transfer coefficients, According te the standard method suggested by
Kline and ~lcKIintock (1953). a typical and most acceptable solution is suitable for a
"single-sample" experimem in engineering \vork, \vhere only one measurement is made
for each point. A reasonabIe estimate of the measurement uncertainty due ta random
error in a single-sample experimem usually is plus or minus half the smallest scale
division (the least count) of the instrument.
Suppose that measurements of independent variables. XI. X:... , '. xn, are made in a
experiment. The relative uncertainty of each independently measured quantity is
estimated as Uj. The measurements are used ta calculate a quantity. Z. for the experiment.
It is desired to analyze how errors in the ~ x, ',. propagate into the calculation of Z from
measured values. ln generaL Z may be expressed mathematically as Z = 2(:'<1. x:.... :'<n).
The effect on Z of an error in measuring an individual x. may be estimated by analogy to
the derivative of a function. :\ variation. 8xj. in x. \vould cause variation 8Z. in Z.
(:\-1 )
The relative variation in R is
(:\-2)
•
Eqn. (A-2) may be used to estimate the relative uncertainty In the result due to
uncertainty in XI' Introducing the notation for relative uncertainty. we can obtain
Xl CL . \ "')UZi = Z~UXI (.""\.-J
CXI
Then the best representation for the relative uncertainty of the result Z is
v
• 'Ir(Xl êl ): (x: èl ): .(Xn Cl ): ~!:Uz=~ Z êx: U i -+- Z êx: U : . -+- ...... Z iXn Un (:-\-4)
If the uncertainties in the independent variables are ail glven in the same odds. the
relative uncertainty in Z having these odds is simplified as
Uz = (( u ;): ~ ( u ::) : ~. o' ~ ( U 3): ] 1 :
Eqn. (:\-5) is especially useful when the differentiatian is complex.
(B). L."ncertainty Ana(vsis ofHeat Transier Coefficient
(:-\-5)
The ability ta evaluate the uncertainty in the heat transfer coefficient. h. lies in the
ability ta determine the uncertainry in each af the independent variables related ta h.
Applying aforementianed method ta the analysis of hx in foHov/ing equatian. in \vhich h
can be expressed mathematically as a function of q.:, Iwo:>: and TJ.
h = qc'< T....x-T (A-6)
• Here \ve treat (Two =-: - TJ ) as ane variable and use Eqn. (:-\-3) ta estimate the relative
uncenainty af h.
By differentiating. \ve can abtain
and
q= dt...hx êq= u. = U:
T....... - T! éh....----'---- U: = - U ::
h.... ê(T.v..~-T;)
(:-\-ï)
(:-\-8)
\vhere Ul and u:: are the relative uncertainty af q..: and (T\\.:..-TJ) respectively. Fram Eqn.
(:-\-5) we can get the relative uncertainty of hx
(:-\-9)
•VI
(.-\-10)
•
•
•
(C). Sample Calculation
In generaL we can use Eqn. (:\-4) to calculate the relative uncertainty of
experimental results. If the differentiation is complex we can simplify the calculation
procedures by assuming that the uncertainties in aIl the independent variables are ail
given with the same odd5 and then use Eqn. (:\-5). In present srudy the heat transtèr
coefficient i5 calculated from following equation (see Chapter 3 for detail):
1 1 ktAr ( ...,hi = f' T ) [1 R - -[ [: - : - [; - : - .:. [/ ) ]
r1s ([/- ..
It is worth mentioning that .-\.; and A" are deflned only for study and thus do not
contribute to the uncenainty in h. Table :\-1 lists an example of estimated values of
relative uncertainty of aIl variables of a real data point from an experiment.
Table .01..-1. Cncertainty analysis of one point of real data
direct variable estimated measured estimated relative times used for
used x. uncertaint or known uncertainty calculation
1
y (~XI) value x. (Ui=~Xl/XI)
tl_[ (JC) 0.2 55.92 1000356 1
1 1
ti (uC)1 0.2
1
55.50 1 0.00360...,-
!JC 0.21
56.26 1 0.00356 1t.-I ( )
TJ
(vC) 0.21
27 .20 0.00735 11
l (amp) 0.51
20 1 0.025 .f
k((W/mK) - 19.5 0.02 l
Tl (nm) - 5.20x 10-9 0.02 l
1. .. - ., indicates t/lat tlu eslimated l'alue is no! available bul relillil'f! uncertaÎlrly etUI be estÎlnaled2. wlren swnming tire square ofa single lYuiable. it is counted the tÏ1ne il appears in the equatiOlr
Using Eqn. (:\-5). the estimated relative uncertainty of h is calculated as 5.83%.
It can be found from the above table that the value of electric current contributes most to
the result because it i5 squared and its uncertainty is 4-8 times more than that of
VII
• temperatures measured by thermocouples. So power supply with accurate, high and
stable current is recommended for less uncertainty (higher accuracy) of experimental
results.
Appendix 2 Estimation of Heat Losses
The net convective heat, Qc, dissipated from the target surface to the air jet. can then be
calculated from Eqn. (A-12). The total power input is Qt which equals (=R.. where l is the
electric current of OC power supply and R is the electric resistance of the steel foil. Qi is
the intemal-energy change of the stainless steel foil. Qk is the conductive heat loss ta the
insulation layer. Qr is the radiative heat loss from the stainless steel sheet surface through
insulation layers. Figure A-l is a schematic of a part of a studied section of impingement
surface which includes three overlapped layers, i.e. stainless foil, fiber glass and other
insulation.
(A-11)
(A-12)1.e.
The objective of estimating heat lasses is to obtain accurate experimental results,
1. e. heat transfer coefficients in this study. As described in Chapter 3, the total heat
generated from the metal foil heater, Qt, is converted into the following four heat transtèr
modes: (1) convective heat dissipated from the heated surface. Qc, i.e. the main part~ (2)
intemal-energy change of the stainless steel sheet, Qi. (3) conductive heat loss. Qk~ and
(4) radiative heat loss, Qr: The heat balance is thus as
Qt = Qc + Qi +- Qk - Qr
Qc = Qt - Qi - Qk - Qr
•
Figure A-l Schematic of a part of studied section of impingement surface•
Ts
TRadiatil'c heat, Qr
L
Top la~'er:
Stainlcss steel roil
l\'liddlc layer:Fiber glass
Bottom layer:Other insulation
VIII
• In Eqn. (A-12), the radiative heat Qr and conductive heat From the heating steel
foil to its underside insulation Qk need to be estimated. Radiative heat can be calculated
using the following equation:
(A-l3)
\vhere E = 0.1 is the emissivity of the smooth stainless steel foiL cr = 5.67 ( 10-8
W/(m 2·K4) is the Stefan-Boltzmann constant. :\ = B< L is the surface area of the studied
section. Ts is the temperature of the heat steel foil and TJ is the jet temperature.
Conductive heat through the fiber glass can be calculated using the following equation.
Q - k ,ts-li.: (.' -1 1)k - F'G·""\. ~l ""\. -t
\vhere krG = 0.038 (W/mK) is the thermal conductivity of fiber glass. \[ = ~mm is the
thickness of the fiber glass layer. ts is the surtàce temperature of the heating foil and tk is
the temperature of the bottom of the tiber glass layer. The samp le calculation is sho\vn as
follows.
Sampie calculation of the radiative heat Or
• ln Eqn. (:\.-13). we have
E = 0.1
G = 5 67 .< 10-~ W!(m2·K~)
A = B ." L = 57 mm ( 375 mm = 2.14 ( 10~ m:
•
TJ :::: (273~27)K =300 K
T s :::: (273~85)K = 358 K
thus \ve can get Qk = 0.01 009\V, dividing it by Qt = 0 7460 Vv', the total po\ver input to
this section at the applied electric current value of 20 amp. i.e.
Qk/Qt:::: 01009/0.7460 :::: 1.3 5~/O
Here Ts = 85°C = 358K is the highest temperature measured on the surtàce. \vhich is the
farthest From the stagnation point. At the stagnation point, where the temperature is the
lowest on the impingement surface, let Ts = 451)C, then
Qk/Qt = 0.002577/0.7460 :::: OJ5~/o
which is just a quarter of that of the highest temperature point.
Though the radiative heat across the impingement surtàce is not even. the
maximum radiative heat loss is generally not great than 1.5°fcJ of the total power input.
LX
• Sampie calculation of the conducti~·e heat Ok
ln Eqn. (A-l4), we have
kFG = 0.038 (W/rnK)
A = 2, l4.< lO-" m:!
:\1 =4 mm = 0.004 m
ts - tl\. = 12°C (the maximum temperature ditference measured)
where to; - tk is the temperature drop across the fiber glass layer. which was obtained by
inserting a few thermocouples under the layer. Thus we can get Qk = 002437 \V.
dividing it by Qt = 0.7460 W. i.e.
Qk/Qt = 0.02437/0.7460 = 3.27~/o
As the same as the radiative heat loss. the conductive heat across the impingement 1
1 surtàce is not even neither. but the maximum conductive heat 105S is generallv less than
14% of the total power input. - .
Though the addition of radiative heat 1055 and conductive heat loss is still \'o'ithin
• the uncertainty limit (below 6~/à) of heat transfer coefficient. they were counted in the
calculation of heat transfer coefficient in arder to get accurate results.
Because the other insulation layers under the fiber glass layer is very thick (more
than SOmm). the conductive heat through them is negligible compared \vith that through
the tiber glass and thus was not counted,
•x
• •Appendix 3 Experimental Data on Local Nusselt Number
•(A) S11100th Nozzle 0.( J'V=5IJUll.(or Planar L)IIJ:!ace
HIW=2.2 HIW=6x!W Re=12000 Re=9000 Re=6000 Re=3000 Re=1500 Re=12000 Re=9000 Re=6000 Re=3000 Re=1500
0 54 50 41 30 18.1 86 74 55 35 19.90.8 54 49 40 30 17.7 85 73 54 34 19.61.5 53 48 39 29 17 64 71 53 33 18.62.3 51 47 38 27 16.1 83 69 51 31 17.43 50 45 36 26 15.3 81 67 49 29 16.3
3.8 49 44 35 25 14.6 79 65 47 28 15.54.5 47 43 34 24 14 76 62 45 27 14.85.3 46 42 33 23 13.5 73 60 43 25 14.26 46 41 32 22 13.1 71 59 42 24 13.7
6.8 46 41 32 22 12.8 69 57 41 24 13.37.5 45 40 31 21 12.4 68 56 40 23 12.88.3 45 40 31 21 12.1 67 55 39 22 12.59 45 40 30 20 11.8 66 55 38 22 12.1
9.8 44 39 30 19.7 11.5 65 54 38 21 11.810.5 44 39 29 19.2 11.2 64 53 37 21 11.511.3 43 38 29 18.8 10.9 63 52 36 20 11.212 43 38 28 18.3 10.7 62 51 36 19.5 11
12.8 42 37 28 17.9 10.4 61 50 35 19.1 10.713.5 42 36 27 17.5 10.2 60 49 34 18.6 10.515 40 35 26 16.6 9.8 58 48 33 17.8 10.118 38 33 24 15.5 9 54 45 30 16.5 9.421 35 31 23 14.4 8.4 50 42 28 15.4 8.824 33 29 21 13.5 7.9 47 39 27 14.4 8.3
XI
• • •HIW=9 HIW=18
xIW Re=12000 Re=9000 Re=6000 Re=3000 Re=1500 Re=12000 Re=9000 Re=6000 Re=3000 Re=15000 91 74 55 34 19.5 56 46 34 19.9 12.1
0.8 91 73 33 33 19.3 55 46 34 19.5 11.81.5 88 71 22 32 18.5 55 46 33 19.2 11.52.3 86 69 17.7 31 17.5 54 45 33 18.9 11.33 83 66 15.8 29 16.6 53 44 32 18.4 11
3.8 80 63 14.5 28 15.8 52 43 31 17.9 10.74.5 77 59 13.6 26 15 50 42 30 17.5 10.55.3 74 57 12.9 25 14.4 49 40 29 17 10.26 71 55 12.3 24 13.8 48 39 29 16.6 10
6.8 69 53 11.8 23 13.3 46 39 28 16.2 9.87.5 67 52 11.4 23 12.9 45 38 27 15.8 9.56.3 66 51 11.1 22 12.5 44 37 27 15.4 9.39 65 50 10.7 21.5 12.1 44 36 26 15.1 9.1
9.8 64 49 10.5 20.9 11.8 43 36 26 14.8 910.5 62 48 10.2 20.3 11.5 42 35 25 14.5 8.811.3 61 47 10 19.9 11.2 41 34 25 14.2 8.612 60 46 9.8 19.4 11 41 34 24 13.9 8.5
12.8 59 45 9.6 19 10.7 40 33 24 13.7 8.313.5 58 45 9.4 18.6 10.5 39 33 23 13.5 8.215 56 43 9.1 17.8 10.1 38 32 23 13.1 7.9
16.5 54 41 8.8 17.1 9.7 37 31 22 12.7 7.718 52 40 8.6 16.5 9.4 36 30 21.3 12.3 7.5
19.5 50 38 8.3 16 9.1 35 29 20.7 12 7.321 48 37 8.1 15.5 8.8 34 28 20.1 11.7 7.1
22.5 47 36 7.9 15 8.6 33 27 19.6 11.4 6.924 45 35 7.8 14.5 8.3 32 27 19.1 11.1 6.8
XII
• • •(8) Snlooth Nozz/e of ~V=7.51111Il.for Il/anar Slll:face
HIW=1.6 HIW= 4xIW Re=12000 Re=9000 Re=6000 Re=3000 Re=1500 Re=12000 Re=9000 Re=6000 Re=3000 Re=1500a 83 71 55 36 24 80 74 56 35 22.5
0.5 81 71 54 36 25 80 73 56 35 22.21 80 70 53 36 24 79 71 55 34 21.7
1.5 79 69 52 35 23 77 69 53 33 20.62 77 67 50 33 22 75 67 51 32 19.5
2.5 74 65 48 32 20.9 73 65 49 31 18.83 72 62 46 30 19.9 71 63 48 30 18.2
3.5 69 60 45 29 19.1 69 61 46 29 17.74 67 58 43 28 18.4 67 59 45 28 17.1
4.5 65 57 42 28 17.9 65 58 43 27 16.65 64 56 41 27 17.4 64 57 42 26 16.2
5.5 63 55 41 26 17 63 56 41 26 15.86 62 54 40 26 16.6 62 55 41 25 15.4
6.5 62 54 39 25 16.2 61 54 40 25 15.17 61 53 39 25 15.9 60 53 39 24 14.8
7.5 60 52 38 24 15.5 59 52 36 24 14.58 59 52 37 24 15.2 58 52 38 23 14.2
8.5 58 51 37 23 14.8 57 51 37 23 13.99 57 50 36 23 14.5 57 50 37 23 13.710 56 49 35 22 14 55 49 36 22 13.211 54 47 33 21 13.4 54 47 35 21 12.712 52 45 32 20.3 12.9 53 46 34 20.7 12.313 50 44 31 19.6 12.5 51 45 33 20.2 1214 49 42 30 19 12.1 49.7 44 32 19.7 11.615 47 41 29 18.4 11.7 48.4 42 31 19.3 11.316 46 40 28 17.9 11.4 47.2 41 31 18.8 11
XIII
• • •H/W=8 HIW=12
xIW Re=12000 Re=9000 Re=6000 Re=3000 Re=1500 Re=12000 Re=9000 Re=6000 Re=3000 Re=15000 81 66 53 34 22 73 59 43 27 18.1
0.5 79 66 53 34 22 73 59 43 26 17.81 77 64 51 33 22 71 58 42 26 17.5
1.5 74 62 49 32 21 69 56 41 25 17.12 71 59 47 31 20 66 54 40 25 16.6
2.5 68 57 45 30 19.6 64 52 39 24 16.33 66 55 44 29 19.1 61 51 38 24 15.9
3.5 64 53 42 28 18.5 59 49 37 23 15.54 62 51 41 27 17.9 57 47 35 22 15.1
4.5 60 49 39 26 17.3 55 46 34 22 14.65 59 48 38 25 16.8 54 45 34 21 14.3
5.5 57 47 37 24 16.3 52 44 33 21 13.96 56 46 37 24 15.9 51 43 32 20 13.6
6.5 55 45 36 23 15.6 50 42 31 19.7 13.37 54 44 35 23 15.2 49 41 31 19.2 13.0
7.5 53 43 35 22 14.9 48 40 30 18,9 12.78 53 43 34 22 14.6 47 39 29 18.5 12.5
8.5 52 42 34 21 14.3 46 39 29 18.2 12.39 51 42 33 21 14.1 45 38 28 17.9 12.010 50 41 32 20 13.6 44 37 28 17.3 11.711 49 40 31 19.5 13,2 42 36 27 16.8 11.312 48 39 31 18.9 12.8 41 35 26 16.4 11.013 46 38 30 18.4 12.5 40 34 25 16.0 10.714 45 37 29 17.9 12.1 39 33 25 15.6 10.515 44 36 29 17.4 11.8 38 32 24 15.2 10.216 43 35 28 17.0 11.5 37 31 24 14.8 10.0
XIV
